ScienceDirect
www.elsevier.com/locate/procedia Procedia CIRP 87 (2020) 491–496
2212-8271 © 2020 The Authors. Published by Elsevier B.V.
This is an open access article under the CC BY-NC-ND license (http://creativecommons.org/licenses/by-nc-nd/4.0/) Peer-review under responsibility of the scientific committee of the 5th CIRP CSI 2020
10.1016/j.procir.2020.02.083
© 2020 The Authors. Published by Elsevier B.V.
This is an open access article under the CC BY-NC-ND license (http://creativecommons.org/licenses/by-nc-nd/4.0/) Peer-review under responsibility of the scientific committee of the 5th CIRP CSI 2020
Keywords: roller burnishing; in-situ imaging; surface integrity
1. Introduction
Roller burnishing is commonly used to improve the surface integrity in the automobile and aero industry [1, 2] for its ability to generate deep compressive residual stresses and work- hardened layers while retaining a relatively smooth surface finish than the shot peening process.
It is widely agreed that the normal burnishing force and feedrate play an important role in the generation of residual stress [3]. Besides, the surface integrity could be enhanced by multi-pass rolling, but caution should be noted with the generation of delamination of the surface layer due to the excessive work hardening [4, 5]. Grain refinement could be also achieved with multi-pass burnishing [6, 7]. The burnishing was also used on the shot-peened work to further improve the surface roughness, microhardness, and fatigue life. Fatigue behavior at elevated temperature (about 450°C) of Ti6Al4V treated by deep rolling and laser shot peening was compared. It was found work-hardened nanoscale grains generated by deep rolling play a critical role in the enhancement of fatigue life
given the fact that the almost complete relaxation of the near- surface residual stress at elevated temperature. However, it was found that excessive compressive residual stress and grain refinement do not lead to the best anti-corrosion performance [8].
To study the relationship between the burnishing parameters and the residual stress, empirical, analytical and finite element models (FEM) have been proposed. Empirical models involving pressure, speed, and feedrate were developed for the prediction of surface roughness and residual stress based on experimental data [9]. An empirical model relating the microhardness and the rolling speed, depth and feedrate was proposed based on the response surface method [4]. However, it should be noted that the applicable ranges of the empirical models are limited to the experimental tests. An analytical residual stress predictive model taking into account the initial stress was proposed based on the Hertz contact theory and elastoplastic theory for the roller burnishing process [10]. The shape and magnitude of the residual stress field were in good accordance with experimental measurements. 2D FEM was
5th CIRP CSI 2020
Experimental and numerical study of the subsurface deformation and residual stress during the roller burnishing process
Dong Zhang
a, Xiao-Ming Zhang
a*, Han Ding
aa School of Mechanical Science and Engineering, Huazhong University of Science and Technology, Wuhan 430074, China
* Corresponding author. Tel.: +86-27-87559842; fax: +86-27-87559842. E-mail address: [email protected], [email protected]
Abstract
Roller burnishing process is being extensively used to enhance the fatigue life of the aircraft engine components by introducing compressive residual stress and work hardening. The temperature rise during the burnishing process could be neglected due to the low friction rolling contact between the roller and workpiece, thus simplifying it an elastoplastic deformation process. In this paper, the digital image correlation technique was adopted to obtain the subsurface deformation field during the roller burnishing process. A 3D finite element model was built up to simulate the burnishing process. The predicted surface integrity parameters including subsurface deformation and residual stress are compared with the experimental measurements.
also used to predict the residual stress [11]. Compared with the experimental measurements, it was found the predicted results at the surface deviate from the measurement while at a deeper depth, matches well. 3D FEM was also used to simulate the multi-pass burnishing process for its reliable predictions near the surface and more realistic tool/workpiece contact [12, 13].
Prediction results by the FEM appeared to agree reasonably well with measurements.
However, it is found that little attention has been paid to the subsurface deformation during the burnishing process, which is another key surface integrity parameter. Experimental measurements would be helpful to validate the numerical model and to understand the formation mechanisms of the surface integrity. Recently, the image correlation techniques, such as digital image correlation (DIC) [14, 15] and particle image velocimetry [16] have been proved to be useful to measure the deformation field during the cutting process.
