Mohamed F El Machado

R M Mansori

Souza

V Seriacopi

S Mezghani

S Crequy

I F Machado

M El Mansori

R M Souza

Study of angular cutting conditions using multiple scratch tests onto low carbon steel: An experimental-numerical approach

Keywords:

Study of angular cutting conditions using multiple scratch tests onto low carbon steel: An experimental-numerical approach

Introduction

Abrasive wear remains as a key phenomenon in many practical applications, not only due to the significant amount of tribological damage that result from the action of hard particles, but also due to the constant need of improvement of manufacturing processes with cutting operations.

Abrasion studies are frequently conducted considering scratch tests, in which a single hard particle contacts a softer slab surface, promoting plastic deformation and/or material removal. The transition between the abrasive mechanisms during scratch tests was characterized and evaluated experimental and/or numerically, based on several approaches, considering: (i) mechanical properties (e.g. hardness, elastic modulus and hardness ratio, strain-hardening, interfacial shear strength) [START_REF] Hokkirigawa | An experimental and theoretical investigation of ploughing, cutting and wedge formation during abrasive wear[END_REF][START_REF] Hokkirigawa | The effect of hardness on the transition of the abrasive wear mechanism of steels[END_REF][START_REF] Machado | The effect of in-service work hardening and crystallographic orientation on the micro-scratch wear analysis of Hadfield steel[END_REF][START_REF] Pöhl | Scratch behavior of soft metallic materials[END_REF][START_REF] Pöhl | Indentation and scratch testing experiment and simulation[END_REF]; (ii) microstructural features (e.g. influence of the second phase particles, crystallographic orientation, deformation-induced phase transformation) [START_REF] Machado | The effect of in-service work hardening and crystallographic orientation on the micro-scratch wear analysis of Hadfield steel[END_REF][START_REF] Seriacopi | Evaluation of Abrasive Mechanisms in Metallic Alloys during Scratch Tests: A Numerical-Experimental Study in Micro-scale[END_REF][START_REF] Franco | Abrasion of Gray Cast Iron: Application to Automotive Engines[END_REF][START_REF] Gahr | Wear by hard particles[END_REF][START_REF] Gahr | Microstructure and wear of materials[END_REF][START_REF] Hu | Synergistic effects of microstructure and abrasion condition on abrasive wear of compositesa modeling study[END_REF][START_REF] Pöhl | Effect of matrix and hard phase properties on the scratch and compound behavior of wear resistant metallic materials containing coarse hard phases[END_REF][START_REF] Tkaya | On some tribological effects of graphite nodules in wear mechanism of SG cast iron: finite element and experimental analysis[END_REF][START_REF] Mendas | Investigating the effects of microstructure on the wear mechanisms in lamellar cast irons via microscratch tests[END_REF]; (iii) material removal factor (f abwhich relates the groove and pile-up areas), degree of wear and degree of penetration [START_REF] Gahr | Wear by hard particles[END_REF][START_REF] Franco | Material removal factor (f ab): a critical assessment of its role in theoretical and practical approaches to abrasive wear of ductile materials[END_REF][START_REF] Kayaba | Analysis of the abrasive wear mechanism by successive observations of wear processes in a scanning electron microscope[END_REF][START_REF] Goel | A numerical study of abrasive wear in tillage tools due to soil-tool interaction[END_REF]; (iv) surface finishing (e.g. ground, mechanical and electrolytic polished surfaces) [START_REF] Seriacopi | Analysis of abrasion mechanisms in the AISI 303 stainless steel: effect of deformed layer[END_REF][START_REF] Avery | Work hardening in relation to abrasion resistance, Climax Molybd[END_REF]; (v) relation between the normal load and geometric parameters of the scratch (e.g. scratch depth and width) [START_REF] Kobrick | Developing abrasion test standards for evaluating lunar construction materials[END_REF][START_REF] Kobrick | Standardization of a volumetric displacement measurement for two-body abrasion scratch test data analysis[END_REF]; (vi) effect of the attack angle of the abrasive particle on the apparent coefficient of friction [START_REF] Mezlini | Experimental, numerical and analytical studies of abrasive wear: correlation between wear mechanisms and friction coefficient[END_REF][START_REF] Tkaya | Influence of the attack angle on the scratch testing of an aluminium alloy by cones: Experimental and numerical studies[END_REF][START_REF] Tkaya | The effect of damage in the numerical simulation of a scratch test[END_REF][START_REF] Wredenberg | On the numerics and correlation of scratch testing[END_REF]; and (vii) specific cutting energy modifications [START_REF] Wang | Specific energy and scratch hardness of gamma titanium aluminides subjected to single-grit pendulum scratching[END_REF][START_REF] Vélez | Evaluation of abrasive wear of ductile cast iron in a single pass pendulum device[END_REF][START_REF] Lee | Experimental Investigation of the Bore Honing Process[END_REF].

