Effects of cutting parameters on surface residual stress and its mechanism in high-speed milling of TB6

Abstract

Formation mechanism and influence of cutting parameters on residual stress in flank milling of TB6 are investigated through orthogonal experiments. The thermal–mechanical coupling effect on forming residual stress is studied by analyzing the microstructure of metamorphic layer. Experiment results show that cutting temperature varies between 300 °C and 518 °C under experiment conditions, while residual compressive stress on machined surface is −199.8 to −41.7 MPa. Tensile residual stress occurs on subsurface when increasing milling parameters excessively. When cutting TB6 with parameters of f_z = 0.02 mm/z, v_c = 60 m/min, a_e = 0.2 mm and a_p = 10 mm, only the compressive residual stress can be detected, which is good for its fatigue life. The depth of compressive residual stress layer is 10–20 µm.

Keywords

TB6 titanium alloy high-speed milling milling force milling temperature residual stress

Introduction

The TB6 titanium alloy is widely used in aerospace industries because of its high specific strength, small anisotropy, low forging temperature and strong resistance to stress corrosion. When using TB6 instead of equal strength 20CrMnSiA and TC4 in aircraft fuselage, wing and landing gear structures, the weight of structural parts can be reduced about 40% and 20%, respectively. So TB6 titanium alloy has a broad application space in aerospace industries.¹

Residual stress on machined surface has an important effect on its static strength, fatigue strength and stress corrosion cracking. Residual stress and its instability will cause stress relaxation and redistribution in machining of thin-walled parts, which will cause unpredictable part deformation. So it is of great importance to research the formation mechanism of surface residual stress. The nature, value, distribution and formation mechanism of residual stress get more and more attention of scholars in various countries. Scholars did active study on surface integrity of high-speed milling titanium alloy. Residual stress is an important index of surface integrity. Jawahir et al.² investigated the surface/subsurface properties involving hardness and residual stress profiles. They concluded that milling processes carry the potential to induce high compressive stress with high depth effect. Thomas et al.³ studied the microstructure damage behavior of high-speed milling titanium alloy TC4. They found that the near-surface microstructure has a direct effect on slip band characteristics of alpha phase. The surface integrity of high-speed milled γ-TiAl alloy was studied by Mantle and Aspinwall.⁴ The depth of hardened layer reaches 300 µm, and the contained compressive stress under the surface is less than 500 MPa. Milling speed has great influence on residual stress. Sridhar et al.⁵ found that material properties and process parameters can affect residual stress on machined surface, which can be alleviated or eliminated by conducting appropriate heat treatment processes. Che-Haron⁶ and Che-Haron and Jawaid⁷ studied the tool life, tool wear, surface roughness, microhardness and metallurgical alterations systematically when cutting TC4. Sun and Guo⁸ investigated the surface integrity of end milling TC4. Compressive residual normal stress can be detected in both cutting and feed directions. Cutting speed and feed speed have different effects on residual stress. Microstructure analysis showed that surface layer grain becomes smaller and distortion appears when the cutting speed is increasing. Puerta Velasqueza et al.⁹ found that the crystallographic texture has negligible effect on residual stress value. The residual stress is tensile in nature for high cutting speed while compressive for lower speed. Du et al.¹⁰ and Yang et al.^11,12 investigated the effects of cutting parameters on surface integrity in high-speed milling. They found that the higher the spindle speed increases from 3000 to 10,000 r/min, the smoother the machined surface becomes. Chen et al.¹³ found that compressive residual stress appears on machined surface when down milling of titanium alloy TC4 due to the burnishing effect between flank and machined surfaces. They also compared the different effects of cutting parameters on residual stress in cutting and feeding directions. Yu et al.¹⁴ found that a higher compressive residual stress can improve fatigue life when stress ratio is constant. When stress ratio increases gradually, the effect of compressive residual stress on fatigue life decreases accordingly. But almost no effect can be found when stress ratio increases to 0.6. John et al.¹⁵ found that 80% of residual stress is reserved in the workpiece when the specimen is exposed in a thermal environment of 399 °C for 100 h. When fatigue loading is close to yield stress, reserved residual stress is 85%, 70% and 50% when the exposing temperature is 260 °C, 316 °C and 399 °C, respectively. Rao et al.¹⁶ focused on the experimental and numerical study of face milling of Ti–6Al–4V titanium alloy. They found that subsurface stress is compressive in nature and the depth extends to around 40 µm below the machined surface. Daymin et al.¹⁷ concluded that mean compressive stress decreases slightly at a higher workpiece angle due to the absence of rubbing effect.

