- Email: [email protected]
Geometrical texture and surface integrity in helical milling and ultrasonic vibration helical milling of Ti-6Al-4V alloy
Journal Pre-proof Geometrical texture and surface integrity in helical milling and ultrasonic vibration helical milling of Ti-6Al-4V alloy Guang Chen, Yunhe Zou, Xuda Qin, Jian Liu, Qiang Feng, Chengzu Ren
PII:
S0924-0136(19)30467-4
DOI:
https://doi.org/10.1016/j.jmatprotec.2019.116494
Reference:
PROTEC 116494
To appear in:
Journal of Materials Processing Tech.
Received Date:
20 April 2019
Revised Date:
30 July 2019
Accepted Date:
4 November 2019
Please cite this article as: Chen G, Zou Y, Qin X, Liu J, Feng Q, Ren C, Geometrical texture and surface integrity in helical milling and ultrasonic vibration helical milling of Ti-6Al-4V alloy, Journal of Materials Processing Tech. (2019), doi: https://doi.org/10.1016/j.jmatprotec.2019.116494
This is a PDF file of an article that has undergone enhancements after acceptance, such as the addition of a cover page and metadata, and formatting for readability, but it is not yet the definitive version of record. This version will undergo additional copyediting, typesetting and review before it is published in its final form, but we are providing this version to give early visibility of the article. Please note that, during the production process, errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain. © 2019 Published by Elsevier.
Geometrical texture and surface integrity in helical milling and ultrasonic vibration helical milling of Ti-6Al-4V alloy Guang Chena,b, Yunhe Zoua, Xuda Qina,b, Jian Liua, Qiang Fengc, Chengzu Rena,b* a
Key Laboratory of Mechanism Theory and Equipment Design of Ministry of Education, Tianjin University, Tianjin 300354,
China; b Tianjin
Key Laboratory of Equipment Design and Manufacturing Technology, Tianjin University, Tianjin 300354, China; c CNPC
Bohai Drilling Engineering Company Limited, Tianjin 300280, China
*Corresponding author: [email protected]
ro of
Abstract
Surface integrity of hole-making for Ti-6Al-4V alloy is very important due to its application in aviation industry. Geometrical and mechanical behaviors of machined holes by helical
-p
milling (HM) and ultrasonic vibration helical milling (UVHM) were investigated. The textures of machined holes by the two processes were analyzed and the distances of adjacent
re
textures were calculated according to the movement of peripheral edges. Due to the different surface geometrical textures, the hole-diameter error and surface roughness in UVHM are
lP
smaller than those of HM. A plastic deformation layer with thickness of 4-6µm was observed by the distorted β phase in the cross-section of hole-surface in UVHM by SEM results. Compared with HM process, UVHM generates larger hardness at the subsurface of machined
na
holes (20-200µm) due to the material deformation caused by ultrasonic vibration. Compared with HM, the compressive residual stress of hole surface was increased larger than 63.5% by UVHM. The compressive stresses in UVHM are larger than 150MPa at the depth less than
ur
10µm, which is consistent with the microstructure evolution caused by mechanical
Jo
deformation.
Keywords: Ti-6Al-4V alloy; Helical milling; Ultrasonic vibration; Surface integrity; Surface texture
1. Introduction Titanium alloys are widely applied in aviation industry owing to the high strength-to-weight 1
ratio and good corrosion resistance. However, due to the characteristics of high chemical reactivity, small thermal conductivity and low deformation coefficient, the machining of titanium alloys always causes serious tool wear. Therefore, titanium alloys normally belong to difficult-to-cut materials as described by Sharif and Rahim (2007). In the aviation industry, hole-making process is normally the ultimate and critical stage in manufacturing that contributes to 40%-60% of material removing during aircraft assembly as demonstrated by Schroeder (1998). The assembling holes produce concentrated stress region that initiates and propagates fatigue cracks. Therefore, surface integrity of machined holes is important for the fatigue life of the aircraft structures, and good surface integrity can ensure
ro of
high reliability and good performance of the aircraft as described by Jawahir et al. (2011). The most widely used hole-making method for titanium alloys is conventional drilling, and many investigations about the performance of this process have been reported. Che-Haron and Jawaid (2005) presented that white layer was formed during rough machining of Ti-6Al-4V alloy. Cantero et al. (2005) reported the micro-hardness in the machined surfaces
-p
at the top 75μm thick layer was increased by 30% compared with the original surface hardness during drilling of Ti-6Al-4V alloy under dry condition. Additionally, it was found
re
that the increase of drilling time led to thicker affected zone. Zeilmann and Weingaertner (2006) measured the cutting temperatures of dry drilling and drilling with MQL for
lP
Ti-6Al-4V alloy using type K thermo-elements. The mechanical resistance of Ti-6Al-4V alloy decreased significantly and it was easier to produce plastic deformation at the temperature higher than 500ºC. Rahim and Sasahara (2010) assessed tool wear evolution,
na
surface integrity in drilling of Ti-6Al-4V alloy at dry and MQL lubricant conditions. The machined surface and subsurface generated deformation due to the effect of thermal softening, which made the micro-hardness at 0.025mm depth from the machined surface smaller than
ur
the average hardness of bulk material. In addition, it was found that the effect of strain hardening was increased with the increasing cutting speed and feed rate. Kuo et al. (2014)
Jo
investigated the influence of feed rate and cutting speed on workpiece surface integrity by single-shot drilling of Ti-6Al-4V/CFRP/Al7050 stacks. The average hole-surface roughness (Ra) of Ti-6Al-4V alloy was 0.60μm, and the feed rate had no obvious influence on the micro-hardness of the machined surface. Li et al. (2008) carried out metallographic observation and nano-indentation test of Ti-6Al-4V workpiece and chips with high-throughput drilling. A peak hardness of 9GPa located at subsurface layer was observed at 15-20μm depth from machined surface, while the bulk material hardness was 4-5 GPa. Shyha et al. (2011) studied the hole quality for drilling of Titanium/CFRP/aluminum stacks. 2
The roughness value (Ra) for Ti-6Al-4V alloy layer was less than 1μm. Additionally, adhered material on the Ti-6Al-4V alloy layer indicated the need of a post-finishing process for drilling the stacks. Micro-hardness results exhibited limited strain hardening until 200μm depth from machined surface along with microstructure deformation. As a relatively new technology, helical milling (HM) is a kind of hole-making process by a milling tool moving along helical path. Compared with the traditional drilling, helical milling has advantages of high accuracy, low axial force, and intermittent cutting at the periphery cut edge as demonstrated by Tian et al. (2017). Therefore, this technology drew more attention due to the potential application in the aerospace industry for assembling manufacture in
ro of
recent years. Kihlman et al. (2002) realized automate orbital drilling on assembly structure using a standard industrial robots in the aerospace industry. Orbital drilling generated lower axial cutting forces by the application of industrial robots than conventional drilling within the required hole tolerances. He et al. (2015) carried out hole-making of CFRP/Ti-6Al-4V alloy stacks by HM with various machining parameters. Lower cutting forces, higher
-p
geometrical accuracy, and better surface finishing (Ra≤0.58 μm for Ti-6Al-4V alloy) were achieved, and the need of reaming or de-burring was eliminated. Olvera et al. (2012)
re
machined holes with the average surface roughness of 1μm by traditional drilling method, while, ball helical milling was capable of producing holes with 0.6μm average surface
lP
roughness for Ti-6Al-4V alloy. Additionally, the thermal softening effect was identified to counter the main strain hardening during helical milling when the cutting speed became higher. Sun et al. (2018) investigated the surface roughness, microstructure evolution and
na
micro-hardness of surface and subsurface for aluminum alloy and Ti-6Al-4V alloy using helical milling and drilling processes. The machined parts by HM had longer (46%) fatigue life than that machined by drilling, as less plastic deformation, lower surface roughness, and
ur
compressive residual stress for the hole surface were machined by helical milling. Similarly, Rasti et al. (2019) compared the micro-hardness along the radial and axial directions in the
Jo
hole-making of 4340 steel by drilling, helical milling, profile milling with and without predrilling. Results indicated that the subsurface (within 100µm) hardness of all experiments were higher than bulk material regardless of cutting methods. Compared with the surface machined by HM, the surface machined by conventional drilling generated thicker white layer, and that was mainly caused by the progressive feed of cutting tool. Zhao et al. (2015) investigated surface integrity for titanium alloy with traditional drilling and HM processes. It was reported that the micro-hardness beneath the machined surface gradually decreased to bulk hardness when the depth was about 200μm. Meanwhile, the residual stress in helical 3
milling behaved as compressive stress. Li et al. (2014) evaluated the tool wear in HM machining of Titanium alloy with TiAlN tool. It was reported that the tool wear is an important factor which affects the hole-making quality for Ti-6Al-4V alloy. The wear rates at the frontal cut edge and the periphery cut edge were analyzed. The severe thermal damage is the main factor that causes the tool fracture, especially at tool nose and the front cutting edge. Additionally, a phenomenon of smearing which may cause cracks and other defects on the hole surface with the worn tool was observed. Therefore, the tool-workpiece contact behavior affects the states of cutting tool and surface integrity of hole surface.
ro of
In order to make cutting edges separate from workpiece material intermittently, one feasible method is applying ultrasonic vibration in helical milling process. Nath and Rahman (2008) imposed vibration with a high frequency to cutting tools and developed an Ultrasonic vibration machining (UVM) manufacturing method. Owing to the superiority of intermittent cutting mechanism, the UVM process can reduce the cutting force; improve heat dissipation,
-p
chip breakability and evacuation. Besides, ultrasonic vibration assisted cutting can remove the adhesive chip on the cutting tool, therefore, better hole-surface quality can be achieved.
re
Bai et al. (2019) applied ultrasonically assisted turning (UAT) in machining of SiCp/Al metal-composite. The cutting forces were reduced with the increase of cutting temperature.
lP
Zou et al. (2019) proposed an axial and torsional ultrasonic assisted helical milling (LTUAHM) for hole-making of Ti-6Al-4V alloy. It was reported that the elliptical vibration trajectory affects the cutting angle and the friction behavior between the rake face and chips.
na
Dvivedi and Kumar (2007) investigated the surface quality of pure titanium and titanium alloy by a process of ultrasonic drilling (USD). The surface roughness of USD was much better than the processes of electro discharge machining (EDM) and laser beam machining
ur
(LBM). It was reported that USD process did not produce crack and recast layer which is benefit to machine holes with straight profile. Pecat and Brinksmeier (2014) performed
Jo
low-frequency vibration assisted drilling (LFVAD) on Ti-6Al-4V alloy/CFRP stacks. A remarkable temperature fall of approximately 43% was observed in comparison with traditional drilling, because of the tool cooling related to the interrupted cutting process. Surface integrity and residual stress in ultrasonic-assisted milling of AISI 316L was investigated by Maurotto and Wickramarachchi (2016). It was found that surface integrity and residual pressure varied within a narrow range at different frequencies. Additionally, the minimum value of residual compressive stress was measured at vibration frequency of 40 KHz. Similarly, Suárez et al. (2016) presented the fatigue life and surface integrity of 4
Ni-alloy 718 machined by ultrasonic vibration assisted milling. Surface hardness of workpieces machined by ultrasonic vibration milling was 3.79% higher than those machined with traditional milling methods, and the fatigue life was also increased by 14.74% compared with those machined with traditional milling methods. Ishida et al. (2014) applied ultrasonic vibration in hole-making of CFRP by helical milling process which resulted in reduction of thrust force. The accuracy of machined hole was also improved due to the reduction of delamination size at the exit of the holes. Recently, ultrasonic vibration helical milling (UVHM) method was proposed in our previous work in machining of Ti-6Al-4V alloy as presented by Chen et al. (2019). The
ro of
material removal mechanism was analyzed by the kinematic modeling of cutting trajectories of the bottom and peripheral cutting edges. Meanwhile, the axial force reduction was modeled and the maximum force reduction is 64% compared with those of HM. Additionally, the surface roughness of hole-surface at different conditions and the residual stress at one condition were measured and analyzed by the material removal mechanism. Although the
-p
material removal mechanism of hole machining by UVHM has been studied, the hole-making quality which is crucial to aviation application was not clear yet. Additionally, Chen et al.
re
(2019) mainly focused on the material removal related with tool and tool-workpiece contact behavior. This work focuses on the geometrical texture and surface integrity for machined
lP
hole, concerning the micro-scale material deformation and surface integrity of Ti-6Al-4V alloy by UVHM and HM. The geometrical accuracy, material micro-structure evolution at the
na
surface and subsurface, micro-hardness and residual stress will be investigated. 2. Movement of cutting edge in UVHM The schematic of tool movement in ultrasonic vibration helical milling (UVHM) is shown
ur
in Fig. 1. The milling tool rotates with the machine spindle, and feeds along helical feed path. Meanwhile, the tool vibrates with a high frequency along the axial direction.