Additionally, a study related to the different orientations of the relative movement of an abrasive tool with respect to a ductile surface can be used to understand manufacturing by machining processes. This approach is based on the previous analyses conducted with multiple parallel scratches. Within this context, Williams and Xie investigated the interaction between parallel scratch tracks using an analytical model and experimental tests. These authors observed that the attack angle of the abrasive asperities played an important role to determine https://doi.org/10.1016/j.wear.2019.01.101

⁎ Corresponding author at: Surface Phenomena Laboratory, Escola Politécnica da Universidade de São Paulo, São Paulo, Brazil. E-mail address: vanessaseriacopi@usp.br (V. Seriacopi).

T

the deformation modes and micro-mechanisms (from ploughing to micromachining) on a flat surface [START_REF] Williams | The generation of wear surfaces by the interaction of parallel grooves[END_REF]. In another work, the same authors identified that the abrasive geometry and the relative lateral displacement between repeated passes onto a copper specimen were prevailing to define the material removal rates. In that case, a map was created to design the dominant wear mechanism and wear rates. After multiple passes, material accumulation was observed, which provided the formation of wear debris [START_REF] Xie | The generation of worn surfaces by the repeated interaction of parallel grooves[END_REF]. Mezlini et al. carried out multiple scratches on aluminum alloys to reproduce repetitive loading conditions of granular materials. A microstructural characterization pointed out that the material removal mechanism by abrasion depends on the wedge interaction of the successive scratches. The subsurface strainhardening was also included in the abrasion model for these materials, resulting in a better reproduction of the experimental results [START_REF] Mezlini | Abrasion of aluminium alloy: effect of subsurface hardness and scratch interaction simulation[END_REF]. Focusing on ceramics and hard metals, Gee tested a scratch array varying the repetition number along the same scratch and using a normal load corresponding to 4 N (low load) with a diamond indenter. In this case, the parameter applied to evaluate the material removal was the scratch width. Furthermore, tangential force measurements, and the microstructural analyses, consisted of important resources to study the evolution of the multiple scratching features, allowing to detect the fracture with debris formation as a major cause of the damage on these materials [START_REF] Gee | Low load multiple scratch tests of ceramics and hard metals[END_REF]. Finally, Silva and Mello [START_REF] Da Silva | Using parallel scratches to simulate abrasive wear[END_REF] used a high-resolution device to examine abrasion effects using superimposed scratches on a surface. The topographic and microstructural approaches allowed to identify the dominant abrasive mechanism. An experimental process simulation was specified according to positions of the parallel scratches, randomly distributed, creating patterns of surface morphology and topography, and considering constant and variable normal load conditions.

Based on this background of works focused on single scratch tests or multiple parallel scratches, the aim of this work consisted in analyzing the cutting forces and burr formation as a result of two consecutive scratches. However, different from previous works in the literature, the

Table 2

Johnson-Cook ductile damage parameters of the AISI 1020 steel [START_REF] Kupchella | Incorporation of material variability in the Johnson Cook model[END_REF].

D 1 D 2 D 3 D 4 D 5 0.24 0.54 -1.50 0 0
Table 3 Mechanical properties of the AISI 1020 steel [START_REF]Properties and Selection: Irons, Steels, and High-Performance Alloys[END_REF].

Elastic modulus [GPa]

Poisson Ratio Density [kg.m -3 ] 200 0.30 7800 consecutive scratches were not parallel, but considered angles of 10, 20 or 30 degrees between them. Experimental and numerical procedures were developed to provide this evaluation, in which the axial, radial and tangential forces, as well as the material removal, were the outputs. Numerical analyses allowed to investigate and compare the material removal in the different conditions and were applied as a support to experimental evaluation.