There is little research on the thermal–mechanical coupling effect on residual stress. Current retrieved research focuses on α+β-phase titanium alloy TC4. But the β-phase titanium alloy TB6 is still rarely studied. TB6 gets an increasing application to reduce aircraft weight, which requires more research focus on surface integrity of components about this material. Surface quality and machining accuracy are mainly determined by finish machining parameters. So the objective of this study is to analyze the effect of milling parameters on milling force, temperature and residual stress and to reveal formation mechanism of residual stress when finish machining TB6. It could be a theoretical basis for surface integrity control and parameter optimization to cut TB6 in high-speed milling.

Plastic bulge deformation effect and surface burnishing effect in milling

Residual stress appears when material is elasto plastic anisotropic or inhomogeneous plastic deformation generation. Residual stress produced by cutting process is related to many factors, such as milling force, thermal load, microstructure transformation and so on. As shown in Figure 1, the tensile and compressive residual stresses on machined surface are induced by plastic bulge deformation effect and surface burnishing effect, respectively.

Figure 1.

Plastic bulge deformation effect and surface burnishing effect.

The plastic bulge deformation effect is the combined effect of thermal load and milling force in the contact region between the cutter and workpiece. On the one hand, compressive plastic deformation occurs due to the constraints of internal material, while surface material is expanding and temperature is rising under a thermal load. Plastic bulge deformation occurs because surface material cannot turn back freely when workpiece surface is cooling down. On the other hand, plastic shrinkage along machined surface and plastic expansion along the vertical direction of machined surface under tension forces T and T′ make another form of plastic bulge deformation effect in the region front of contact points on cutting edge. Obviously, plastic bulge deformation effect in this region is most significant all over the workpiece. It is caused by the combined effect of mechanical and thermal deformations. The surface burnishing effect is generated only by the physical force occurring between the cutting tool and workpiece. Compressive residual stress can be produced in cutting process under the burnishing effect of friction force F and burnishing force P, which is perpendicular to the surface. The final residual stress depends on the balance of plastic bulge deformation effect and burnishing effect.

Cutting experiments are designed and conducted in this research to reveal the influence of cutting parameters on cutting force and temperature. The thermal–mechanical coupling effect on residual stress formation is also investigated. Symbols adopted in this research are listed in Table 1.

Table 1.

List of symbols.

v_c	Milling speed	F_x	Milling force in x-direction
f_z	Feed per tooth	F_y	Milling force in y-direction
a_e	Milling width	F_z	Milling force in z-direction
a_p	Milling depth	σ	Residual stress
v_f	Feed rate	σ_x	Residual stress in x-direction
n	Spindle speed	σ_z	Residual stress in z-direction
T	Milling temperature	h	Distance below surface

Experiment material and procedure

Workpiece material

The workpiece material used in all experiments is TB6 alloy (Ti–10V–2Fe–3Al), one kind of β-phase titanium alloy with high strength and high toughness. Chemical composition of TB6 is shown in Table 2. Mechanical properties in normal and high temperatures are shown in Table 3.

Table 2.

Chemical composition of TB6 (mass fraction, %).

Element	H	O	N	Fe	Al	V	Ti
Wt%	0.01	0.03	0.03	1.93	2.95	10.15	Balance

Table 3.

Mechanical properties of TB6.