Jo
In UVHM, the movement of cutting edges can be characterized in the XYZ coordinate
system. To characterize the mechanism of material removal in the UVHM, the previous work in Chen et al. (2019) investigated the trajectories of the peripheral and bottom cutting edges. The trajectory of specific point in the peripheral cutting edge can be calculated by
5
nrevt Dt nrot t Dh Dt x 2 cos 30 2 cos 30 n t Dh Dt nrevt Dt sin rot y sin 30 2 30 2 z vfa t A sin(2 ft )
(1)
where Dh is the diameter of the target hole; Dt is the diameter of the tool; nrot is the rotation speed of spindle; nrev is orbital revolution speed; vfa is the velocity of axial feed and f is the
re
-p
ro of
frequency of vibration.
lP
Fig.1 Movement of the milling tool in UVHM
To indicate the material removal in the hole-surface during UVHM, the schematic of movement of the peripheral cutting edge and hole-surface is shown in Fig. 2. MN (in Fig. 2a)
na
is one of the peripheral cutting edges of the milling tool. The hole-surface can be analyzed in the coordinate of sOZ. While, Os is the circumferential direction of machined hole; the cutting edge will transform to a short line M1N1, as shown in Fig. 2b. In the machining
ur
process, due to the self-rotation and the vibration along the direction of OZ, the cutting edge M1N1will move with periodic path in the coordinate of sOZ. In addition, the cutting edge can
Jo
separate from the chip periodically at specific process conditions (Chen et al., 2019). In coordinate of XOY, the cutting edge will leave a circular arc M2N2 (Fig. 2c). As the diameter of the tool and hole are different, the arc M2N2 has different curvature compared with the target machined hole, which will affect the geometrical morphology of hole-surface, that is, surface roughness, hole-diameter and roundness.
6
Fig. 2 Relationship of peripheral cutting edge and material removal of hole-surface in UVHM (a) XYZ coordinate (b) sOZ coordinate and (c) XOY coordinate
ro of
3. Experimental procedure Helical milling experiments for Ti-6Al-4V alloy with/without vibration (HM and UVHM) were proposed to study the surface integrity of hole-surface. The experimental setup of UVHM is shown in Fig. 3. The workpiece is machined with dimension of
-p
258mm×125mm×5mm. The experiment was carried out on a machining center DMG®-60 mono BLOCK. Four edges cemented carbides milling tools with composite of WC-12%Co
re
were applied in the HM and UVHM experiments. The ɸ6mm milling tool is used to machine through-holes with nominal diameter of 10mm on workpiece with 5mm thickness. The
lP
detailed geometrical parameters of the milling tool are listed in Table 1. Table 1 Geometrical parameters of the milling tool Rake angle of bottom edge (º)
Rake angle of peripheral edge (º)
Clearance angle of bottom edge (º)
Clearance angle of peripheral edge (º)
Axial length of peripheral edge (mm)
Tool nose radius (mm)
38
0
5
8
10
7
0.5
ur
na
Helix angle (º)
The equipment used to generate the ultrasonic vibration consists of ultrasonic power,
Jo
upper ring with electromagnetic induction coil and a tool holder with bottom ring. A non-contact upper ring electromagnetic unit fixed on the spindle shaft is utilized to transform the electrical signal. The bottom ring with electromagnetic induction coil is installed on the tool holder. When the tool holder is installed on the spindle, there is approximately 1mm distance between the upper ring and bottom ring. Ultrasonic power (type: SZ12) is connected to the upper ring which can transfer the electrical signal to the bottom ring by electromagnetic transmission. Therefore, the tool holder can receive the electrical signal and drive the piezoelectric ceramic generating high frequency axial vibration. The vibration 7
frequency is approximately 34 KHz and the vibration amplitude is approximately 4-6 μm. The cutting forces were measured during the machining process with a three-component dynamometer (Kistler® 9527A) and a charge amplifier (Type: 5070). The signals of cutting forces were measured with a sampling frequency of 20 KHz. The HM and UVHM processes
-p
ro of
were established under dry conditions. The processing parameters are listed in Table 2.
re
Fig. 3 Experimental setup for Ti-6Al-4V alloy machining using HM and UVHM
Test No.
lP
Table 2 Cutting conditions for hole-making of Ti-6Al-4V alloy by HM and UVHM Spindle
speed
Axial feed
Tangential feed
ap(mm/rev)
ft(mm/tooth)
na
nrot (rpm) 2000
0.15
0.03
2
2500
0.15
0.03
3
3000
0.15
0.03
4
3500
0.15
0.03
5
4000
0.15
0.03
6
2500
0.15
0.02
7
2500
0.15
0.04
8
2500
0.15
0.05
9
2500
0.15
0.06
10
2500
0.1
0.03
11
2500
0.2
0.03
Jo
ur
1
8
12
2500
0.25
0.03
13
2500
0.3
0.03
To study the geometry morphology and surface integrity of machined holes using HM and UVHM, surface morphology was measured by a ZYGOTM (type: 7300) scanning white light interferometer. The hole-diameter errors were measured by Coordinate Measuring Machining (CMM) system. The wire electro discharge machining (WEDM) was utilized to section the machined holes in circumferential and axial directions to measure the surface and subsurface microstructure for the machined holes. The samples were mounted and ground by
ro of
abrasive papers from 400 to 2000 grit to remove the surface material. Afterwards, the mounted samples were polished using a Tegramin-25 polishing system for microstructural observation of the cross-section of machined holes. Hardness values at different depths from the machined surface were tested by a MHV digital micro-hardness Tester (type: 2000) using
-p
load of 100mN with 10s duration. The residual stresses were measured by μ-X360n® X ray diffractometer. Cr tube was used to determine the elastic strains for the {1 0 3} diffraction planes of the HCP crystal structure for Ti-6Al-4V alloy. The X-Ray tube voltage and current
re
are 30kV and 1.0mA, respectively. The diffraction peak of X-Ray at the Bragg angle of 140º was measured using a spot with diameter of 1 mm. The Young’s modulus and Poisson’s ratio
lP
are 115.7 GPa and 0.321, respectively. The machined holes were sectioned across and along the axial direction by WEDM before tests, and the axial and circumferential residual stresses at different depths from machined surface were measured. Similar residual tress
na
measurements for machined hole surface were presented in Kwong et al. (2009). Successive layer of material was moved away by electro-polishing to prevent the surface from
ur
regenerating additional residual stresses in order to obtain stress values at different depths from machined surface. Material relaxation during the layer electro-polishing (5~50µm for each layer) was neglected due to the smaller depth removed relative to the thick bulk material
Jo
(3~4mm).
4. Hole-surface topography, surface integrity and tool wear 4.1 Topography of machined Hole-surface The surface topography of machined hole surfaces at four conditions (No.2, No.5, No.11 and No.13.) is shown in Fig. 4. The selected surface area is 0.11 mm× 0.11 mm. The surface 9
ro of -p re lP na ur Jo Fig. 4 Comparison of surface topography of machined hole-surfaces by HM and UVHM at tests of No.2 (a, a′), No.5 (b, b′), No.11 (c, c') and No.13 (d, d'), where a-d represent the results in HM; a'-d' represent the results in UVHM 10
roughness was measured with seven linear positions along circumferential direction (Fig. 4a), and the mean Ra value was presented in each condition. Note that the mean surface roughness of hole surface in UVHM is smaller than that machined in HM when the machining parameters (in Table 2) are identical. For example, the Ramean in No.2' (UVHM) is 0.232 μm, while the value in No.2 (HM) is 0.386μm. Similar phenomena can be observed for Tests No.5, No.11 and No.13. As shown in Fig. 4a-d, periodic convex peaks are observed in the surface generated by HM. Compared with the surface of HM, the surface in UVHM generates periodic convex textures with smaller distance, as shown in Figs. 4a'-d'. In addition, the convex textures appear an
ro of
angle with the axial direction, which is generated by the helix angle of peripheral cutting edges. Compared with HM, the peripheral cutting edge in UVHM generates periodic vibration along the axial direction, as illustrated in Fig. 2. The vibration of peripheral edge will generate a combined effect of friction and compression in UVHM, which can reduce the larger peaks generated in HM. Therefore, the periodic convex textures of hole-surface
-p
indicate micro-scale plastic deformation and it affects the geometrical accuracy of machined hole and may affect the physical behaviors (hardness, residual stresses) and tribological
re
properties of hole-surface.
For the texture machined in HM, the periodic surface curve is mainly caused by the tangential
Jo
ur
na
is presented in Fig. 5.
lP
feed of cutting edges. The schematic of the generation of periodic curves on the hole-surface
Fig. 5 Schematic of curves generated at hole-surface by peripheral cutting edges
During the helical milling feed process, the peripheral cutting edges will leave convex curves on the hole-surface which is generated by adjacent cutting edges, as shown in Figs.4a-d. With the self-rotation and orbital rotation, the points of tangency for the milling tool and the hole surface along the peripheral edge will move along an incline line due to the orbital feed. In 11
another word, the separation points along the peripheral cutting edge will leave a vertical line on the hole-surface if the tool just does self-rotation. Due to the orbital revolution, the separation points for the same peripheral cutting edge will leave an incline line on the machined hole-surface, as shown in Figs.4a-d. The distance of adjacent curves can be analyzed by the machined curves generated by the peripheral edges on the XOY coordinate
ur
na
lP
re
-p
ro of
system.
Jo
Fig. 6 Machined texture generated by peripheral cutting edges (a) and (b) formation of curve A0A1A2A3A4 by same peripheral edge (c) repeated curves generated by adjacent edges (d) the textures generated at the bottom of machined hole
In HM, each cutting edge rotates with an angle speed of 𝜔2 along the clockwise direction, as shown in Fig. 6a. The center of cutting tool rotates along the counter clockwise direction with an angle speed of 𝜔1. As the cutting edge O0A0 rotates to the position of O1A1and O2A2, a trajectory A0A1A2 is formed by the end of bottom cutting edge represented by the solid lines 12
(Fig. 6a). After identical time interval, the cutting edge will rotate from O2A2 to the positions of O3A3and O4A4 continuously and, the cutting edge will leave a trajectory A2A3A4, as shown in Fig. 6b. In this process, the point in the peripheral edge will form the cutting trajectory A0A1A2A3A4 in the XOY coordinate. Therefore, the adjacent cutting edges will leave repeated trajectories with identical shape, as shown in Fig. 6c. Due to the repeated machined trajectories, the machined hole-surface will generate surface texture like BDAEC, which is identical with the experimental textures leaved at the bottom hole surface (Fig. 6d). Therefore, the distance of machined curves (Δl1) in the hole-surface of HM process (arc DE in Fig. 6) can be calculated by the adjacent points of tangency of the peripheral edge (arc AB in Fig. 6)
lP
re
-p
ro of
and the hole-surface, such as the arc of A1B2 in Fig. 7.
na
Fig. 7 Calculation of the distance of the adjacent convex textures
In Fig. 7, O1A1 and OA1 are in the same line, that is, A1 is the point of tangency between the
ur
tool and the hole surface. After time interval of Δt, the adjacent cutting edge O1B1 rotates an angle (γ) to O2B2, where B2 is the next tangency point between the tool and hole-surface, that is, O2B2 will leave another curve on the hole-surface due to the orbital feed, as shown in Fig.
Jo
6c. During time interval of Δt, each cutting edge rotates an angle of γ, at the mean time, the center of milling tool rotates an angle of θ (Fig. 7). Therefore, distance of adjacent curves on the hole-surface can be calculated according to the geometrical relationship described in Chen et al. (2019) as
2π / Z e 2 t t 1 13
(2)
where Ze=4 is the number of cutting edges. Then, the distance of arc A1B2 (Δl1 in Fig. 4d) can be calculated by
l1`
2π1 (e r ) Z e (1 2 )
(3)
where e equals to the length of OO1 (Fig. 7), r is the diameter of the cutting tool. In addition, the distance of adjacent curves (such as Δl2 in Fig. 4d') generated by UVHM (Fig.4a'-4d') can be calculated according to Chen et al. (2019) by the instantaneous speed of peripheral edge along the circumferential direction as t
s vt dt
(5)
-p
v v 2 v 2 2v v cos 1 2 1 2 t v1 1e v2 2 r 2πn 60 rev 1 2 2πnrot 60
ro of
(4)
0
(6)
lP
re
3 π 2 π 3 + 2 1 2 1 1t 2 2t
na
However, as the speed of orbital revolution is much lower than that of self-rotation speed (ω1<<ω2), the linear speed of the peripheral edge along the trajectory of A1A2A3 can be considered as vt≈ω2r=2πnrotr/60. Therefore, the distance of curves in UVHM (Fig. 4d') can
Jo
ur
be calculated as
l2
2 r
(7)
f
where f is the actual frequency in UVHM. Then, the calculated distances of adjacent curves in Fig. 4 are shown in Table 3. It can be seen that the distances of adjacent textures machined by UVHM are smaller than those of HM. The ratios of the distances in HM and UVHM are 3.186 (for 2500rpm) and 2.059 (for 4000rpm), respectively. Therefore, it was mostly affected by the tool rotation speed. In addition, the calculated distances in HM for Test Nos. 2, 5, 11 and 13 are identical (Δl1=74.328 µm), while, the calculated distances in UVHM for Test Nos. 2, 5, 11 and 13 have little difference which is caused by the variation of vibration frequency. 14
The calculated results agree with the measured morphology of machined hole-surface in Fig. 4.