Materials and methods

Angular scratches were conducted on a slab of low carbon steel, with hardness of 126 ± 6 HV 30 . This carbon steel was selected for the analyses since single scratch tests in microscale were previously reported for other soft materials, including annealed AISI 1045 steel, and well-discussed in the literature. Pöhl, Hardes and Theisen [START_REF] Pöhl | Scratch behavior of soft metallic materials[END_REF] pointed out a more pronounced trend to micro-cutting on AISI 1045, in comparison with pure iron, due to its greater work hardening, which avoided a considerable pile-up formation.

In the present work, the abrasive tests were carried out using a modified shaper machine tool [START_REF] Xu | Wear characteristics of polycrystalline diamond tools in orthogonal cutting of CFRP/Ti stacks[END_REF]. An initial set of parallel scratches was conducted on the slab, followed by a second set of scratches with different orientations (10, 20 or 30°) with respect to the previous one. The parameters of the test were: scratch velocity corresponding to 50 m min -1 and a constant depth of cut of 100 μm. The tool used to conduct scratch tests was made of cemented carbide (WC-Co), presenting a cono-spherical geometry with 233 ± 53 μm tip radius and 30°apex angle. Based on these settings, and considering the geometry of the tool, the slab material can be simplified as homogeneous. The experimental apparatus can be seen in Fig. 1. Five replicas for each condition (10, 20 and 30°) were tested to analyze the reproducibility and for comparison with the numerical approach.

The numerical model was prepared using the Finite Element Method (FEM) with Abaqus/Explicit ® 2016 code. As displayed in Fig. 2, the tribosystem was modelled assuming that the tool was rigid, with a 235 μm tip radius, and the slab presented an elastic-plastic behavior. Johnson-Cook constitutive -Eq. ( 1) [START_REF] Johnson | Cook A constitutive model and data for metals subjected to large strains, high strain rates and high temperatures[END_REF] -and ductile damage -Eq. ( 2) [START_REF] Johnson | Fracture characteristics of three metals subjected to various strains, strain rates, temperatures and pressures[END_REF] -models were assigned to this slab to allow plastic deformation and material removal.

⎜ ⎟ ⎜ ⎟ = + ⎛ ⎝ + ⎞ ⎠ ⎡ ⎣ ⎢ -⎛ ⎝ - - ⎞ ⎠ ⎤ ⎦ ⎥ σ A Bε Cln ε ε T T T T ̅ ( ) 1 ̇̇1 n r m r m 0 (1)
Where: σ is the equivalent stress, ε is the equivalent plastic strain, ε̇is the strain rate, ε0 is the reference strain rate, A is the initial yield stress, B is the hardening modulus, C is the strain rate coefficient, n is the strain-hardening coefficient, m is the thermal softening coefficient, T is the process temperature, T r is the room temperature and T m is the melting temperature of the material [START_REF] Johnson | Cook A constitutive model and data for metals subjected to large strains, high strain rates and high temperatures[END_REF]. In which: ε f is the equivalent strain to failure, -D D 1 5 are experimental failure constants of the Johnson-Cook damage equation, p is the pressure, q is the von Mises stress. The other parameters are described according to the terms in Eq. ( 1) [START_REF] Johnson | Fracture characteristics of three metals subjected to various strains, strain rates, temperatures and pressures[END_REF].

⎜ ⎟ ⎜ ⎟ = ⎡ ⎣ ⎢ + ⎛ ⎝ ⎞ ⎠ ⎤ ⎦ ⎥ ⎡ ⎣ ⎢ + ⎤ ⎦ ⎥ ⎡ ⎣ ⎢ + ⎛ ⎝ - - ⎞ ⎠ ⎤ ⎦ ⎥ ε D De x p D p q D ln ε ε D T T T T 1 ̇̇1 f r m r 1 2 3 4 0 5 (2)
The Johnson-Cook parameters of AISI 1020 steel used as modelling inputs are displayed in Table 1 [START_REF] Kui | The dynamic characteristics and constitutive model of 1020 steel[END_REF] and Table 2 [START_REF] Kupchella | Incorporation of material variability in the Johnson Cook model[END_REF]. Furthermore, the other mechanical properties assigned to the carbon steel specimen are shown in Table 3 [START_REF]Properties and Selection: Irons, Steels, and High-Performance Alloys[END_REF].