Temperature	σ_b (MPa)	σ_0.2 (MPa)	δ (%)	Ψ (%)	a_k (J/cm²)
Normal	1145	1082	11.1	64.4	60
350 °C	963		16.8	70.8

The measurement of milling force and temperature in high-speed milling

A semi-artificial thermocouple composed by TB6 and constantan (Φ = 0.3 mm) is used to measure milling temperature. It has small heat inertia and a wide temperature measurement range. Linear expansion coefficient of TB6 titanium is 10.2 × 10⁻⁶ when temperature increases from 20 °C to 500 °C. Temperature error measured with this method has little influence on thermal stresses. The fabricating procedure of a TB6–constantan thermocouple, as shown in Figure 2, can be briefly described as follows: machine two pieces of TB6 specimen with the size of 20 mm × 15 mm × 6 mm; carve a groove with the width and depth both close to 0.3 mm (the diameter of constantan wire) on one surface of one specimen; squeeze the constantan wire with insulation paint layer about 0.01 mm into the groove; coat a layer of glue on one surface for another specimen; bind two specimens together to make the constantan wire be clamped by two workpieces; stay this sample being clamped by vise moderately for 8–12 h in room temperature; and get off the sample from vise and check the insulativity of constantan wire and uncarved TB6 workpiece. This TB6–constantan thermocouple is fixed on a computer numerical control (CNC) machine worktable with vise to measure the cutting temperature.

Figure 2.

Temperature measurement system of semi-artificial thermocouple.

Because TB6–constantan thermocouple is a nonstandard thermocouple, a calibration needs to be conducted to get the correspondence between temperature and thermoelectric potential. A comparative method is adopted to calibrate the TB6–constantan thermocouple, which is described as follows: place a standard S type thermocouple and this TB6–constantan thermocouple in a same temperature environment; change the temperature and record different thermoelectric potentials of each thermocouple at some sample times; transfer thermoelectric potentials of standard thermocouple to actual temperatures directly with a temperature controller, so the actual temperature at these sample times can be exactly obtained; and find the correlation curve (calibration curve) between actual temperatures and thermoelectric potentials of TB6–constantan thermocouple. This calibration curve, as shown in Figure 3, can be used to predict cutting temperature by measuring thermoelectric potential in cutting experiments.

Figure 3.

Calibration curve of TB6–constantan thermocouple.

The calibration curve is fitted to be a polynomial

T = 42.1 + 30.3 t_{e} - 0.69 {t_{e}}^{2} + 0.0099 {t_{e}}^{3}

(1)

where T is the temperature of hot junction, while t_e is the corresponding thermoelectric potential. The determination coefficient R² of fitting result is 0.99985, standard deviation (SD) is 2.62755, p < 0.0001. So the calibration curve is acceptable and reliable. This TB6–constantan thermocouple can be used to measure the cutting temperature in experiments.

Orthogonal experiments with three factors and three levels are designed and conducted to get the cutting force, temperature and residual stress. The cutting experiments are carried out with emulsified liquid in JOHNFORD VMC 850. The cutting tool is a 4-flute uncoated cemented carbide end mill. Rake angle, relief angle, helix angle and diameter of this tool are 12°, 12°, 35° and 10 mm, respectively. Three levels of milling speed, feed per tooth and milling width are selected to compare the different influence when the milling depth a_p is kept constant at 10 mm. Milling parameters are given in Table 4.

Table 4.

Measurement results of milling force and temperature for milling TB6 titanium alloy.

No.	f_z (mm/z)	v_c (m/min)	a_e (mm)	F_x (N)	F_y (N)	F_z (N)	T (°C)
1#	0.02	60	0.2	54	67	25	300
2#	0.02	100	0.4	90	117	48	401
3#	0.02	140	0.6	93	159	71	431
4#	0.06	60	0.6	200	232	114	377
5#	0.06	140	0.4	235	290	138	510
6#	0.06	100	0.2	180	263	70	456
7#	0.1	60	0.4	250	237	122	409
8#	0.1	100	0.6	395	455	231	518
9#	0.1	140	0.2	285	289	63	476

The measurement system of cutting force consists of a Kistler 9255B dynamic piezoelectric dynamometer, a Kistler 5017A charge amplifier and a recorder. Milling forces in three directions are represented as F_x, F_y and F_z, respectively, as shown in Figure 5. Mean force by averaging 50 peak values of continuous transient cutting force is used as the experiment result of cutting force, as shown in Table 4.