In this paper, the DIC was used to measure the subsurface deformation during the roller burnishing process. The 3D FEM simulation was conducted to predict the subsurface deformation and residual stress. Comparisons with the experimental measurements are conducted. The remainder of the paper offers the following.
Firstly, the experimental setup including the burnishing system and camera-lighting system will be introduced in detail.
Besides, the parameters of the burnishing process, the data acquisition, the image correlation, and the X-ray diffractometer (XRD) will be described.
Secondly, the FEM simulation of the burnishing process including the material model, friction and the derivation of the residual stress will be given.
Lastly, the measured subsurface deformation fields and residual stress of the workpiece during the burnishing process will be given and compared with the FEM simulations.
2. Experimental set-up
The roller burnishing tests were conducted on a computer numerically controlled lathe. The experimental setup consisted of a burnishing system and a camera-lighting system as shown in Fig. 1.
Contact switch
DAQ systemPulse generator DriverLED
Dynamometer
Roller Workpiece
Double shutter camera ω
Fixture
LED
Work LED
Lens
Roller
ω b
a
Fig. 1. Structure of the experimental setup and its working principle
2.1. Roller burnishing conditions
The workpiece used in this paper was made of Al 7075-T351.
The 40 x 18 x 4 mm workpiece had been grinded, polished and then heat-treated at 450 °C for 3 hours. After that, the side
surface was polished using an 800 grit sandpaper. The workpiece hardness was measured as HRA 18.8 ± 3.1. The burnishing tool adopted was a roller bearing (SKF 6200-2Z) with a cylindrical surface and a diameter of 30 mm. The hardness of the cylindrical surface was measured as HRC 63.9
± 2.4.
To reduce the thermal softening and strain rate hardening of the workpiece material, low burnishing velocity is preferred.
Due to the minimum speed limit of the spindle, single-pass burnishing tests were conducted manually without any coolant or lubricate. The burnishing velocity was estimated to be less than 1 m/min which was measured by a high-speed counter parallelly connected to the spindle’s encoder. Two burnishing tests were conducted for each level of set burnishing depth and the measured specific normal burnishing forces (normal force divided by the width of the workpiece) are given in Table 1. It should be noted that the actual burnishing depth could be different from the set value due to the flexibility of the burnishing system and the large normal force.
Table 1. Burnishing depth and the measured specific normal force Level Set burnishing depth d (μm) Specific normal force F (N/mm)
1 50 64 ± 3.5
2 100 134.4 ± 3.1
3 150 188.8 ± 1.3
2.2. Acquisition devices
A double shutter camera was used to measure the subsurface deformation with a sensor size of 1392 x 1040 pixels and the spatial resolution was calibrated as 5.73 µm per pixel. A high power white LED was used to illuminate the scene. Three images were recorded for each burnishing test, corresponding to the workpiece before, during and after the burnishing process as shown in Fig. 2.
Before
burnishing During
burnishing After
burnishing
Residual deformation Burnishing
deformation
Fig. 2. Illustration of the DIC process for the burnishing process
An open-source DIC software [12] was utilized to conduct the image correlation to derive the subsurface deformation during and after the burnishing process. The subset used for correlation was a circle with a radius of 20 pixels. The
reliability of the DIC results is characterized by the zero-mean normalized cross-correlation (ZNCC) which always falls into a range of [-1, 1]. The higher the ZNCC gets, the more are those two images correlated [17]. For the images in this paper, it was found that the ZNCC was commonly around 0.95, indicating an accurate result.
2.3. Residual stress measurement
After the burnishing process, the depth-distributed residual stress in the hoop direction σxx was measured using a Proto iXRD X-ray diffractometer based on the sin2ψ technique and the successive layer removal was realized by electro-polishing.
The XRD parameters are given in Table 2.