The finite element mesh assigned to the slab presented the following features: domain of 2500 × 1800 × 300 μm³ , discretized by 1,580,625 linear hexahedral elements of the C3D8 type. 32 CPUs available in high performance massively parallel machines of a cluster have been used for this modelling, resulting in approximately 18 h of computational time for each case of angular scratch. Regarding the boundary conditions, the numerical model was planned based on the following steps of the tool movement with respect to the slab:

(i) First scratch: the tool treads 3000 μm along the length (negative xdisplacement -Fig. 2) with a constant depth of penetration of 100 μm (negative y-displacement -Fig. 2); (ii) (ii-) Tool Moving: the tool goes through 650 μm along the width (negative z-displacement -Fig. 2); (iii) (iii-) Angular scratches: the tool is moved back towards its origin, resulting in a second scratch, which considered the 10, 20 or 30°a ngle (x-z plane -Fig. 2). The results were evaluated in terms of cutting forces and burr formation. In this context, burr was defined as the chip formed at the intersection of two angular scratches.

During the experimental tests, a Kistler platform was mounted on the shaper machine tool and used to acquire the cutting forces (radial, axial and tangential forces). Moreover, the signal of the experimental cutting forces was treated using a signal denoiser. This noise filter was employed, consisting of an important resource to separate the wavelet and vibration effects. Additionally, after the scratch tests, the topography of the specimens was characterized using an optical profilometer to provide the information about the burr. Optical and scanning electron microscopy analyses were conducted on the cross sections at the intersection of consecutive scratches to investigate the influence of the abrasive mechanisms on the material subsurface.

Results and discussion

Fig. 3 presents an example of the experimentally calculated cutting forces as a function of time. When the tool crossed a previous scratch (groove) with a certain orientation, the cutting forces were reduced locally (points (a) and (b) in Fig. 3), as expected. In addition, Fig. 4 shows the average values of the experimental cutting forces, after the signal treatment. The lowest angle (10°) resulted in lower average values of the cutting forces.

Table 4 shows a comparison between experimental and numerical results of the cutting forces, indicating good agreement between numerical and experimental data. The differences are probably related to slight divergences in the mechanical properties and the approximated tool tip radius assumed in the modelling. Furthermore, despite several replacements of the tool, it could present wear caused by the successive scratches during the experimental tests, which was not considered in the numerical model.

Concerning the angular scratch features, the average width and pileup height were characterized using optical profilometry. A theoretical analysis of the scratch width was conducted based on the literature [START_REF] Kobrick | Developing abrasion test standards for evaluating lunar construction materials[END_REF]. These geometrical results are found in Table 5 and denote a good correlation of the numerical and experimental findings. Nevertheless, the difference in view of the pile-up height can be assigned to a slightly larger strain-hardening coefficient assumed in the FEM modelling, reducing the global plastic deformation when compared to the experimental one.

To evaluate the burr formation and its consequences in the manufacturing processes, two distinct analyses were conducted: (a) experimental characterization of the topography, to verify the height effects (Fig. 5); and (b-) numerical result regarding the equivalent plastic strain at the nucleation of the damage, to observe the cumulative material (Fig. 6).

Figs. 5 and6 show that the experimental burr formation was qualitatively reproduced by the numerical approach. The highest values of length and width occurred for the 20°orientation, which was also critical in terms of the cutting forces (Table 4). However, the burr height was more pronounced for the 10°orientation. Observing the experimental and numerical results, the following tendency can be reported: (i-) a large burr formation was found when the 10°angular scratch was conducted; (ii-) for the 20°condition, burr formation was accentuated, as well as pile-up occurrence; and (iii-) the burr was removed when the 30°scratch was conducted.

Considering that the modelling was validated, a numerical evaluation based on the material removal and dissipated energy during the cut can be used to understand the critical case corresponding to 20°scratch (cutting forces), and the burr formation (Figs. 5 and6). Fig. 7 shows that the removed volume per distance decreased in the 10°-20°range, in which the plastic dissipation energy increases. This fact agrees with the higher cutting forces verified for the 20°condition (Table 4), i.e. the force should increase to provide the cutting with the occurrence of large plastic deformation of the carbon steel. Consequently, the lowest removed volume was observed in the 20°case (Fig. 7), in which a very high plastic dissipation energy was necessary to proceed the scratch, leading to a slight cutting energy (damage) decrease (Table 6).