The measurement system of cutting temperature is constructed by a semi-artificial TB6–constantan thermocouple, an A11S11 signal disposal module, a UA305 high-speed USB data collector and a computer, as shown in Figure 2. Potential signals are recorded during milling experiments. Figure 4 shows the potential signal of No. 1#. It can be transformed to an actual milling temperature with the assistance of the calibration curve obtained above. The milling temperature of No. 1# is about 300 °C as a mean value of several tests. Milling temperatures of nine samples can be obtained with same method, as shown in Table 4.

Figure 4.

Potential signal of first milling.

Specimen preparation, measurement of residual stress and microstructure

In order to investigate the influence of milling parameters on residual stress, another nine specimens are prepared. The size of specimen is 55 mm × 50 mm × 30 mm, as shown in Figure 5. Surface residual stresses measured in experiments are listed in Table 5.

Figure 5.

Schematic diagram of flank milling.

Table 5.

Surface residual stresses measured in experiments.

No.	f_z (mm/z)	v_c (m/min)	a_e (mm)	σ_x (MPa)	σ_z (MPa)
1#	0.02	60	0.2	−64.4	−121.5
2#	0.02	100	0.4	−95.7	−138.7
3#	0.02	140	0.6	−167.1	−41.7
4#	0.06	60	0.6	−73.2	−120.5
5#	0.06	140	0.4	−103.4	−134.1
6#	0.06	100	0.2	−182.6	−67.2
7#	0.1	60	0.4	−50.9	−90.8
8#	0.1	100	0.6	−59.4	−119.1
9#	0.1	140	0.2	−196.1	−199.8

Residual stresses are measured in XStress 3000 with Cu-Kα radiation using the Ψ-tilt X-ray method. In fact, the measured result is residual strain. But it can be transferred to residual stress after a calculation based on Hooke’s law. When residual stress exists in the sample, the change of interatomic spacing causes a shift in the diffracted X-ray peak position. An X-ray with wavelength of $λ$ radiates to the sample at several different incidence angles. Measure the corresponding diffraction angle $2 θ$ and find the slope of $2 θ$ to $\sin 2 ψ$ . By resolving the angular peak shift and applying the Bragg law to quantify the interatomic spacing, residual stress on the machined surface can be calculated using the theory of linear elasticity. A mean value of three points along the feed direction on machined surface is adopted as the result of σ_x and σ_z. Electrochemical method is used to delaminate specimens to measure residual stress on subsurface. Each electrochemical processing is 15 s with a delaminating depth of 510 µm. Repeat this process until the residual stress measured is close to the body material. The residual stress measurement parameters are listed in Table 6.

Table 6.

The residual stress measurement parameters.

Material	TITANIUM(ALPHA)2 (HCP, hkl-213)		Units	MPa
Tube	Cu-Kα		Bragg angle	139.69
Wavelength	1.542		Peak location	Gaussian 80%
KV; mA	21.00; 30.00		Gain correction	P/G(s)
Beta Osc	5.00		PeakShift	Absolute peak
Phi Osc	0.00		(1/2)S2	11.89E−6 (1/MPa)
X Osc	0.00		−S1	2.83E−6 (1/MPa)
Y Osc	0.00		DSpacing (Angstroms)	0.8212065
Exp. time	2.00	2.00	Lorentz-polarization and absorption (LPA) correction	Yes
Nr of exp	10	10	Psi zero assignment (pixel)	262.72	255.80
Background static fit
Left start pixel	130	130	Left stop pixel	140	140
Right start pixel	379	379	Right stop pixel	389	389
ROI left	132	131	ROI right	388	388

ROI: region of interest.

The microstructure of specimen’s crossing section was observed using a DMI 5000M inverted optical microscope. The metallographic test sample is made by cutting the machined specimen along feed direction. The cutting section is corroded after abrading and polishing in a corrodent matched by HNO₃:HF:H₂O = 1:1:10.