Table 2. XRD parameters for the residual stress measurements
Radiation source Cr-Kα
lattice plane (hkl) 311
Bragg angle (°) 139
Incident angle ψ (°) ±20.5, ±11.52, ±1.52, 0
3. FEM modeling of the roller burnishing process
In this paper, the commercially available finite element simulation code (Simulia Abaqus 6.14-explicit) was adopted to simulate the burnishing process as demonstrated in Fig. 3. Due to the free traction boundary of the side surface in the z- direction, a 3D FEM model was set up with a workpiece size of L25 x H8 x B4 mm3.
F V
Work Roller
L H B
C R
x
y y
z
Fig. 3. Demonstration of the 3D FEM simulation of the burnishing process
The workpiece material’s plastic behavior is described using the following equation
σ= +A Bεn (1)
where A is the initial yield stress, B is the hardening modulus, and n is the hardening exponent. The elastic properties of the workpiece are as follows: the Young’s modulus E = 71.7 GPa and the Poisson’s ratio ν = 0.33. The plastic parameters were identified as A = 112.06 MPa, B = 685.96 MPa and n = 0.504 using two tensile tests with a maximum strain around 0.1. Rigid body constraint was assumed for the roller due to its high Young’s modulus and hardness.
The roller-workpiece contact was set as frictionless due to the free rotation of the roller. Furthermore, the temperature effects i.e. material softening and thermal expansion and strain rate hardening were neglected for the low burnishing velocity.
The average burnishing forces as shown in Table 1 and a horizontal velocity constraint were exerted on the roller simultaneously. Then the residual stress could be obtained after the unloading of the roller.
4. Results and discussion
In this section, the experimental measured and FEM simulated subsurface deformation and residual stress are analyzed.
4.1. Subsurface deformation
The DIC measured subsurface deformation fields during and after the burnishing process are given in Fig. 4 - Fig. 6.
a b
2000 μ m 2000 μ m
4 3.5 3 2.5 2 1.5 1 0.5 0 Fig. 4. DIC measured displacement Uy (μm) (a) during the burnishing process and (b) after the burnishing process for set burnishing depth d = 50
μm and specific normal force F = 60.5 N/mm
a b
6 5 4 3 2 1 0
2000 μ m 2000 μ m
Fig. 5. DIC measured displacement Uy (μm) (a) during the burnishing process and (b) after the burnishing process for set burnishing depth d = 100
μm and specific burnishing force F = 131.25 N/mm
a b
12 10 8 6 4 2 0
2000 μ m 2000 μ m
Fig. 6. DIC measured displacement Uy (μm) (a) during the burnishing process and (b) after the burnishing process for set burnishing depth d = 150
μm and specific normal force F = 190 N/mm
It could be found from the experimental measurements that the subsurface deformation was well captured using the DIC technique and it increases with the increasing burnishing forces.
The residual subsurface deformation is commonly less than that during the burnishing process due to the unloading of the burnishing force.
For comparison, the FEM simulated subsurface deformation fields of the side surface are given in Fig. 7 - Fig. 9 with the same color bar as the experimental measurements.
a 2000 μm b 2000 μm
Fig. 7. FEM simulated displacement Uy (μm) (a) during the burnishing process and (b) after the burnishing process for specific normal force F = 64
N/mm
a 2000 μm b 2000 μm
Fig. 8. FEM simulated displacement Uy (μm) (a) during the burnishing process and (b) after the burnishing process for specific normal force F =
134.25 N/mm
a 2000 μm b 2000 μm
Fig. 9. FEM simulated displacement Uy (μm) (a) during the burnishing process and (b) after the burnishing process for specific normal force F =
188.75 N/mm
It is found that the FEM simulated displacement Uy both during and after the burnishing process is less than the measurements. Besides, comparing the measured and simulated deformations during the burnishing process, it could be found that the shapes of measured data are more broad and flattened, which may be attributed to the neglect of the roller- workpiece friction in the FEM simulation.
For a detailed comparison of the residual subsurface
deformation during the burnishing, line interpolations of the data along the depth direction at both the middle and side surfaces for the FEM simulation and the side surface for the measurements were conducted and the results are given in Fig.
10.
a
b
c
Depth (μm)
Depth (μm)
Depth (μm) (μm)(μm) (μm)
Fig. 10. The depth distributed residual subsurface deformation with different specific normal force F (N/mm). ‘I’ and ‘O’ stand for middle and
side surface, resp.