Table 3 Calculated distances of adjacent textures in the machined hole-surface Experimental
Distance of adjacent
Distance
of
adjacent
conditions
curves in HM Δl1(µm)
curves in UVHM
Ratio Δl1/Δl2
Δl2(µm) 74.327901
23.327687
3.186252464
No. 5
74.327901
36.094028
2.05928526
No. 11
74.327901
23.356144
3.186252464
No. 13
74.327901
23.307601
3.186252464
ro of
No. 2
4.2 Hole-diameter error
-p
Hole-diameter error presents the deviation of the actual diameter and nominal diameter which is related with the geometrical morphology of hole-surface. The coordinates of 8 points on
re
the hole-surface at 1 mm depth from the entry and outlet of the machined holes were applied to evaluate the hole-diameter error. The hole-diameter errors near the entry and outlet of hole-surfaces are presented in Fig. 8. The test positions are also illustrated in Fig. 8b. In this
lP
work, the hole-diameter error represents the relative error between the measured diameter and the target diameter (ϕ10mm), and the measured values were the average errors of the repeated experiments. It can be seen that the hole-diameter errors are less than 0.03mm for both HM
na
and UVHM. The values of entry area are larger than those of outlet area for most conditions. In addition, the hole-diameter error decreases with the increasing cutting speed at both of
ur
entry and outlet areas. Similarly, with increase of tangential feed (from 0.02 to 0.06mm/r), the hole-diameter errors near entry are mostly positive and decrease gradually with tangential
Jo
feed. While, the hole-diameter errors near outlet area are mostly negative and also decrease with the increasing tangential feed. Additionally, for most conditions, the hole-diameter errors of UVHM are slightly smaller than those of HM, especially for the hole-diameter errors at outlet area. Note that the hole-diameter errors were tested on the circumferential direction (as indicated in Fig. 8b), and it should be related with the convex textures formed along the circumferential direction, as shown in Fig. 4c. As discussed in section 4.1, the surface convex textures in HM are generated by the movement of peripheral edges. In UVHM, a rubbing effect was caused 15
during the vibration of peripheral edges, which will affect the geometrical morphology and thermal-mechanical behavior of machined surface. The geometrical morphology can be characterized by the distances of convex textures by Eqs. (3) and (7). In HM, the tangential feed per tooth fzt is related with the tool self-rotation nrot and orbital feed nrev.
2πenrev f zt Ze nrot
(8)
Thus, the distance of adjacent textures in HM can be calculated based on Eqs.(3)(5)(8),
2π(e r ) 2 e Z e (1 ) Z e f zt
ro of
(9)
Jo
ur
na
lP
re
-p
l1`
Fig. 8 Comparison of hole-diameter errors in HM and UVHM at different (a) spindle speeds, (b) tangential feed per tooth and (c) axial feeds.
Therefore, the adjacent distance of the textures in HM (Δl1) will increase with the increasing tangential feed per tooth (fzt). Geometrically, the dimension of Δh will increase with the adjacent distance Δl1, which will lead to decrease of the hole-diameter. This agrees with the measured results in Fig. 8b. With the increase of cutting speed from 2000rpm to 4000rpm, the 16
tangential feed per tooth is kept constant, the hole-diameter decreased slightly with the spindle speed (Fig. 8a). That should be related with the thermal deformation, because the cutting temperature normally increases with instantaneous cutting speed. In addition, when the axial feed increases, the rotation speed and tangential feed keep constant, the hole-diameter remains stable due to the same Δl1 and similar heat generation. Therefore, the surface topography for machined hole affects the geometrical accuracy. To further indicate the variation of hole diameter along the axial direction, the hole diameters at different depths (1-4mm) along axial direction were presented in Fig. 9. Both of the hole diameters machined by HM and UVHM decrease with the increasing axial depth, from 1mm to 4mm along axial direction. In detail, when the depth along the axial direction increases
ro of
from 1mm to 4mm, the hole diameter in UVHM decreases from ɸ9.994mm to ɸ9.982mm,
lP
re
-p
while the values in HM decreases from ɸ9.979mm to ɸ9.959mm.
na
Fig. 9 Variation of hole diameters along axial direction for the condition of Test No.9 (nrot=2500rpm, ft=0.06mm/tooth, ap=0.15mm/rev)
ur
The decrease of hole diameter along the axial direction is mainly related with two factors. Firstly, it is related with the tool deflection which is caused by the resultant forces due to the
Jo
tangential feed. The tool deflection may accumulate with the increase of axial feed due to the support from machined hole surface. Secondly, with the increase of axial feed, the cutting temperature near the cutting zone accumulates and the thermal elastic deformation of the workpiece will increase during the cutting process. Therefore, the hole diameter will change accordingly when the workpiece is cooled down after machining. Similar with the results of the hole diameter errors in Fig. 8, the hole diameters machined by UVHM are closer to ɸ10mm than those machined by HM. Although the hole diameter decreases with the 17
increasing depth along axial direction, both of the hole diameters in HM and UVHM are close to target value, and the largest error is approximately 0.04mm. 4.3 Microstructure of sub-surface in HM and UVHM Fig. 10 shows comparison of cross-sectional SEM images of surface and subsurface microstructure produced by HM and UVHM under the machining conditions of No.1 (a, a′ and b, b′), No.2 (c and c′) and No.5 (d and d′). As shown in Fig.10a-10d, there is no obvious distorted β phase on the subsurface in HM, that is, no obvious plastic deformation occurred in HM process along the axial direction. This phenomenon is similar to the microstructure
ro of
results presented by Sun et al. (2018). In UVHM, distorted β phase are observed under the machined hole-surface. As illustrated in Fig. 10a′-d′, the distorted β phase mainly occurs between the hole-surface and a line “L”. The β phase does not exhibit deformation in the area below the line “L”. Therefore, a plastic deformation layer with thickness of approximately
-p
4-6μm exist between the surface and the line “L” in UVHM.
In addition, it is worth noting that the periodic concave textures can be observed on the hole surface in UVHM process, as indicated in Fig. 10a′-d′. The periodic concave texture should
re
be caused by the impulsive sinusoidal cutting trajectory which is related with vibration. Additionally, the periodic surface textures are also observed in the surface topography along
lP
the axial direction, similar with the results in Fig. 4.
Note that the morphology of top surface in SEM graph represents the topology along the axial edges (in Fig. 4c and 4c'). As presented in Fig. 4c, the speed of peripheral edge in HM is
na
along the circumferential direction, which means the instantaneous cutting speed of the points in the peripheral cutting edge is normal to the top surfaces of SEM results in HM. Therefore,
ur
the microstructure of subsurface did not exhibit plastic deformation (distorted β grains). In UVHM, due to the axial vibration, the machined hole-surface generates a thin plastic deformation layer. It should be noted that the instantaneous speed direction of peripheral edge
Jo
has an incline angle with the axial edge (in Fig. 4c'). That is, the distorted grains in Fig. 10a'-d' just represent the deformation state of axial direction, but not the largest deformation direction. The microstructure evolution on the surface and subsurface will contributed to physical-mechanical properties of the hole-surface, i.e., micro-hardness and residual stress states.
18
ro of -p re lP na ur Jo Fig. 10 Comparison of cross-section SEM images of surface microstructure at conditions of No.1 (a, a′ and b, b′), No.2 (c and c') and No.5 (d and d'), where a-d represent the results in HM; a'-d' represent the results in UVHM. 19
4.4 Micro-hardness of machined hole-surface Micro-hardness is normally determined by the coupling effect of mechanical and thermal loads. When the titanium alloy is subjected to high temperature and high pressure during the dry cutting process, a competing effect of work hardening and thermal softening exists and finally leads to specific thermal-mechanical behaviors of the machined surface. This phenomenon can be characterized by the cutting force, temperature and micro-hardness, which are caused by plastic deformation, friction and tool-workpiece contact compression, etc. To investigate the mechanical behaviors of machined hole-surface, the micro-hardness of
ro of
subsurface from 20μm to 250μm depth from the machined surface was tested, as illustrated in Fig. 11. The micro-hardness at each depth was measured three times and the average value
lP
re
-p
was considered as the hardness for different depths from machined surface.
na
Fig. 11 Measured indentation beneath the machined surface at test No.1 (nrot=2000rpm, ft=0.03mm/tooth, ap=0.15mm/rev)
ur
Fig. 12 presents measured micro-hardness of machined subsurface in HM and UVHM at the conditions of Test Nos.1, 2, and 5. The bulk hardness is approximately 333-343HV. The
Jo
micro-hardness in HM and UVHM processes at tests Nos. 1, 2, and 5 exhibit work hardening effects from 0-200µm depth from machined surface, which means the hardness of machined surface or subsurface is higher than that of the bulk material. For example, the hardness of UVHM at 40µm depth from machined surface is 363HV in Test No.1, while, the harnesses at depth larger than 150µm are close to the bulk hardness. However, due to various processing parameters, the largest hardness appeared at different depths for the given conditions. The largest hardness of UVHM with rotation speed of 2500rpm is 362HV which located at 60µm depth. That should be related to the processing parameters (cutting speed, tangential and axial 20
feed rates) and also related with the contact behavior of tool-hole surface. As illustrated in Fig. 4, the material removal and tool-hole surface contact behavior are not identical along the circumferential direction. Therefore, the surface hardness may behave a little fluctuation for different conditions. In addition, compared with the subsurface hardness in HM, the hardness in UVHM at depth less than 200µm is enhanced. For example, for Test No. 1, the hardness of UVHM at 20-80μm depths is approximately 10HV larger than those of HM. Besides, for Test No. 2, the hardness of UVHM at 20-200μm depths is 20-30HV larger than those of HM,
ur
na
lP
re
-p
ro of
indicating a hardness rise by 5.9%-9.1%.
Fig. 12 Comparison of micro-hardness at (a) Test No.1 (nrot=2000rpm, ft=0.03mm/tooth, ap=0.15mm/rev); (b) Test No.2 (nrot=2500rpm, ft=0.03mm/tooth, ap=0.15mm/rev); (c) Test No.5 (nrot=4000rpm,
Jo
ft=0.03mm/tooth, ap=0.15mm/rev).
The increase of subsurface hardness should be caused by the plastic deformation at the machined surface and subsurface which can be observed in the metallurgical behavior of the machined subsurface in Fig. 10. The deformation layer in the top machined surface in UVHM will result in more work hardening effect and higher hardness than HM process. Chen et al. (2019) reported that the vibration can generate rubbing and friction effects to the surface of 21
machined holes. In addition, the vibration leads to distorted β phase in the deformation layer (Figs. 10a'-d') and cause more work hardening effect in UVHM. Besides, it was reported the instantaneous highest cutting speed of peripheral edges in UVHM was much higher than that in HM, which can cause more strain and strain rate hardening in UVHM, increasing the hardness of machined surface and subsurface. Similar work (or strain) hardening phenomena were also reported in machining of titanium and other light-weight alloys due to the effects of rubbing, burnishing and the cutting-extrusion in milling by Yao et al. (2013) and vibration assisted burnishing proposed by Teimouri et al. (2018).
ro of
4.5 Residual stress of hole-surface in HM and UVHM Residual stress of machined surface and subsurface is normally produced by the coupled effects of the thermal-mechanical loading. In general, tensile residual stress normally involves higher cutting temperature and intensive plastic bulge deformation. Sun and Guo (2009) stated that the compressive residual stress was dominated by mechanical deformation.
-p
Compared with conventional drilling, helical milling process can remove chips by larger hole space and dissipate cutting heat. Therefore, the mechanical loading is predominant in HM,
according to Zhao et al. (2015).
re
and the compressive residual stress was observed for machining of Ti-6Al-4V alloy by HM
lP
To further investigate the mechanical behavior of machined hole-surface by HM and UVHM, the residual stresses along the circumferential and axial directions of two conditions (Test No. 2 and No.5) are illustrated in Fig. 13. Note that the residual stresses of both HM and UVHM
na
processes generate compressive stresses which are beneficial to enhance the fatigue strength
Jo
ur
of hole-surface.