Formation mechanism of residual stress

Effect of milling parameters on surface residual stress

1. Effect of milling parameters on residual stress

Mathematical model of residual stress is built using multiple linear regression analysis based on the data in Table 5

{\begin{matrix} σ_{x} = 0.5 & f_{z}^{- 0.08} v_{c}^{1.03} a_{e}^{- 0.39} \\ σ_{z} = 214.6 & f_{z}^{0.22} v_{c}^{- 0.07} a_{e}^{- 0.27} \end{matrix}

(2)

With this mathematic model, the changing sensitivity of residual stress by the change of cutting parameters can be obtained. σ_x is most sensitive to milling speed, while least sensitive to feed per tooth. σ_z is most sensitive to milling width, while least sensitive to milling speed.

The effect of milling parameters on surface residual stress is studied by extremum difference analysis. As shown in Figure 6, compressive residual stress is observed in both directions on machined surface. The effect of feed per tooth on surface residual stress is shown in Figure 6(a). It can be seen that σ_x increases with increase of f_z from 0.02 to 0.06 mm/z and decreases when f_z is from 0.06 to 0.10 mm/z. σ_z keeps decreasing when f_z increases from 0.02 to 0.10 mm/z. σ_z is a little larger than σ_x at a relatively low feed per tooth. The changing range of residual stress in both directions is within [−140 MPa, −100 MPa]. As for the effect of v_c in Figure 6(b), residual stress almost increases linearly with increase of v_c in both directions. This is mainly because higher milling speed leads to a higher temperature that tends to induce thermal plastic deformation. The effect of v_c on residual stress is greater than that of f_z when comparing Figure 6(b) with Figure 6(a). Figure 6(c) shows that residual stress decreases with increase of a_e, and the changing range of residual stress in both directions is [−150 MPa, −80 MPa]. The effect of a_e on residual stress is between that of v_c and f_z.

Figure 6.

Effect of milling parameters on surface residual stress. (a) Effect of f_z on residual stress, (b) effect of v_c on residual stress and (c) effect of a_e on residual stress.

2. Interaction effects of milling parameters on residual stress

In order to choose a better combination of milling parameters, the interaction effect of milling parameters on residual stress must be investigated. The interaction effect of f_z and v_c on residual stress is shown in Figure 7. Figure 7(a) shows that σ_x higher than −180 MPa appears in the region of “high milling speed and low feed per tooth” (region A). This is because a large thermal plastic deformation is prone to occur near the surface layer of workpiece under this milling condition. σ_x lower than −100 MPa appears in the region of “low milling speed and high feed per tooth” (region C). The thermal plastic deformation is small under this cutting condition. The range of σ_x under other experiment cutting conditions, as signed with region B, is [−100 MPa, −180 MPa]. Figure 7(b) shows that σ_z higher than −140 MPa appears in the region of “low milling speed and high feed per tooth” (region D). This is because the increase of feed per tooth induces higher milling temperature that tends to produce a large thermal plastic deformation on machined surface. σ_z lower than −110 MPa appears in the region of “high milling speed and low feed per tooth” (region F). The range of σ_z under other experiment cutting conditions, as signed with region E, is [−110 MPa, −140 MPa].

Figure 7.

The interaction effect of f_z and v_c on residual stress. (a) Interaction effect of f_z and v_c on σ_x and (b) interaction effect of f_z and v_c on σ_z.

The interaction effect of a_e and v_c on residual stress is shown in Figure 8. Figure 8(a) shows that σ_x higher than −160 MPa appears in the region of “high milling speed and low milling width” (region G). σ_x lower than −100 MPa appears in the region of “low milling speed and high milling width” (region I). The range of σ_x under other experiment cutting conditions, as signed with region H, is [−100 MPa, −160 MPa]. Figure 8(b) shows that σ_z higher than −100 MPa appears in the region of “low milling speed and low milling width” (region J). σ_z lower than −80 MPa appears in the region of “high milling speed and high milling width” (region L).

Figure 8.

The interaction effect of a_e and v_c on residual stress. (a) Interaction effect of a_e and v_c on σ_x and (b) interaction effect of a_e and v_c on σ_z.