Although lower than the 10°case, significant material removal is achieved for the 30°condition (Fig. 7) because of the aggressive scratch setup. The geometrical settings lead to a large amount of energy dissipated due to the damage (cutting energy in Table 4), in conjunction with a reduction of the plastic dissipation energy.

Fig. 8(a-d) present the characterization of the cross sections at the intersection of two angled scratches obtained experimentally (section indicated in Fig. 9a). The observation of the etched material subsurface using optical microscopy (Fig. 8a) allowed verifying the presence of a deformed layer, whose thickness is dependent on the angle between scratches. This characterization shows that the thinnest deformed layer was observed at 20°, which indicated a higher material hardening and a reduced removed volume in this case. SEM images in Fig. 8(b-c) provide more details of the deformed layers obtained in each case. The mean values and standard deviations of five representative measurements resulting from this experimental procedure are also provided (Table 7). The thickness of the deformed layer is in agreement with the behavior of the cutting forces (Table 4):

(i) 10°case (Fig. 8b)maximum deformed layer thickness could be associated with the minimum forces necessary to cut the carbon steel. For this geometrical setup, the material strain-hardening was not favored during the combination of scratches, resulting in an intensified removed volume with massive plastic deformation (Fig. 7); (ii) 20°case (Fig. 8c) -minimum deformed layer thickness corresponded to the maximum cutting forces, which corroborated the hypothesis of material strain-hardening with the lowest material removal (Fig. 7); (iii) 30°case (Fig. 8d)consisted of an intermediary condition between 10°and 20°responses. The deformed layer thickness was within the range determined by the lower limit (20°) and the higher limit (10°). Conversely, the opposite could be thought based on the cutting forces. Therefore, as expected, an intermediate value of the removed volume was predicted by FEM for the 30°angle case (Fig. 7). Fig. 9 presents a numerical analysis of the plastic strain fields at the material subsurface, indicating the section at which the experimental and numerical analyses were effected (Fig. 9a). For all angles, the equivalent plastic strain decreased from the surface to the subsurface reaching a minimum value of 0.009. Moreover, the numerical thickness of the deformed layer followed the experimental behavior (Table 7), i.e. the scratches at 30°resulted in thicknesses lower than those at 10°and higher than those at 20°. A result similar to the 20°critical case was found by Mezlini et al. [START_REF] Mezlini | Abrasion of aluminium alloy: effect of subsurface hardness and scratch interaction simulation[END_REF] for parallel scratch tests conducted onto an annealed aluminum alloy. Those authors observed that a significant subsurface strain-hardened layer was formed in addition to an accumulation of deformed material on the scratch sides as a function of the alloy ductile properties. Therefore, the experimental-numerical results of the 20°case could be explained using two approaches:

(a) mechanical properties -subsurface layer analyses (Table 7) denote the material strain-hardening influence; (b) damage properties -the higher values related to burr length and width (Figs. 5 and6) imply in a greater material plastic deformation on the scratch crossing, due to material ductile behavior.

Conclusions

The experimental and numerical studies developed in this work, to analyze angular scratches (10, 20 and 30°) on low carbon steel, provide the following conclusions:

• The numerical model was able to reproduce the experimental scratch features, cutting forces and deformed layer thickness formed on the carbon steel, for all angular scratches evaluated in this work;

• Cutting forces, burr formation and subsurface deformed layer values indicated that the 20°case was the most critical. The material strainhardening was more accentuated when a 20°angle was used. This outcome is corroborated with a higher plastic dissipation energy as well as a lower removed volume for the 20°condition, when compared to the others;

• Conversely, considering minimum and maximum scratch angles (10°a nd 30°), it was noticed that the cutting forces showed the trend to decrease for 10°. Moreover, the plastic dissipation energy and removed material analyses indicated 30°case as an intermediary condition between 10°and 20°angular scratches, which agrees with the deformed layer thickness evaluation (material mechanical properties) and burr removal (ductile damage).