Fig. 13 Comparison of residual stress at (a) Test No.2 (nrot=2500rpm, ft=0.03mm/tooth, ap=0.15mm/rev); and (b) Test No.5 (nrot =4000rpm, ft=0.03mm/tooth, ap=0.15mm/rev). 22
For test No. 2, the residual stress of hole surface machined by HM along axial direction is -127MPa, whereas the stresses in UVHM along axial direction is -216MPa. While, the residual stress of hole surface machined by HM along circumferential direction is -104MPa, whereas the stresses in UVHM along circumferential direction is -170MPa. That is, compared with HM, UVHM increases the compressive surface residual stress by 70.1% and 63.5% for axial and circumferential directions, respectively. Similarly, for test No.5, the compressive residual stress of hole-surface along the axial direction is increased by 70.2%, while the stress along the circumferential direction is increased by 172.1%. Note that most of the compressive
ro of
stresses which are larger than 150MPa in UVHM located at the top layer with depth smaller than 10µm. That is related to the grain deformation caused by mechanical deformation in UVHM, as the thickness of top deformation layer (Fig. 10) in UVHM is approximately 4-6μm. Besides, comparing the residual stress with the rotation speed of 2500 and 4000rpm, the rotation speed has little influence on residual stresses in HM and UVHM. Due to the
-p
plastic deformation caused by ultrasonic vibration, the UVHM process can generate higher
4.6 Tool wear in HM and UVHM
re
compressive residual stress which can increase the fatigue strength of machined hole-surface.
lP
Tool wear is an important factor which affects the performance of machining quality and machining cost. In this work, to quantify the wear of the HM and UVHM processes, the average flank wear (VB) was measured to characterize the tool wear in HM and UVHM at
na
the condition of Test No. 2, as shown in Fig. 14. During the experiments, new tools were applied in HM and UVHM. As the axial feed is 0.15mm/rev, the length of peripheral edge
ur
involved in cutting is much smaller than that of the bottom cutting edge. Then, the average flank wear of bottom cutting edge was measured by microscope when each five holes were machined. As shown in Fig. 14, the tool wear in HM and UVHM increase with similar trends.
Jo
That is, tool wear increases until 10 holes, and then remains stable until 25 holes, and the tool wear increases rapidly with the hole numbers until 40. Additionally, the flank wears in UVHM are more rapidly than those in HM. When 40 holes was machined, the average flank wears are approximately 65µm and 100µm in HM and UVHM, respectively. The speed of tool wear should be related with the instantaneous cutting speed. According to the study of Li et al. (2014), tool wear increases when higher cutting speed was employed. In UVHM, due to the high frequency axial vibration, the instantaneous speed of bottom edge 23
generates periodical sinusoidal fluctuation. The highest instantaneous of bottom cutting edge is much higher than that in HM. Besides, the material removal in HM is continuous, while the cutting of bottom edge in UVHM was intermitted due to the vibration, as stated by Chen et al. (2019). The intermitted cutting in UVHM will change the thickness of uncut chip thickness and will generate impact effect between the flank surface of bottom edge and the workpiece. That is, the highest instantaneous cutting speed, changeable uncut chip thickness and impact effect can lead to higher flank tool wear in UVHM. Additionally, the average flank wear in HM and UVHM are far less than 200µm when 40 holes were machined. That is, during the surface integrity investigation in HM and UVHM (i.e., 13 holes in HM and 13 holes in
ro of
UVHM) experiments, the tool wear should be in the stable stage, that is, less than 70µm as
lP
re
-p
indicated in Fig. 14.
ur
5. Conclusions
na
Fig. 14 Flank wear versus machined hole number with HM and UVHM at Test No.2 (nrot=2500rpm, ft=0.03mm/tooth, ap=0.15mm/rev)
In this paper, comparative study for geometrical topography and surface integrity of Ti-6Al-4V alloy by helical milling (HM) and ultrasonic vibration helical milling (UVHM)
Jo
was proposed. The surface textures of machined surfaces in HM and UVHM were analyzed. The geometrical and physical behaviors of machined holes, including the hole-diameter error, micro-hardness, microstructure evolution and residual stresses in two different processes were presented. The main conclusions are summarized as follow: 1. The distance of adjacent textures of machined surface in HM and UVHM were calculated. The adjacent textures in HM are larger than those generated by UVHM. The textures in HM were related to the tangential feed of peripheral edges, while, the 24
textures in UVHM were caused by the helical feed and the ultrasonic vibration along the axial direction. 2. The hole-diameter errors in UVHM are smaller than those in HM. The hole-diameter error is related with the textures along the circumferential direction. The hole-diameter value decreases with the increasing tangential feed per tooth. 3. Due to the movement of peripheral cutting edges, approximately 4-6µm deformation layers with distorted β phase were observed by the cross-section of SEM images in UVHM, however, no deformation layer along axial direction was generated in HM. 4. Compared with HM, UVHM generates more work hardening effect at the subsurface of
ro of
machined holes (20-200µm). The vibration load generates more plastic deformation at the surface and subsurface of machined hole, increasing the work hardening effect. The micro-hardness increases with the spindle speed, tangential and axial feeds.
5. The compressive residual stress of hole-surface was increased larger than 63.5% by UVHM for given conditions compared with HM process. The compressive stresses that
-p
are larger than 150MPa in UVHM located at the top layer with depth smaller than 10µm, which is consistent with the microstructure evolution caused by mechanical
re
deformation in UVHM, as illustrated by SEM images.
6. The tool wear in UVHM is larger than that in HM, due to the highest instantaneous
lP
cutting speed, variable uncut chip thickness and the impact effect between the cutting edge and machined surface.
na
Acknowledgements
The authors would like to acknowledge the financial support received from the National
ur
Natural Science Foundation of China (No. 51575384).
Jo
References
Bai, W., Roy, A., Sun, R., Silberschmidt, V.V., 2019. Enhanced machinability of SiC-reinforced metal-matrix composite with hybrid turning. J. Mater. Process. Technol. 268, 149-161. Cantero, J.L., Tardío, M.M., Canteli, J.A., Marcos, M., Miguélez, M.H., 2005. Dry drilling of alloy Ti–6Al–4V. Int. J. Mach. Tools. Manuf. 45 (11), 1246-1255. Che-Haron, C., Jawaid, A., 2005. The effect of machining on surface integrity of titanium 25
alloy Ti–6% Al–4% V. J. Mater. Process. Technol. 166 (2), 188-192. Chen, G., Ren, C., Zou, Y., Qin, X., Lu, L., Li, S., 2019. Mechanism for material removal in ultrasonic vibration helical milling of Ti-6Al-4V alloy. Int. J. Mach. Tools. Manuf. 138, 1-13. Dvivedi, A., Kumar, P., 2007. Surface quality evaluation in ultrasonic drilling through the Taguchi technique. Int. J. Adv. Manuf. Technol. 34 (1-2), 131-140. He, G., Li, H., Jiang, Y., Qin, X., Zhang, X., Guan, Y., 2015. Helical milling of CFRP/Ti-6Al-4V stacks with varying machining parameters. Trans. Tianjin Univ. 21 (1), 56-63.
ro of
Ishida, T., Noma, K., Kakinuma, Y., Aoyama, T., Hamada, S., Ogawa, H., Higaino, T., 2014. Helical milling of carbon fiber reinforced plastics using ultrasonic vibration and liquid nitrogen. Procedia CIRP. 24, 13-18.
Jawahir, I.S., Brinksmeier, E., M'saoubi, R., Aspinwall, D., Outeiro, J., Meyer, D., Umbrello,
CIRP. Ann-Manuf. Technol. 60 (2), 603-626.
-p
D., Jayal, A., 2011. Surface integrity in material removal processes: Recent advances.
Kihlman, H., Eriksson, I., Ennis, M., 2002. Robotic orbital drilling of structures for aerospace
re
applications. SAE. Tech. Pap. No. 2002-01-2636.
Kuo, C.L., Soo, S., Aspinwall, D., Thomas, W., Bradley, S., Pearson, D., M'saoubi, R., Leahy,
lP
W., 2014. The effect of cutting speed and feed rate on hole-surface integrity in single-shot drilling of metallic-composite stacks. Procedia CIRP. 13, 405-410. Kwong, J., Axinte, D.A., Withers, P.J. 2009. The sensitivity of Ni-based superalloy to hole
na
making operations: Influence of process parameters on subsurface damage and residual stress. J. Mater. Process. Technol. 209(8), 3968-3977. Li, H., He, G., Qin, X., Wang, G., Lu, C., Gui, L., 2014. Tool wear and hole quality
ur
investigation in dry helical milling of Ti-6Al-4V alloy. Int. J. Adv. Manuf. Technol. 71 (5-8), 1511-1523.
Jo
Li, R., Riester, L., Watkins, T.R., Blau, P.J., Shih, A.J., 2008. Metallurgical analysis and nanoindentation characterization of Ti–6Al–4V workpiece and chips in high-throughput drilling. Mater. Sci. Eng. A. 472 (1-2), 115-124.
Maurotto, A., Wickramarachchi, C., 2016. Experimental investigations on effects of frequency in ultrasonically-assisted end-milling of AISI 316L: A feasibility study. Ultrasonics. 65, 113-120. Nath, C., Rahman, M., 2008. Effect of machining parameters in ultrasonic vibration cutting. Int. J. Mach. Tools. Manuf. 48 (9), 965-974. 26
Olvera, D., de Lacalle, L.N.L., Urbikain, G., Lamikiz, A., Rodal, P., Zamakona, I., 2012. Hole making using ball helical milling on titanium alloys. Mach. Sci. Technol. 16 (2), 173-188. Pecat, O., Brinksmeier, E., 2014. Tool wear analyses in low frequency vibration assisted drilling of CFRP/Ti6Al4V stack material. Procedia CIRP. 14, 142-147. Rahim, E.A., Sasahara, H., 2010. Effect of MQL liquids on surface integrity when high speed drilling titanium alloy. Key. Eng. Mater. 443, 359-364. Rahim, E.A., Sasahara, H., 2011. Investigation of tool wear and surface integrity on MQL machining of Ti-6AL-4V using biodegradable oil. Proc. Inst. Mech. Eng. Part B-J. Eng.
ro of
Manuf, 225(9), 1505-1511. Rasti, A., Sadeghi, M.H., Farshi, S.S., 2019. An investigation into the effect of surface integrity on the fatigue failure of AISI 4340 steel in different drilling strategies. Eng. Fail. Anal. 95, 66-81.
Schroeder, P., 1998. Widening Interest In Twist Drills: How wide-web drills with
-p
high-performance points can increase efficiency. Mod. Mach. Shop. 71, 106-113.
Sharif, S., Rahim, E., 2007. Performance of coated-and uncoated-carbide tools when drilling
re
titanium alloy–Ti–6Al4V. J. Mater. Process. Technol. 185 (1-3), 72-76. Shyha, I., Soo, S.L., Aspinwall, D., Bradley, S., Perry, R., Harden, P., Dawson, S., 2011. Hole
Manuf. 51 (7-8), 569-578.
lP
quality assessment following drilling of metallic-composite stacks. Int. J. Mach. Tools.
Suárez, A., Veiga, F., de Lacalle, L.N.L., Polvorosa, R., Lutze, S., Wretland, A., 2016. Effects
na
of ultrasonics-assisted face milling on surface integrity and fatigue life of Ni-Alloy 718. J. Mater. Eng. Perform. 25 (11), 5076-5086. Sun, J., Guo, Y.B., 2009. A comprehensive experimental study on surface integrity by end
ur
milling Ti–6Al–4V. J. Mater. Process. Technol. 209 (8), 4036-4042. Sun, D., Lemoine, P., Keys, D., Doyle, P., Malinov, S., Zhao, Q., Qin, X., Jin, Y., 2018.
Jo
Hole-making processes and their impacts on the microstructure and fatigue response of aircraft alloys. Int. J. Adv. Manuf. Technol. 94 (5-8), 1719-1726.
Teimouri R., Amini S., Bami A.B., 2018. Evaluation of optimized surface properties and residual stress in ultrasonic assisted ball burnishing of AA6061-T6, Measurement 116, 129–139. Tian, Y., Liu, Y., Wang, F., Jing, X., Zhang, D., Liu, X., 2017. Modeling and analyses of helical milling process. Int. J. Adv. Manuf. Technol. 90 (1-4), 1003-1022. Yao C.F., Wu D.X., Tian L., Ren J.X., Shi K.N., Yang Z.C., 2013. Surface Integrity and 27
Fatigue Behavior for High-Speed Milling Ti–10V–2Fe–3Al Titanium Alloy, J. Fail. Anal. Prev. 14(1), 102-112. Zeilmann, R.P., Weingaertner, W.L., 2006. Analysis of temperature during drilling of Ti6Al4V with minimal quantity of lubricant. J. Mater. Process. Technol. 179 (1-3), 124-127. Zhao, Q., Qin, X., Ji, C., Li, Y., Sun, D., Jin, Y., 2015. Tool life and hole-surface integrity studies for hole-making of Ti6Al4V alloy. Int. J. Adv. Manuf. Technol. 79 (5-8), 1017-1026. Zou, Y., Chen, G., Lu, L., Qin, X., Ren, C.Z., 2019. Kinematic view of cutting mechanism in
ro of
hole-making process of longitude-torsional ultrasonic assisted helical milling, Int. J. Adv.