The interaction effect of a_e and f_z on residual stress is shown in Figure 9. Figure 9(a) shows that σ_x higher than −190 MPa appears in the region of “low milling width and low feed per tooth” (region M). σ_x lower than −130 MPa appears in the region of “high milling width and high feed per tooth” (region O). The range of σ_x under other experiment cutting conditions, as signed with region N, is [−130 MPa, −190 MPa]. Figure 9(b) shows that σ_z higher than −120 MPa appears in the region of “low milling width and high feed per tooth” (region P). This is because the increase of feed per tooth induces a large thermal plastic deformation on machined surface. σ_z lower than −90 MPa appears in the region of “high milling width and low feed per tooth” (region R). The range of σ_z under other experiment cutting conditions, as signed with region Q, is [−90 MPa, −120 MPa].

Figure 9.

The interaction effect of a_e and f_z on residual stress: (a) Interaction effect of a_e and f_z on σ_x and (b) interaction effect of a_e and f_z on σ_z.

3. Selection of milling speed and feed per tooth

Compressive residual stresses with the range of [−199.8 MPa, −41.7 MPa] appear on all machined surfaces typically. When surface roughness and tool wear are under cutting requirements, milling parameters should be selected properly to obtain a higher compressive residual stress. Figure 6 shows that the changing range of σ_x is larger than that of σ_z. This means milling parameters should be selected mainly based on σ_x, which is easier to be affected by milling parameters. As demonstrated in equation (2), milling speed has the most significant influence on σ_x, while feed per tooth has the least. So cutting speed should be selected first as high as possible. After the cutting speed has been selected, a higher feed per tooth is recommended to improve machining efficiency.

Effect of milling parameters on milling force and milling temperature

1. Milling force and temperature model based on milling parameters

Mathematical model of milling force and temperature is built using multiple linear regression analysis based on the data in Table 4, as shown in equations (3) and (4), respectively

{\begin{matrix} F_{x} = 623.2 & f_{z}^{0.86} v_{c}^{0.35} a_{e}^{0.29} \\ F_{y} = 204.2 & f_{z}^{0.69} v_{c}^{0.53} a_{e}^{0.35} \\ F_{z} = 451.3 & f_{z}^{0.65} v_{c}^{0.25} a_{e}^{0.86} \end{matrix}

(3)

As shown in equation (3), feed per tooth and milling speed both have a significant effect on milling force. The effect of milling width on milling force is insignificant under experiment conditions

T = 155.6 f_{z}^{0.14} v_{c}^{0.33} a_{e}^{0.08}

(4)

Milling temperature is most sensitive to the variation of milling speed and least sensitive to milling width. Table 4 shows that the range of milling forces F_x, F_y, and F_z is [54 N, 395 N], [67 N, 455 N] and [25 N, 231 N], respectively. The effect of milling parameters on milling force and milling temperature is studied by extremum difference analysis. The result is shown in Figure 10.

Figure 10.

Effect of milling parameters on milling force and temperature. (a) Effect of f_z on milling force and temperature, (b) effect of v_c on milling force and temperature and (c) effect of a_e on milling force and temperature.

2. Effect of milling parameters on milling force

The effect of feed per tooth on average milling force is shown in Figure 10(a). It can be seen that milling forces increase with increase of f_z in three directions. As shown in Figure 10(b), milling force increases with increase of v_c from 60 to 100 m/min and decreases when v_c is from 100 to 140 m/min. This may be explained as follows. First, milling temperature increases with the increase of milling speed. The increasing temperature decreases friction coefficient and material deformation coefficient, which finally leads to a decrease of milling force. Second, with increase of milling speed, shear angle increases and shear plane becomes smaller. The shear force decreases by the change of shear plane under a constant shear strength. The increase of high-frequency impact force is less than the decrease of milling force at the meantime. As a consequence, the milling force shows a decreasing tendency. As shown in Figure 10(c), when a_e changes from 0.2 to 0.6 mm, average milling force only increases slightly. The range of milling forces F_x, F_y and F_z are [173 N, 229 N], [200 N, 280 N] and [53 N, 139 N], respectively.