The residual subsurface deformation distribution was fitted using the following equation
exp d
y a b
æ ö÷
= çççè- ÷÷ø (2)
where a and b are the fitting parameters in μm, d is the depth in μm. a represents the deformation at the surface and b indicates the influence depth of the burnishing process. Nonlinear least square fitting was conducted and the results are given in Table 3.
Table 3. Fitting parameters of the residual subsurface deformation.
Designation Source Surface F (N/mm) a (μm) b (μm)
FEM_I_F64 FEM Middle 64 0.16 1600
FEM_O_F64 FEM Side 64 0.85 1094
EXP_O_F60.5 EXP Side 60.5 3.8 613.5
FEM_I_F134.25 FEM Middle 134.25 1.13 298.4
FEM_O_F134.25 FEM Side 134.25 2.98 1002
EXP_O_F131.25 EXP Side 131.25 6.12 963.8 FEM_I_F188.75 FEM Middle 188.75 2.45 378.2
FEM_O_F188.75 FEM Side 188.75 4.49 1039
EXP_O_F190 EXP Side 190 12.63 1137
It could be seen that the b value of the experimental
measurements increases with the increasing burnishing forces.
However, it stays almost the same for the side surface of the FEM simulations. Comparing the subsurface deformation on the middle and the side surfaces of the FEM simulations, it could be found that the middle surface deforms less than the side surface, indicating a 3D FEM simulation is needed to predict the subsurface on the side surface.
4.2. Residual stress
The comparisons of the FEM and experimental measured residual stress are shown in Fig. 11. For the FEM simulation results, the depth distributed residual stress was extracted on the middle plane but at two different locations. The average values with mean differences are given in this figure. Two experimental tests were conducted for the same set burnishing depth, and the error bar indicates the fitting error generated by the analysis software. Moreover, the two measurements for the same set burnishing depth were averaged and the results are also shown in the figure.
a
b
c
Fig. 11. The FEM simulated and measured depth distributed residual stress with the specific normal forces F (N/mm) in the legends.
It could be seen that both the FEM simulations and measurements produce hook-shaped distributed residual stresses. The measurements are about 80 MPa less than the simulation.
From the FEM simulation results, it could be seen that the surface residual stress tends to become tensile as the increasing the burnishing forces. However, the peak compressive residual
stress stays steady after the burnishing force increases to 134.25 N/mm. The depth of the peak compressive residual stress increases as the increasing burnishing forces.
From the average values of the experimental measurements, it is found that the peak compressive residual stress stays around -125 MPa for all the three set burnishing depths.
Moreover, it is also found that the depth of the peak compressive residual stress increases as the increasing burnishing force. Besides, the residual stress in the bulk workpiece is around -40 to -60 MPa, indicating the initial residual stress was not fully relieved by the heat treatment described in Section 2.1. With the unrelieved initial compressive stress, material yielding could occur in advance under the compressive loading of the roller. This could explain why the measured deformations are greater than the simulations. Moreover, the initial compressive residual stress in the bulk workpiece could also change the residual stress distribution in the subsurface according to the equilibrium requirement.
5. Conclusions and outlook
In this paper, the subsurface deformation and residual stress during the roller burnishing process were experimentally and numerically studied. It is found that the DIC successfully captured the subsurface deformation during and after the burnishing process. For the experimentally measured subsurface deformation, the influence depth increases with the increasing burnishing forces. However, it stays almost the same for the side surface of the FEM simulations. Besides, the FEM simulated subsurface deformations are less than the measured ones.
Both the FEM simulations and measurements produce hook-shaped distributed residual stresses and the measurements are about 80 MPa less than the simulations. It is confirmed by both the FEM simulations and the experiments that the depth of the peak compressive residual stress increases as the increasing burnishing force.
In the future, the relief of the initial stress should be improved and the correlation between the subsurface deformation and the residual stress will be studied.
Acknowledgments
This work was partially supported by the National Natural Science Foundation of China (Grant Nos. 51722505 and 51721092).
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