Jo
ur
na
lP
re
-p
Manuf. Technol. 103 (1-4), 267-280.
28
PII:
S0924-0136(19)30467-4
DOI:
https://doi.org/10.1016/j.jmatprotec.2019.116494
Reference:
PROTEC 116494
To appear in:
Journal of Materials Processing Tech.
Received Date:
20 April 2019
Revised Date:
30 July 2019
Accepted Date:
4 November 2019
Please cite this article as: Chen G, Zou Y, Qin X, Liu J, Feng Q, Ren C, Geometrical texture and surface integrity in helical milling and ultrasonic vibration helical milling of Ti-6Al-4V alloy, Journal of Materials Processing Tech. (2019), doi: https://doi.org/10.1016/j.jmatprotec.2019.116494
This is a PDF file of an article that has undergone enhancements after acceptance, such as the addition of a cover page and metadata, and formatting for readability, but it is not yet the definitive version of record. This version will undergo additional copyediting, typesetting and review before it is published in its final form, but we are providing this version to give early visibility of the article. Please note that, during the production process, errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain. © 2019 Published by Elsevier.
Geometrical texture and surface integrity in helical milling and ultrasonic vibration helical milling of Ti-6Al-4V alloy Guang Chena,b, Yunhe Zoua, Xuda Qina,b, Jian Liua, Qiang Fengc, Chengzu Rena,b* a
Key Laboratory of Mechanism Theory and Equipment Design of Ministry of Education, Tianjin University, Tianjin 300354,
China; b Tianjin
Key Laboratory of Equipment Design and Manufacturing Technology, Tianjin University, Tianjin 300354, China; c CNPC
Bohai Drilling Engineering Company Limited, Tianjin 300280, China
*Corresponding author: [email protected]
ro of
Abstract
Surface integrity of hole-making for Ti-6Al-4V alloy is very important due to its application in aviation industry. Geometrical and mechanical behaviors of machined holes by helical
-p
milling (HM) and ultrasonic vibration helical milling (UVHM) were investigated. The textures of machined holes by the two processes were analyzed and the distances of adjacent
re
textures were calculated according to the movement of peripheral edges. Due to the different surface geometrical textures, the hole-diameter error and surface roughness in UVHM are
lP
smaller than those of HM. A plastic deformation layer with thickness of 4-6µm was observed by the distorted β phase in the cross-section of hole-surface in UVHM by SEM results. Compared with HM process, UVHM generates larger hardness at the subsurface of machined
na
holes (20-200µm) due to the material deformation caused by ultrasonic vibration. Compared with HM, the compressive residual stress of hole surface was increased larger than 63.5% by UVHM. The compressive stresses in UVHM are larger than 150MPa at the depth less than
ur
10µm, which is consistent with the microstructure evolution caused by mechanical
Jo
deformation.
Keywords: Ti-6Al-4V alloy; Helical milling; Ultrasonic vibration; Surface integrity; Surface texture
1. Introduction Titanium alloys are widely applied in aviation industry owing to the high strength-to-weight 1
ratio and good corrosion resistance. However, due to the characteristics of high chemical reactivity, small thermal conductivity and low deformation coefficient, the machining of titanium alloys always causes serious tool wear. Therefore, titanium alloys normally belong to difficult-to-cut materials as described by Sharif and Rahim (2007). In the aviation industry, hole-making process is normally the ultimate and critical stage in manufacturing that contributes to 40%-60% of material removing during aircraft assembly as demonstrated by Schroeder (1998). The assembling holes produce concentrated stress region that initiates and propagates fatigue cracks. Therefore, surface integrity of machined holes is important for the fatigue life of the aircraft structures, and good surface integrity can ensure
ro of
high reliability and good performance of the aircraft as described by Jawahir et al. (2011). The most widely used hole-making method for titanium alloys is conventional drilling, and many investigations about the performance of this process have been reported. Che-Haron and Jawaid (2005) presented that white layer was formed during rough machining of Ti-6Al-4V alloy. Cantero et al. (2005) reported the micro-hardness in the machined surfaces
-p
at the top 75μm thick layer was increased by 30% compared with the original surface hardness during drilling of Ti-6Al-4V alloy under dry condition. Additionally, it was found
re
that the increase of drilling time led to thicker affected zone. Zeilmann and Weingaertner (2006) measured the cutting temperatures of dry drilling and drilling with MQL for
lP
Ti-6Al-4V alloy using type K thermo-elements. The mechanical resistance of Ti-6Al-4V alloy decreased significantly and it was easier to produce plastic deformation at the temperature higher than 500ºC. Rahim and Sasahara (2010) assessed tool wear evolution,
na
surface integrity in drilling of Ti-6Al-4V alloy at dry and MQL lubricant conditions. The machined surface and subsurface generated deformation due to the effect of thermal softening, which made the micro-hardness at 0.025mm depth from the machined surface smaller than
ur
the average hardness of bulk material. In addition, it was found that the effect of strain hardening was increased with the increasing cutting speed and feed rate. Kuo et al. (2014)
Jo
investigated the influence of feed rate and cutting speed on workpiece surface integrity by single-shot drilling of Ti-6Al-4V/CFRP/Al7050 stacks. The average hole-surface roughness (Ra) of Ti-6Al-4V alloy was 0.60μm, and the feed rate had no obvious influence on the micro-hardness of the machined surface. Li et al. (2008) carried out metallographic observation and nano-indentation test of Ti-6Al-4V workpiece and chips with high-throughput drilling. A peak hardness of 9GPa located at subsurface layer was observed at 15-20μm depth from machined surface, while the bulk material hardness was 4-5 GPa. Shyha et al. (2011) studied the hole quality for drilling of Titanium/CFRP/aluminum stacks. 2
The roughness value (Ra) for Ti-6Al-4V alloy layer was less than 1μm. Additionally, adhered material on the Ti-6Al-4V alloy layer indicated the need of a post-finishing process for drilling the stacks. Micro-hardness results exhibited limited strain hardening until 200μm depth from machined surface along with microstructure deformation. As a relatively new technology, helical milling (HM) is a kind of hole-making process by a milling tool moving along helical path. Compared with the traditional drilling, helical milling has advantages of high accuracy, low axial force, and intermittent cutting at the periphery cut edge as demonstrated by Tian et al. (2017). Therefore, this technology drew more attention due to the potential application in the aerospace industry for assembling manufacture in
ro of
recent years. Kihlman et al. (2002) realized automate orbital drilling on assembly structure using a standard industrial robots in the aerospace industry. Orbital drilling generated lower axial cutting forces by the application of industrial robots than conventional drilling within the required hole tolerances. He et al. (2015) carried out hole-making of CFRP/Ti-6Al-4V alloy stacks by HM with various machining parameters. Lower cutting forces, higher
-p
geometrical accuracy, and better surface finishing (Ra≤0.58 μm for Ti-6Al-4V alloy) were achieved, and the need of reaming or de-burring was eliminated. Olvera et al. (2012)
re
machined holes with the average surface roughness of 1μm by traditional drilling method, while, ball helical milling was capable of producing holes with 0.6μm average surface
lP
roughness for Ti-6Al-4V alloy. Additionally, the thermal softening effect was identified to counter the main strain hardening during helical milling when the cutting speed became higher. Sun et al. (2018) investigated the surface roughness, microstructure evolution and
na
micro-hardness of surface and subsurface for aluminum alloy and Ti-6Al-4V alloy using helical milling and drilling processes. The machined parts by HM had longer (46%) fatigue life than that machined by drilling, as less plastic deformation, lower surface roughness, and
ur
compressive residual stress for the hole surface were machined by helical milling. Similarly, Rasti et al. (2019) compared the micro-hardness along the radial and axial directions in the
Jo
hole-making of 4340 steel by drilling, helical milling, profile milling with and without predrilling. Results indicated that the subsurface (within 100µm) hardness of all experiments were higher than bulk material regardless of cutting methods. Compared with the surface machined by HM, the surface machined by conventional drilling generated thicker white layer, and that was mainly caused by the progressive feed of cutting tool. Zhao et al. (2015) investigated surface integrity for titanium alloy with traditional drilling and HM processes. It was reported that the micro-hardness beneath the machined surface gradually decreased to bulk hardness when the depth was about 200μm. Meanwhile, the residual stress in helical 3
milling behaved as compressive stress. Li et al. (2014) evaluated the tool wear in HM machining of Titanium alloy with TiAlN tool. It was reported that the tool wear is an important factor which affects the hole-making quality for Ti-6Al-4V alloy. The wear rates at the frontal cut edge and the periphery cut edge were analyzed. The severe thermal damage is the main factor that causes the tool fracture, especially at tool nose and the front cutting edge. Additionally, a phenomenon of smearing which may cause cracks and other defects on the hole surface with the worn tool was observed. Therefore, the tool-workpiece contact behavior affects the states of cutting tool and surface integrity of hole surface.
ro of
In order to make cutting edges separate from workpiece material intermittently, one feasible method is applying ultrasonic vibration in helical milling process. Nath and Rahman (2008) imposed vibration with a high frequency to cutting tools and developed an Ultrasonic vibration machining (UVM) manufacturing method. Owing to the superiority of intermittent cutting mechanism, the UVM process can reduce the cutting force; improve heat dissipation,
-p
chip breakability and evacuation. Besides, ultrasonic vibration assisted cutting can remove the adhesive chip on the cutting tool, therefore, better hole-surface quality can be achieved.
re
Bai et al. (2019) applied ultrasonically assisted turning (UAT) in machining of SiCp/Al metal-composite. The cutting forces were reduced with the increase of cutting temperature.
lP
Zou et al. (2019) proposed an axial and torsional ultrasonic assisted helical milling (LTUAHM) for hole-making of Ti-6Al-4V alloy. It was reported that the elliptical vibration trajectory affects the cutting angle and the friction behavior between the rake face and chips.
na
Dvivedi and Kumar (2007) investigated the surface quality of pure titanium and titanium alloy by a process of ultrasonic drilling (USD). The surface roughness of USD was much better than the processes of electro discharge machining (EDM) and laser beam machining
ur
(LBM). It was reported that USD process did not produce crack and recast layer which is benefit to machine holes with straight profile. Pecat and Brinksmeier (2014) performed
Jo
low-frequency vibration assisted drilling (LFVAD) on Ti-6Al-4V alloy/CFRP stacks. A remarkable temperature fall of approximately 43% was observed in comparison with traditional drilling, because of the tool cooling related to the interrupted cutting process. Surface integrity and residual stress in ultrasonic-assisted milling of AISI 316L was investigated by Maurotto and Wickramarachchi (2016). It was found that surface integrity and residual pressure varied within a narrow range at different frequencies. Additionally, the minimum value of residual compressive stress was measured at vibration frequency of 40 KHz. Similarly, Suárez et al. (2016) presented the fatigue life and surface integrity of 4
Ni-alloy 718 machined by ultrasonic vibration assisted milling. Surface hardness of workpieces machined by ultrasonic vibration milling was 3.79% higher than those machined with traditional milling methods, and the fatigue life was also increased by 14.74% compared with those machined with traditional milling methods. Ishida et al. (2014) applied ultrasonic vibration in hole-making of CFRP by helical milling process which resulted in reduction of thrust force. The accuracy of machined hole was also improved due to the reduction of delamination size at the exit of the holes. Recently, ultrasonic vibration helical milling (UVHM) method was proposed in our previous work in machining of Ti-6Al-4V alloy as presented by Chen et al. (2019). The
ro of
material removal mechanism was analyzed by the kinematic modeling of cutting trajectories of the bottom and peripheral cutting edges. Meanwhile, the axial force reduction was modeled and the maximum force reduction is 64% compared with those of HM. Additionally, the surface roughness of hole-surface at different conditions and the residual stress at one condition were measured and analyzed by the material removal mechanism. Although the
-p
material removal mechanism of hole machining by UVHM has been studied, the hole-making quality which is crucial to aviation application was not clear yet. Additionally, Chen et al.
re
(2019) mainly focused on the material removal related with tool and tool-workpiece contact behavior. This work focuses on the geometrical texture and surface integrity for machined
lP
hole, concerning the micro-scale material deformation and surface integrity of Ti-6Al-4V alloy by UVHM and HM. The geometrical accuracy, material micro-structure evolution at the
na
surface and subsurface, micro-hardness and residual stress will be investigated. 2. Movement of cutting edge in UVHM The schematic of tool movement in ultrasonic vibration helical milling (UVHM) is shown
ur
in Fig. 1. The milling tool rotates with the machine spindle, and feeds along helical feed path. Meanwhile, the tool vibrates with a high frequency along the axial direction.