3. Effect of milling parameters on milling temperature

Figure 10 shows that average temperatures increase with the increase of all milling parameters. The range of milling temperature is about [300 °C, 518 °C]. Such a high milling temperature causes the following effects. On the one hand, due to the uneven distribution of temperature on workpiece surface, thermal stress appears on machined surface when the subsurface restricts the expansion of workpiece volume. When thermal stress exceeds the yield limit of material, compression plastic deformation appears on machined surface. Because surface material cannot turn back freely with the constraint of subsurface, tensile residual stress occurs when workpiece surface is cooling down. On the other hand, surface tempering temperature is close to or even more than the critical temperature to lead the phase transition. Residual stress is produced due to the volume change of microstructure.

Formation mechanism of residual stress on machined surface

In order to improve fatigue life of part, the consideration of residual stress value on machined surface is not enough. Distribution of residual stresses on subsurface plays a more significant role on the fatigue life. Specimens machined with cutting parameters of No.1# and 9# are selected to study residual stresses distribution. Milling speed and feed per tooth in machining specimen 1# are both lower than that of specimen 9#. Surface roughness of specimens 1# and 9# are 0.36 and 0.91 µm, respectively.

The residual stress distribution on machined surface and subsurface of specimens 1# and 9# are shown in Figure 11, where h is the distance below machined surface. Figure 11(a) shows that compressive residual stresses appears in both directions on machined surface with a depth about 10~20 µm. σ_x and σ_z on machined surface are −70 and −120 MPa, respectively. Residual stress tends to be close to zero gradually with the increase of h. Figure 11(b) shows that σ_x and σ_z on surface are −200 and −190 MPa, respectively. The maximum tensile stress is about 100 MPa, which appears 10 µm beneath the machined surface. Tensile residual stress reduces gradually with increase of h. The depth of residual stress layer is about 30~40 µm.

Figure 11.

Distribution of residual stresses in subsurface. (a) Residual stresses distribution of 1# and (b) residual stresses distribution of 9#.

Figure 10 shows that milling temperature range is [360 °C, 470 °C], which does not exceed the phase transition temperature of 817 °C. The range of F_x and F_y is [75 N, 300 N] and [120 N, 330 N], respectively. This is because severe plastic deformation and shear effect induce separation grain bear tension, which leads to larger tensile stress in the first deformation zone. In the third deformation zone, tensile plastic deformation occurs under the severe friction and squeeze between cutting tool flank and workpiece. Small elastic modulus of titanium alloy induces larger rebound on machined surface that leads to strong surface burnishing effect on tool flank surface. Compared with the plastic bulge deformation effect by milling force and heat, surface burnishing effect by milling force is more significant. Both effects result in compressive residual stresses on machined surface.

Figure 11(a) shows that almost only compressive residual stress appears on subsurface. Figure 12(a) shows that plastic deformation near the machined surface is not obvious with a depth less than 10 µm. This is mainly because strain hardening on subsurface caused by thermal–mechanical coupling effect is not obvious when F_x, F_y and F_z and milling temperature are 54, 67 and 25 N and 300 °C, respectively. Figure 11(b) shows that the maximum tensile residual stress appears at h = 10 µm on specimen 9#. Figure 12(b) shows that plastic deformation of specimen 9# is larger than that of specimen 1# along the feeding direction. The depth of deteriorative layer is about 25 µm. This is mainly because strain hardening on subsurface is significant when F_x, F_y and F_z and milling temperature are 285, 289 and 63 N and 476 °C, respectively.

Figure 12.

Variation of surface microstructure: (a) specimen 1# (f_z = 0.02 mm/z, v_c = 60 m/s, a_e = 0.2 mm, a_p = 10 mm) and (b) specimen 9# (f_z = 0.1 mm/z, v_c = 100 m/s, a_e = 0.6 mm, a_p = 10 mm).

With the increase of milling parameters, milling force increases rapidly, while milling temperature increases slowly. Plastic deformation layer becomes deeper, and grain fibrosis increases aggravatingly along feed direction. The thermal–mechanical coupling effect leads to residual stress, which increases with increase of milling parameters. However, when milling parameters increases too largely, tensile residual stress appears on the subsurface. It is because the maximum of plastic deformation caused by thermal–mechanical coupling effect appears on the subsurface, which leads to large strain hardening effect in this region.