Jo
In UVHM, the movement of cutting edges can be characterized in the XYZ coordinate
system. To characterize the mechanism of material removal in the UVHM, the previous work in Chen et al. (2019) investigated the trajectories of the peripheral and bottom cutting edges. The trajectory of specific point in the peripheral cutting edge can be calculated by
5
nrevt Dt nrot t Dh Dt x 2 cos 30 2 cos 30 n t Dh Dt nrevt Dt sin rot y sin 30 2 30 2 z vfa t A sin(2 ft )
(1)
where Dh is the diameter of the target hole; Dt is the diameter of the tool; nrot is the rotation speed of spindle; nrev is orbital revolution speed; vfa is the velocity of axial feed and f is the
re
-p
ro of
frequency of vibration.
lP
Fig.1 Movement of the milling tool in UVHM
To indicate the material removal in the hole-surface during UVHM, the schematic of movement of the peripheral cutting edge and hole-surface is shown in Fig. 2. MN (in Fig. 2a)
na
is one of the peripheral cutting edges of the milling tool. The hole-surface can be analyzed in the coordinate of sOZ. While, Os is the circumferential direction of machined hole; the cutting edge will transform to a short line M1N1, as shown in Fig. 2b. In the machining
ur
process, due to the self-rotation and the vibration along the direction of OZ, the cutting edge M1N1will move with periodic path in the coordinate of sOZ. In addition, the cutting edge can
Jo
separate from the chip periodically at specific process conditions (Chen et al., 2019). In coordinate of XOY, the cutting edge will leave a circular arc M2N2 (Fig. 2c). As the diameter of the tool and hole are different, the arc M2N2 has different curvature compared with the target machined hole, which will affect the geometrical morphology of hole-surface, that is, surface roughness, hole-diameter and roundness.
6
Fig. 2 Relationship of peripheral cutting edge and material removal of hole-surface in UVHM (a) XYZ coordinate (b) sOZ coordinate and (c) XOY coordinate
ro of
3. Experimental procedure Helical milling experiments for Ti-6Al-4V alloy with/without vibration (HM and UVHM) were proposed to study the surface integrity of hole-surface. The experimental setup of UVHM is shown in Fig. 3. The workpiece is machined with dimension of
-p
258mm×125mm×5mm. The experiment was carried out on a machining center DMG®-60 mono BLOCK. Four edges cemented carbides milling tools with composite of WC-12%Co
re
were applied in the HM and UVHM experiments. The ɸ6mm milling tool is used to machine through-holes with nominal diameter of 10mm on workpiece with 5mm thickness. The
lP
detailed geometrical parameters of the milling tool are listed in Table 1. Table 1 Geometrical parameters of the milling tool Rake angle of bottom edge (º)
Rake angle of peripheral edge (º)
Clearance angle of bottom edge (º)
Clearance angle of peripheral edge (º)
Axial length of peripheral edge (mm)
Tool nose radius (mm)
38
0
5
8
10
7
0.5
ur
na
Helix angle (º)
The equipment used to generate the ultrasonic vibration consists of ultrasonic power,
Jo
upper ring with electromagnetic induction coil and a tool holder with bottom ring. A non-contact upper ring electromagnetic unit fixed on the spindle shaft is utilized to transform the electrical signal. The bottom ring with electromagnetic induction coil is installed on the tool holder. When the tool holder is installed on the spindle, there is approximately 1mm distance between the upper ring and bottom ring. Ultrasonic power (type: SZ12) is connected to the upper ring which can transfer the electrical signal to the bottom ring by electromagnetic transmission. Therefore, the tool holder can receive the electrical signal and drive the piezoelectric ceramic generating high frequency axial vibration. The vibration 7
frequency is approximately 34 KHz and the vibration amplitude is approximately 4-6 μm. The cutting forces were measured during the machining process with a three-component dynamometer (Kistler® 9527A) and a charge amplifier (Type: 5070). The signals of cutting forces were measured with a sampling frequency of 20 KHz. The HM and UVHM processes
-p
ro of
were established under dry conditions. The processing parameters are listed in Table 2.
re
Fig. 3 Experimental setup for Ti-6Al-4V alloy machining using HM and UVHM
Test No.
lP
Table 2 Cutting conditions for hole-making of Ti-6Al-4V alloy by HM and UVHM Spindle
speed
Axial feed
Tangential feed
ap(mm/rev)
ft(mm/tooth)
na
nrot (rpm) 2000
0.15
0.03
2
2500
0.15
0.03
3
3000
0.15
0.03
4
3500
0.15
0.03
5
4000
0.15
0.03
6
2500
0.15
0.02
7
2500
0.15
0.04
8
2500
0.15
0.05
9
2500
0.15
0.06
10
2500
0.1
0.03
11
2500
0.2
0.03
Jo
ur
1
8
12
2500
0.25
0.03
13
2500
0.3
0.03
To study the geometry morphology and surface integrity of machined holes using HM and UVHM, surface morphology was measured by a ZYGOTM (type: 7300) scanning white light interferometer. The hole-diameter errors were measured by Coordinate Measuring Machining (CMM) system. The wire electro discharge machining (WEDM) was utilized to section the machined holes in circumferential and axial directions to measure the surface and subsurface microstructure for the machined holes. The samples were mounted and ground by
ro of
abrasive papers from 400 to 2000 grit to remove the surface material. Afterwards, the mounted samples were polished using a Tegramin-25 polishing system for microstructural observation of the cross-section of machined holes. Hardness values at different depths from the machined surface were tested by a MHV digital micro-hardness Tester (type: 2000) using
-p
load of 100mN with 10s duration. The residual stresses were measured by μ-X360n® X ray diffractometer. Cr tube was used to determine the elastic strains for the {1 0 3} diffraction planes of the HCP crystal structure for Ti-6Al-4V alloy. The X-Ray tube voltage and current
re
are 30kV and 1.0mA, respectively. The diffraction peak of X-Ray at the Bragg angle of 140º was measured using a spot with diameter of 1 mm. The Young’s modulus and Poisson’s ratio
lP
are 115.7 GPa and 0.321, respectively. The machined holes were sectioned across and along the axial direction by WEDM before tests, and the axial and circumferential residual stresses at different depths from machined surface were measured. Similar residual tress
na
measurements for machined hole surface were presented in Kwong et al. (2009). Successive layer of material was moved away by electro-polishing to prevent the surface from
ur
regenerating additional residual stresses in order to obtain stress values at different depths from machined surface. Material relaxation during the layer electro-polishing (5~50µm for each layer) was neglected due to the smaller depth removed relative to the thick bulk material
Jo
(3~4mm).
4. Hole-surface topography, surface integrity and tool wear 4.1 Topography of machined Hole-surface The surface topography of machined hole surfaces at four conditions (No.2, No.5, No.11 and No.13.) is shown in Fig. 4. The selected surface area is 0.11 mm× 0.11 mm. The surface 9
ro of -p re lP na ur Jo Fig. 4 Comparison of surface topography of machined hole-surfaces by HM and UVHM at tests of No.2 (a, a′), No.5 (b, b′), No.11 (c, c') and No.13 (d, d'), where a-d represent the results in HM; a'-d' represent the results in UVHM 10
roughness was measured with seven linear positions along circumferential direction (Fig. 4a), and the mean Ra value was presented in each condition. Note that the mean surface roughness of hole surface in UVHM is smaller than that machined in HM when the machining parameters (in Table 2) are identical. For example, the Ramean in No.2' (UVHM) is 0.232 μm, while the value in No.2 (HM) is 0.386μm. Similar phenomena can be observed for Tests No.5, No.11 and No.13. As shown in Fig. 4a-d, periodic convex peaks are observed in the surface generated by HM. Compared with the surface of HM, the surface in UVHM generates periodic convex textures with smaller distance, as shown in Figs. 4a'-d'. In addition, the convex textures appear an
ro of
angle with the axial direction, which is generated by the helix angle of peripheral cutting edges. Compared with HM, the peripheral cutting edge in UVHM generates periodic vibration along the axial direction, as illustrated in Fig. 2. The vibration of peripheral edge will generate a combined effect of friction and compression in UVHM, which can reduce the larger peaks generated in HM. Therefore, the periodic convex textures of hole-surface
-p
indicate micro-scale plastic deformation and it affects the geometrical accuracy of machined hole and may affect the physical behaviors (hardness, residual stresses) and tribological
re
properties of hole-surface.
For the texture machined in HM, the periodic surface curve is mainly caused by the tangential
Jo
ur
na
is presented in Fig. 5.
lP
feed of cutting edges. The schematic of the generation of periodic curves on the hole-surface
Fig. 5 Schematic of curves generated at hole-surface by peripheral cutting edges
During the helical milling feed process, the peripheral cutting edges will leave convex curves on the hole-surface which is generated by adjacent cutting edges, as shown in Figs.4a-d. With the self-rotation and orbital rotation, the points of tangency for the milling tool and the hole surface along the peripheral edge will move along an incline line due to the orbital feed. In 11
another word, the separation points along the peripheral cutting edge will leave a vertical line on the hole-surface if the tool just does self-rotation. Due to the orbital revolution, the separation points for the same peripheral cutting edge will leave an incline line on the machined hole-surface, as shown in Figs.4a-d. The distance of adjacent curves can be analyzed by the machined curves generated by the peripheral edges on the XOY coordinate
ur
na
lP
re
-p
ro of
system.
Jo
Fig. 6 Machined texture generated by peripheral cutting edges (a) and (b) formation of curve A0A1A2A3A4 by same peripheral edge (c) repeated curves generated by adjacent edges (d) the textures generated at the bottom of machined hole
In HM, each cutting edge rotates with an angle speed of 𝜔2 along the clockwise direction, as shown in Fig. 6a. The center of cutting tool rotates along the counter clockwise direction with an angle speed of 𝜔1. As the cutting edge O0A0 rotates to the position of O1A1and O2A2, a trajectory A0A1A2 is formed by the end of bottom cutting edge represented by the solid lines 12
(Fig. 6a). After identical time interval, the cutting edge will rotate from O2A2 to the positions of O3A3and O4A4 continuously and, the cutting edge will leave a trajectory A2A3A4, as shown in Fig. 6b. In this process, the point in the peripheral edge will form the cutting trajectory A0A1A2A3A4 in the XOY coordinate. Therefore, the adjacent cutting edges will leave repeated trajectories with identical shape, as shown in Fig. 6c. Due to the repeated machined trajectories, the machined hole-surface will generate surface texture like BDAEC, which is identical with the experimental textures leaved at the bottom hole surface (Fig. 6d). Therefore, the distance of machined curves (Δl1) in the hole-surface of HM process (arc DE in Fig. 6) can be calculated by the adjacent points of tangency of the peripheral edge (arc AB in Fig. 6)
lP
re
-p
ro of
and the hole-surface, such as the arc of A1B2 in Fig. 7.
na
Fig. 7 Calculation of the distance of the adjacent convex textures
In Fig. 7, O1A1 and OA1 are in the same line, that is, A1 is the point of tangency between the
ur
tool and the hole surface. After time interval of Δt, the adjacent cutting edge O1B1 rotates an angle (γ) to O2B2, where B2 is the next tangency point between the tool and hole-surface, that is, O2B2 will leave another curve on the hole-surface due to the orbital feed, as shown in Fig.
Jo
6c. During time interval of Δt, each cutting edge rotates an angle of γ, at the mean time, the center of milling tool rotates an angle of θ (Fig. 7). Therefore, distance of adjacent curves on the hole-surface can be calculated according to the geometrical relationship described in Chen et al. (2019) as
2π / Z e 2 t t 1 13
(2)
where Ze=4 is the number of cutting edges. Then, the distance of arc A1B2 (Δl1 in Fig. 4d) can be calculated by
l1`
2π1 (e r ) Z e (1 2 )
(3)
where e equals to the length of OO1 (Fig. 7), r is the diameter of the cutting tool. In addition, the distance of adjacent curves (such as Δl2 in Fig. 4d') generated by UVHM (Fig.4a'-4d') can be calculated according to Chen et al. (2019) by the instantaneous speed of peripheral edge along the circumferential direction as t
s vt dt
(5)
-p
v v 2 v 2 2v v cos 1 2 1 2 t v1 1e v2 2 r 2πn 60 rev 1 2 2πnrot 60
ro of
(4)
0
(6)
lP
re
3 π 2 π 3 + 2 1 2 1 1t 2 2t
na
However, as the speed of orbital revolution is much lower than that of self-rotation speed (ω1<<ω2), the linear speed of the peripheral edge along the trajectory of A1A2A3 can be considered as vt≈ω2r=2πnrotr/60. Therefore, the distance of curves in UVHM (Fig. 4d') can
Jo
ur
be calculated as
l2
2 r
(7)
f
where f is the actual frequency in UVHM. Then, the calculated distances of adjacent curves in Fig. 4 are shown in Table 3. It can be seen that the distances of adjacent textures machined by UVHM are smaller than those of HM. The ratios of the distances in HM and UVHM are 3.186 (for 2500rpm) and 2.059 (for 4000rpm), respectively. Therefore, it was mostly affected by the tool rotation speed. In addition, the calculated distances in HM for Test Nos. 2, 5, 11 and 13 are identical (Δl1=74.328 µm), while, the calculated distances in UVHM for Test Nos. 2, 5, 11 and 13 have little difference which is caused by the variation of vibration frequency. 14
The calculated results agree with the measured morphology of machined hole-surface in Fig. 4.
Table 3 Calculated distances of adjacent textures in the machined hole-surface Experimental
Distance of adjacent
Distance
of
adjacent
conditions
curves in HM Δl1(µm)
curves in UVHM
Ratio Δl1/Δl2
Δl2(µm) 74.327901
23.327687
3.186252464
No. 5
74.327901
36.094028
2.05928526
No. 11
74.327901
23.356144
3.186252464
No. 13
74.327901
23.307601
3.186252464
ro of
No. 2
4.2 Hole-diameter error
-p
Hole-diameter error presents the deviation of the actual diameter and nominal diameter which is related with the geometrical morphology of hole-surface. The coordinates of 8 points on
re
the hole-surface at 1 mm depth from the entry and outlet of the machined holes were applied to evaluate the hole-diameter error. The hole-diameter errors near the entry and outlet of hole-surfaces are presented in Fig. 8. The test positions are also illustrated in Fig. 8b. In this
lP
work, the hole-diameter error represents the relative error between the measured diameter and the target diameter (ϕ10mm), and the measured values were the average errors of the repeated experiments. It can be seen that the hole-diameter errors are less than 0.03mm for both HM
na
and UVHM. The values of entry area are larger than those of outlet area for most conditions. In addition, the hole-diameter error decreases with the increasing cutting speed at both of
ur
entry and outlet areas. Similarly, with increase of tangential feed (from 0.02 to 0.06mm/r), the hole-diameter errors near entry are mostly positive and decrease gradually with tangential
Jo
feed. While, the hole-diameter errors near outlet area are mostly negative and also decrease with the increasing tangential feed. Additionally, for most conditions, the hole-diameter errors of UVHM are slightly smaller than those of HM, especially for the hole-diameter errors at outlet area. Note that the hole-diameter errors were tested on the circumferential direction (as indicated in Fig. 8b), and it should be related with the convex textures formed along the circumferential direction, as shown in Fig. 4c. As discussed in section 4.1, the surface convex textures in HM are generated by the movement of peripheral edges. In UVHM, a rubbing effect was caused 15
during the vibration of peripheral edges, which will affect the geometrical morphology and thermal-mechanical behavior of machined surface. The geometrical morphology can be characterized by the distances of convex textures by Eqs. (3) and (7). In HM, the tangential feed per tooth fzt is related with the tool self-rotation nrot and orbital feed nrev.
2πenrev f zt Ze nrot
(8)
Thus, the distance of adjacent textures in HM can be calculated based on Eqs.(3)(5)(8),
2π(e r ) 2 e Z e (1 ) Z e f zt
ro of
(9)
Jo
ur
na
lP
re
-p
l1`
Fig. 8 Comparison of hole-diameter errors in HM and UVHM at different (a) spindle speeds, (b) tangential feed per tooth and (c) axial feeds.
Therefore, the adjacent distance of the textures in HM (Δl1) will increase with the increasing tangential feed per tooth (fzt). Geometrically, the dimension of Δh will increase with the adjacent distance Δl1, which will lead to decrease of the hole-diameter. This agrees with the measured results in Fig. 8b. With the increase of cutting speed from 2000rpm to 4000rpm, the 16
tangential feed per tooth is kept constant, the hole-diameter decreased slightly with the spindle speed (Fig. 8a). That should be related with the thermal deformation, because the cutting temperature normally increases with instantaneous cutting speed. In addition, when the axial feed increases, the rotation speed and tangential feed keep constant, the hole-diameter remains stable due to the same Δl1 and similar heat generation. Therefore, the surface topography for machined hole affects the geometrical accuracy. To further indicate the variation of hole diameter along the axial direction, the hole diameters at different depths (1-4mm) along axial direction were presented in Fig. 9. Both of the hole diameters machined by HM and UVHM decrease with the increasing axial depth, from 1mm to 4mm along axial direction. In detail, when the depth along the axial direction increases
ro of
from 1mm to 4mm, the hole diameter in UVHM decreases from ɸ9.994mm to ɸ9.982mm,
lP
re
-p
while the values in HM decreases from ɸ9.979mm to ɸ9.959mm.
na
Fig. 9 Variation of hole diameters along axial direction for the condition of Test No.9 (nrot=2500rpm, ft=0.06mm/tooth, ap=0.15mm/rev)
ur
The decrease of hole diameter along the axial direction is mainly related with two factors. Firstly, it is related with the tool deflection which is caused by the resultant forces due to the
Jo
tangential feed. The tool deflection may accumulate with the increase of axial feed due to the support from machined hole surface. Secondly, with the increase of axial feed, the cutting temperature near the cutting zone accumulates and the thermal elastic deformation of the workpiece will increase during the cutting process. Therefore, the hole diameter will change accordingly when the workpiece is cooled down after machining. Similar with the results of the hole diameter errors in Fig. 8, the hole diameters machined by UVHM are closer to ɸ10mm than those machined by HM. Although the hole diameter decreases with the 17
increasing depth along axial direction, both of the hole diameters in HM and UVHM are close to target value, and the largest error is approximately 0.04mm. 4.3 Microstructure of sub-surface in HM and UVHM Fig. 10 shows comparison of cross-sectional SEM images of surface and subsurface microstructure produced by HM and UVHM under the machining conditions of No.1 (a, a′ and b, b′), No.2 (c and c′) and No.5 (d and d′). As shown in Fig.10a-10d, there is no obvious distorted β phase on the subsurface in HM, that is, no obvious plastic deformation occurred in HM process along the axial direction. This phenomenon is similar to the microstructure
ro of
results presented by Sun et al. (2018). In UVHM, distorted β phase are observed under the machined hole-surface. As illustrated in Fig. 10a′-d′, the distorted β phase mainly occurs between the hole-surface and a line “L”. The β phase does not exhibit deformation in the area below the line “L”. Therefore, a plastic deformation layer with thickness of approximately
-p
4-6μm exist between the surface and the line “L” in UVHM.
In addition, it is worth noting that the periodic concave textures can be observed on the hole surface in UVHM process, as indicated in Fig. 10a′-d′. The periodic concave texture should
re
be caused by the impulsive sinusoidal cutting trajectory which is related with vibration. Additionally, the periodic surface textures are also observed in the surface topography along
lP
the axial direction, similar with the results in Fig. 4.
Note that the morphology of top surface in SEM graph represents the topology along the axial edges (in Fig. 4c and 4c'). As presented in Fig. 4c, the speed of peripheral edge in HM is
na
along the circumferential direction, which means the instantaneous cutting speed of the points in the peripheral cutting edge is normal to the top surfaces of SEM results in HM. Therefore,
ur
the microstructure of subsurface did not exhibit plastic deformation (distorted β grains). In UVHM, due to the axial vibration, the machined hole-surface generates a thin plastic deformation layer. It should be noted that the instantaneous speed direction of peripheral edge
Jo
has an incline angle with the axial edge (in Fig. 4c'). That is, the distorted grains in Fig. 10a'-d' just represent the deformation state of axial direction, but not the largest deformation direction. The microstructure evolution on the surface and subsurface will contributed to physical-mechanical properties of the hole-surface, i.e., micro-hardness and residual stress states.
18
ro of -p re lP na ur Jo Fig. 10 Comparison of cross-section SEM images of surface microstructure at conditions of No.1 (a, a′ and b, b′), No.2 (c and c') and No.5 (d and d'), where a-d represent the results in HM; a'-d' represent the results in UVHM. 19
4.4 Micro-hardness of machined hole-surface Micro-hardness is normally determined by the coupling effect of mechanical and thermal loads. When the titanium alloy is subjected to high temperature and high pressure during the dry cutting process, a competing effect of work hardening and thermal softening exists and finally leads to specific thermal-mechanical behaviors of the machined surface. This phenomenon can be characterized by the cutting force, temperature and micro-hardness, which are caused by plastic deformation, friction and tool-workpiece contact compression, etc. To investigate the mechanical behaviors of machined hole-surface, the micro-hardness of
ro of
subsurface from 20μm to 250μm depth from the machined surface was tested, as illustrated in Fig. 11. The micro-hardness at each depth was measured three times and the average value
lP
re
-p
was considered as the hardness for different depths from machined surface.
na
Fig. 11 Measured indentation beneath the machined surface at test No.1 (nrot=2000rpm, ft=0.03mm/tooth, ap=0.15mm/rev)
ur
Fig. 12 presents measured micro-hardness of machined subsurface in HM and UVHM at the conditions of Test Nos.1, 2, and 5. The bulk hardness is approximately 333-343HV. The
Jo
micro-hardness in HM and UVHM processes at tests Nos. 1, 2, and 5 exhibit work hardening effects from 0-200µm depth from machined surface, which means the hardness of machined surface or subsurface is higher than that of the bulk material. For example, the hardness of UVHM at 40µm depth from machined surface is 363HV in Test No.1, while, the harnesses at depth larger than 150µm are close to the bulk hardness. However, due to various processing parameters, the largest hardness appeared at different depths for the given conditions. The largest hardness of UVHM with rotation speed of 2500rpm is 362HV which located at 60µm depth. That should be related to the processing parameters (cutting speed, tangential and axial 20
feed rates) and also related with the contact behavior of tool-hole surface. As illustrated in Fig. 4, the material removal and tool-hole surface contact behavior are not identical along the circumferential direction. Therefore, the surface hardness may behave a little fluctuation for different conditions. In addition, compared with the subsurface hardness in HM, the hardness in UVHM at depth less than 200µm is enhanced. For example, for Test No. 1, the hardness of UVHM at 20-80μm depths is approximately 10HV larger than those of HM. Besides, for Test No. 2, the hardness of UVHM at 20-200μm depths is 20-30HV larger than those of HM,
ur
na
lP
re
-p
ro of
indicating a hardness rise by 5.9%-9.1%.
Fig. 12 Comparison of micro-hardness at (a) Test No.1 (nrot=2000rpm, ft=0.03mm/tooth, ap=0.15mm/rev); (b) Test No.2 (nrot=2500rpm, ft=0.03mm/tooth, ap=0.15mm/rev); (c) Test No.5 (nrot=4000rpm,
Jo
ft=0.03mm/tooth, ap=0.15mm/rev).
The increase of subsurface hardness should be caused by the plastic deformation at the machined surface and subsurface which can be observed in the metallurgical behavior of the machined subsurface in Fig. 10. The deformation layer in the top machined surface in UVHM will result in more work hardening effect and higher hardness than HM process. Chen et al. (2019) reported that the vibration can generate rubbing and friction effects to the surface of 21
machined holes. In addition, the vibration leads to distorted β phase in the deformation layer (Figs. 10a'-d') and cause more work hardening effect in UVHM. Besides, it was reported the instantaneous highest cutting speed of peripheral edges in UVHM was much higher than that in HM, which can cause more strain and strain rate hardening in UVHM, increasing the hardness of machined surface and subsurface. Similar work (or strain) hardening phenomena were also reported in machining of titanium and other light-weight alloys due to the effects of rubbing, burnishing and the cutting-extrusion in milling by Yao et al. (2013) and vibration assisted burnishing proposed by Teimouri et al. (2018).
ro of
4.5 Residual stress of hole-surface in HM and UVHM Residual stress of machined surface and subsurface is normally produced by the coupled effects of the thermal-mechanical loading. In general, tensile residual stress normally involves higher cutting temperature and intensive plastic bulge deformation. Sun and Guo (2009) stated that the compressive residual stress was dominated by mechanical deformation.
-p
Compared with conventional drilling, helical milling process can remove chips by larger hole space and dissipate cutting heat. Therefore, the mechanical loading is predominant in HM,
according to Zhao et al. (2015).
re
and the compressive residual stress was observed for machining of Ti-6Al-4V alloy by HM
lP
To further investigate the mechanical behavior of machined hole-surface by HM and UVHM, the residual stresses along the circumferential and axial directions of two conditions (Test No. 2 and No.5) are illustrated in Fig. 13. Note that the residual stresses of both HM and UVHM
na
processes generate compressive stresses which are beneficial to enhance the fatigue strength
Jo
ur
of hole-surface.