Conclusion

All machined surfaces perform compressive residual stress under experiment conditions. The range of surface compressive residual stress is [−199.8 MPa, −41.7 MPa]. Milling speed has the most significant influence on residual stress. With increase of milling parameters, surface compressive residual stress increases and tensile residual stress appears on the subsurface.

The range of milling temperature is [300 °C, 518 °C], which is below the phase transition temperature. With increase of milling parameters, milling force increases rapidly, while milling temperature increases slowly. Larger strain hardening effect on the subsurface is the main cause for tensile residual stress.

When f_z = 0.02 mm/z, v_c = 60 m/min, a_e = 0.2 mm and a_p = 10 mm, surface residual stress is about −100 MPa. Compressive residual stress appears on machined surface and subsurface. The depth of compressive residual stress layer is 1020 µm.

Footnotes

Funding

The authors wish to thank the support of the National Natural Science Foundation of China (grant numbers 50975237 and 51005184) and National Science and Technology Major Project of the Ministry of Science and Technology of China (grant number 2013ZX04011031) and graduate starting seed fund of Northwestern Polytechnical University (grant number Z2013048).

References

Wang

Meng

Liao

. Present state of research on Ti-1023 titanium alloy. Dev Appl Mater 2009; 24: 66–69 (in Chinese).

Jawahir

Brinksmeier

M’Saoubi

. Surface integrity in material removal processes: recent advances. CIRP Ann Manuf Techn 2011; 60: 603–626.

Thomas

Turner

Jackson

. Microstructural damage during high-speed milling of titanium alloys. Scripta Mater 2010; 62: 250–253.

Mantle

Aspinwall

. Surface integrity of a high speed milled gamma titanium aluminide. J Mater Process Tech 2001; 118: 143–150.

Sridhar

Devananda

Ramachandra

. Effect of machining parameters and heat treatment on the residual stress distribution in titanium alloy IMI-834. J Mater Process Tech 2003; 139: 628–634.

Che-Haron

. Tool life and surface integrity in turning titanium alloy. J Mater Process Tech 2001; 118: 231–237.

Che-Haron

Jawaid

. The effect of machining on surface integrity of titanium alloy Ti-6%Al-4%V. J Mater Process Tech 2005; 166: 188–192.

Sun

Guo

. A comprehensive experimental study on surface integrity by end milling Ti–6Al–4V. J Mater Process Tech 2009; 209: 4036–4042.

Puerta

Velasqueza JD

Tidua

Bollea

. Sub-surface and surface analysis of high speed machined Ti-6Al-4V alloy. Mat Sci Eng A: Struct 2010; 527: 2572–2578.

10.

Ren

. Study on surface morphology and microstructure of titanium alloy TC4 under high-speed milling. Acta Aeronautica et Astronautica Sinica 2008; 29: 1710–1715 (in Chinese).

11.

Yang

Zhang

Yao

. Effects of high-speed milling parameters on surface integrity of TC4 titanium alloy. J Northwest Polytech Univ 2009; 27: 538–543 (in Chinese).

12.

Yang

Zhang

Yao

. Influence mechanism of milling speed on TC4 titanium alloy surface integrity in high speed milling. J Nan Jing Univ Aeronaut Astronaut 2009; 41: 644–648 (in Chinese).

13.

Chen

Sun

. Surface residual stress of titanium alloy induced by milling. J Mech Strength 2010; 32: 53–57 (in Chinese).

14.

Zheng

. Effects of residual stresses on the high cycle fatigue behavior of Ti-6Al-4V. In: ASME Conf. Proc. / ASME 2010 pressure vessels and piping conference, Bellevue, Washington, USA, July 18–22, 2010, volume 5

15.

John

Buchanan

Jha

. Stability of shot-peen residual stresses in an α+β titanium alloy. Scripta Mater 2009; 61: 343–346.

16.

Rao

Dandekar

Shin

. An experimental and numerical study on the face milling of Ti–6Al–4V alloy: tool performance and surface integrity. J Mater Process Tech 2011; 211: 294–304.

17.

Daymin

Boujelbene

Ben

Amara A

. Surface integrity in high speed end milling of titanium alloy Ti-6Al-4V. Mater Sci Tech Ser 2011; 27: 387–394.