Fig. 13 Comparison of residual stress at (a) Test No.2 (nrot=2500rpm, ft=0.03mm/tooth, ap=0.15mm/rev); and (b) Test No.5 (nrot =4000rpm, ft=0.03mm/tooth, ap=0.15mm/rev). 22
For test No. 2, the residual stress of hole surface machined by HM along axial direction is -127MPa, whereas the stresses in UVHM along axial direction is -216MPa. While, the residual stress of hole surface machined by HM along circumferential direction is -104MPa, whereas the stresses in UVHM along circumferential direction is -170MPa. That is, compared with HM, UVHM increases the compressive surface residual stress by 70.1% and 63.5% for axial and circumferential directions, respectively. Similarly, for test No.5, the compressive residual stress of hole-surface along the axial direction is increased by 70.2%, while the stress along the circumferential direction is increased by 172.1%. Note that most of the compressive
ro of
stresses which are larger than 150MPa in UVHM located at the top layer with depth smaller than 10µm. That is related to the grain deformation caused by mechanical deformation in UVHM, as the thickness of top deformation layer (Fig. 10) in UVHM is approximately 4-6μm. Besides, comparing the residual stress with the rotation speed of 2500 and 4000rpm, the rotation speed has little influence on residual stresses in HM and UVHM. Due to the
-p
plastic deformation caused by ultrasonic vibration, the UVHM process can generate higher
4.6 Tool wear in HM and UVHM
re
compressive residual stress which can increase the fatigue strength of machined hole-surface.
lP
Tool wear is an important factor which affects the performance of machining quality and machining cost. In this work, to quantify the wear of the HM and UVHM processes, the average flank wear (VB) was measured to characterize the tool wear in HM and UVHM at
na
the condition of Test No. 2, as shown in Fig. 14. During the experiments, new tools were applied in HM and UVHM. As the axial feed is 0.15mm/rev, the length of peripheral edge
ur
involved in cutting is much smaller than that of the bottom cutting edge. Then, the average flank wear of bottom cutting edge was measured by microscope when each five holes were machined. As shown in Fig. 14, the tool wear in HM and UVHM increase with similar trends.
Jo
That is, tool wear increases until 10 holes, and then remains stable until 25 holes, and the tool wear increases rapidly with the hole numbers until 40. Additionally, the flank wears in UVHM are more rapidly than those in HM. When 40 holes was machined, the average flank wears are approximately 65µm and 100µm in HM and UVHM, respectively. The speed of tool wear should be related with the instantaneous cutting speed. According to the study of Li et al. (2014), tool wear increases when higher cutting speed was employed. In UVHM, due to the high frequency axial vibration, the instantaneous speed of bottom edge 23
generates periodical sinusoidal fluctuation. The highest instantaneous of bottom cutting edge is much higher than that in HM. Besides, the material removal in HM is continuous, while the cutting of bottom edge in UVHM was intermitted due to the vibration, as stated by Chen et al. (2019). The intermitted cutting in UVHM will change the thickness of uncut chip thickness and will generate impact effect between the flank surface of bottom edge and the workpiece. That is, the highest instantaneous cutting speed, changeable uncut chip thickness and impact effect can lead to higher flank tool wear in UVHM. Additionally, the average flank wear in HM and UVHM are far less than 200µm when 40 holes were machined. That is, during the surface integrity investigation in HM and UVHM (i.e., 13 holes in HM and 13 holes in
ro of
UVHM) experiments, the tool wear should be in the stable stage, that is, less than 70µm as
lP
re
-p
indicated in Fig. 14.
ur
5. Conclusions
na
Fig. 14 Flank wear versus machined hole number with HM and UVHM at Test No.2 (nrot=2500rpm, ft=0.03mm/tooth, ap=0.15mm/rev)
In this paper, comparative study for geometrical topography and surface integrity of Ti-6Al-4V alloy by helical milling (HM) and ultrasonic vibration helical milling (UVHM)
Jo
was proposed. The surface textures of machined surfaces in HM and UVHM were analyzed. The geometrical and physical behaviors of machined holes, including the hole-diameter error, micro-hardness, microstructure evolution and residual stresses in two different processes were presented. The main conclusions are summarized as follow: 1. The distance of adjacent textures of machined surface in HM and UVHM were calculated. The adjacent textures in HM are larger than those generated by UVHM. The textures in HM were related to the tangential feed of peripheral edges, while, the 24
textures in UVHM were caused by the helical feed and the ultrasonic vibration along the axial direction. 2. The hole-diameter errors in UVHM are smaller than those in HM. The hole-diameter error is related with the textures along the circumferential direction. The hole-diameter value decreases with the increasing tangential feed per tooth. 3. Due to the movement of peripheral cutting edges, approximately 4-6µm deformation layers with distorted β phase were observed by the cross-section of SEM images in UVHM, however, no deformation layer along axial direction was generated in HM. 4. Compared with HM, UVHM generates more work hardening effect at the subsurface of
ro of
machined holes (20-200µm). The vibration load generates more plastic deformation at the surface and subsurface of machined hole, increasing the work hardening effect. The micro-hardness increases with the spindle speed, tangential and axial feeds.
5. The compressive residual stress of hole-surface was increased larger than 63.5% by UVHM for given conditions compared with HM process. The compressive stresses that
-p
are larger than 150MPa in UVHM located at the top layer with depth smaller than 10µm, which is consistent with the microstructure evolution caused by mechanical
re
deformation in UVHM, as illustrated by SEM images.
6. The tool wear in UVHM is larger than that in HM, due to the highest instantaneous
lP
cutting speed, variable uncut chip thickness and the impact effect between the cutting edge and machined surface.
na
Acknowledgements
The authors would like to acknowledge the financial support received from the National
ur
Natural Science Foundation of China (No. 51575384).
Jo
References
Bai, W., Roy, A., Sun, R., Silberschmidt, V.V., 2019. Enhanced machinability of SiC-reinforced metal-matrix composite with hybrid turning. J. Mater. Process. Technol. 268, 149-161. Cantero, J.L., Tardío, M.M., Canteli, J.A., Marcos, M., Miguélez, M.H., 2005. Dry drilling of alloy Ti–6Al–4V. Int. J. Mach. Tools. Manuf. 45 (11), 1246-1255. Che-Haron, C., Jawaid, A., 2005. The effect of machining on surface integrity of titanium 25
alloy Ti–6% Al–4% V. J. Mater. Process. Technol. 166 (2), 188-192. Chen, G., Ren, C., Zou, Y., Qin, X., Lu, L., Li, S., 2019. Mechanism for material removal in ultrasonic vibration helical milling of Ti-6Al-4V alloy. Int. J. Mach. Tools. Manuf. 138, 1-13. Dvivedi, A., Kumar, P., 2007. Surface quality evaluation in ultrasonic drilling through the Taguchi technique. Int. J. Adv. Manuf. Technol. 34 (1-2), 131-140. He, G., Li, H., Jiang, Y., Qin, X., Zhang, X., Guan, Y., 2015. Helical milling of CFRP/Ti-6Al-4V stacks with varying machining parameters. Trans. Tianjin Univ. 21 (1), 56-63.
ro of
Ishida, T., Noma, K., Kakinuma, Y., Aoyama, T., Hamada, S., Ogawa, H., Higaino, T., 2014. Helical milling of carbon fiber reinforced plastics using ultrasonic vibration and liquid nitrogen. Procedia CIRP. 24, 13-18.
Jawahir, I.S., Brinksmeier, E., M'saoubi, R., Aspinwall, D., Outeiro, J., Meyer, D., Umbrello,
CIRP. Ann-Manuf. Technol. 60 (2), 603-626.
-p
D., Jayal, A., 2011. Surface integrity in material removal processes: Recent advances.
Kihlman, H., Eriksson, I., Ennis, M., 2002. Robotic orbital drilling of structures for aerospace
re
applications. SAE. Tech. Pap. No. 2002-01-2636.
Kuo, C.L., Soo, S., Aspinwall, D., Thomas, W., Bradley, S., Pearson, D., M'saoubi, R., Leahy,
lP
W., 2014. The effect of cutting speed and feed rate on hole-surface integrity in single-shot drilling of metallic-composite stacks. Procedia CIRP. 13, 405-410. Kwong, J., Axinte, D.A., Withers, P.J. 2009. The sensitivity of Ni-based superalloy to hole
na
making operations: Influence of process parameters on subsurface damage and residual stress. J. Mater. Process. Technol. 209(8), 3968-3977. Li, H., He, G., Qin, X., Wang, G., Lu, C., Gui, L., 2014. Tool wear and hole quality
ur
investigation in dry helical milling of Ti-6Al-4V alloy. Int. J. Adv. Manuf. Technol. 71 (5-8), 1511-1523.
Jo
Li, R., Riester, L., Watkins, T.R., Blau, P.J., Shih, A.J., 2008. Metallurgical analysis and nanoindentation characterization of Ti–6Al–4V workpiece and chips in high-throughput drilling. Mater. Sci. Eng. A. 472 (1-2), 115-124.
Maurotto, A., Wickramarachchi, C., 2016. Experimental investigations on effects of frequency in ultrasonically-assisted end-milling of AISI 316L: A feasibility study. Ultrasonics. 65, 113-120. Nath, C., Rahman, M., 2008. Effect of machining parameters in ultrasonic vibration cutting. Int. J. Mach. Tools. Manuf. 48 (9), 965-974. 26
Olvera, D., de Lacalle, L.N.L., Urbikain, G., Lamikiz, A., Rodal, P., Zamakona, I., 2012. Hole making using ball helical milling on titanium alloys. Mach. Sci. Technol. 16 (2), 173-188. Pecat, O., Brinksmeier, E., 2014. Tool wear analyses in low frequency vibration assisted drilling of CFRP/Ti6Al4V stack material. Procedia CIRP. 14, 142-147. Rahim, E.A., Sasahara, H., 2010. Effect of MQL liquids on surface integrity when high speed drilling titanium alloy. Key. Eng. Mater. 443, 359-364. Rahim, E.A., Sasahara, H., 2011. Investigation of tool wear and surface integrity on MQL machining of Ti-6AL-4V using biodegradable oil. Proc. Inst. Mech. Eng. Part B-J. Eng.
ro of
Manuf, 225(9), 1505-1511. Rasti, A., Sadeghi, M.H., Farshi, S.S., 2019. An investigation into the effect of surface integrity on the fatigue failure of AISI 4340 steel in different drilling strategies. Eng. Fail. Anal. 95, 66-81.
Schroeder, P., 1998. Widening Interest In Twist Drills: How wide-web drills with
-p
high-performance points can increase efficiency. Mod. Mach. Shop. 71, 106-113.
Sharif, S., Rahim, E., 2007. Performance of coated-and uncoated-carbide tools when drilling
re
titanium alloy–Ti–6Al4V. J. Mater. Process. Technol. 185 (1-3), 72-76. Shyha, I., Soo, S.L., Aspinwall, D., Bradley, S., Perry, R., Harden, P., Dawson, S., 2011. Hole
Manuf. 51 (7-8), 569-578.
lP
quality assessment following drilling of metallic-composite stacks. Int. J. Mach. Tools.
Suárez, A., Veiga, F., de Lacalle, L.N.L., Polvorosa, R., Lutze, S., Wretland, A., 2016. Effects
na
of ultrasonics-assisted face milling on surface integrity and fatigue life of Ni-Alloy 718. J. Mater. Eng. Perform. 25 (11), 5076-5086. Sun, J., Guo, Y.B., 2009. A comprehensive experimental study on surface integrity by end
ur
milling Ti–6Al–4V. J. Mater. Process. Technol. 209 (8), 4036-4042. Sun, D., Lemoine, P., Keys, D., Doyle, P., Malinov, S., Zhao, Q., Qin, X., Jin, Y., 2018.
Jo
Hole-making processes and their impacts on the microstructure and fatigue response of aircraft alloys. Int. J. Adv. Manuf. Technol. 94 (5-8), 1719-1726.
Teimouri R., Amini S., Bami A.B., 2018. Evaluation of optimized surface properties and residual stress in ultrasonic assisted ball burnishing of AA6061-T6, Measurement 116, 129–139. Tian, Y., Liu, Y., Wang, F., Jing, X., Zhang, D., Liu, X., 2017. Modeling and analyses of helical milling process. Int. J. Adv. Manuf. Technol. 90 (1-4), 1003-1022. Yao C.F., Wu D.X., Tian L., Ren J.X., Shi K.N., Yang Z.C., 2013. Surface Integrity and 27
Fatigue Behavior for High-Speed Milling Ti–10V–2Fe–3Al Titanium Alloy, J. Fail. Anal. Prev. 14(1), 102-112. Zeilmann, R.P., Weingaertner, W.L., 2006. Analysis of temperature during drilling of Ti6Al4V with minimal quantity of lubricant. J. Mater. Process. Technol. 179 (1-3), 124-127. Zhao, Q., Qin, X., Ji, C., Li, Y., Sun, D., Jin, Y., 2015. Tool life and hole-surface integrity studies for hole-making of Ti6Al4V alloy. Int. J. Adv. Manuf. Technol. 79 (5-8), 1017-1026. Zou, Y., Chen, G., Lu, L., Qin, X., Ren, C.Z., 2019. Kinematic view of cutting mechanism in
ro of
hole-making process of longitude-torsional ultrasonic assisted helical milling, Int. J. Adv.
Jo
ur
na
lP
re
-p
Manuf. Technol. 103 (1-4), 267-280.
28