As a final remark, it can be said that the approach implemented here is of technological importance to design and reproduce surfaces manufactured by machining processes. The numerical procedure and the experimental characterizations and measurements are essential to understand the machined surface and to design processes for its modification.

Table 6

Numerical results in terms of some parcels of the dissipated energy during the cutting.

Fig. 1 .Fig. 2 .

 12 Fig. 1. Experimental setup of the scratch tests onto the 1020 steel conducted in the sequence a-d. Parallel scratches were carried out (a-c) and later a second set of parallel scratches was run at a specific angle (10, 20 or 30°) with respect to the previous one (d).

Fig. 3 .

 3 Fig.3. Typical graph of the experimental cutting forces x time in the following conditions: without and with signal treatment (denoise). According to Fig.1d, the angular scratch crossed some previous scratches. Thus, the points (a) and (b) show the influence of these events in the experimental cutting forces.

Fig. 4 .

 4 Fig. 4. Average experimental cutting forces obtained from the different angular scratch conditions.

ig. 5 .

 5 Experimental results: topography characterization for all orientations. The burr featureslength, width and height -are also displayed here.

ig. 6 .

 6 Numerical results: Equivalent plastic strain at the onset of the fracture, defined by the Johnson-Cook damage criterion (JCCRT) for all orientations. The burr featureslength, width and heightobtained from the numerical analyses are available here.

Fig. 7 .

 7 Fig. 7. Numerical results of the material removal and plastic dissipation energy as a function of the orientation of the angular scratch.

  (a) OM -Angle 20° example (b) SEM: Angle 10°( c) SEM: Angle 20°(d) SEM: Angle 30°F ig. 8. Characterization by optical and scanning electron microscopies (OM and SEM) of the cross sections at the intersection of two consecutive scratches with different orientations onto the AISI 1020 steel: (a) OMexample for angle 20°; SEM -(b) angle 10°, (c) angle 20°; and (d) angle 30°.

Table 1

 1 Johnson-Cook plasticity parameters of the AISI 1020 steel[START_REF] Kui | The dynamic characteristics and constitutive model of 1020 steel[END_REF].

	A [MPa]	B [MPa]	C	n	m	Tm [°C]	Tr [°C]	ε0 [s -1 ]
	213	53	0.055	0.345	0.81	1350	20	0.004

Table 4

 4 Experimental and numerical results of the tangential, axial and radial forces obtained from the angular scratches. Exp. = experimental, and Num.= numerical.

	Condition Tangential force [N]	Axial force [N]	Radial force [N]
	Exp.	Num.	Exp.	Num.	Exp.	Num.
	10°228 ± 14	220.95	8.7 ± 1.8	7.20	217 ± 9	248.14
	20°385 ± 16	313.04	14.2 ± 1.3 13.13 365 ± 10 346.17
	30°267 ± 68	249.96	10 ± 4	8.42	296 ± 45 278.83

Table 5

 5 Scratch features calculated and normalized according with the schematic representation.

Acknowledgments

The authors would like to acknowledge the Brazilian research support agencies: Sao Paulo Research Foundation (FAPESP) Grant number 2016/18014-6, National Council for Scientific and Technological Development (CNPq), and Foundation for the Support of University of Sao Paulo (FUSP).

Fig. 9. Indication of the cross section evaluated throughout numerical and experimental characterizations (a). In addition, the equivalent plastic strain, obtained from the numerical simulation, is shown along the subsurface layer of the material (b).


Chao Li

Jinyang Xu
email: xujinyang@sjtu.edu.cn

Ming Chen

Qinglong An

Mohamed El Mansori

Fei Ren

Tool wear processes in low frequency vibration assisted drilling of CFRP/ Ti6Al4V stacks with forced air-cooling

Keywords:

Tool wear processes in low frequency vibration assisted drilling of CFRP/Ti6Al4V stacks with forced air-cooling

Introduction

The hybrid structure of composites and metal alloys has been widely used in the aerospace industry for the purpose of reducing the weight of airframes and of improving the structural strength [START_REF] Park | Tool wear in drilling of composite/ titanium stacks using carbide and polycrystalline diamond tools[END_REF]. Combination of the single materials with individual advantages generates enhanced properties including high strength-to-weight ratio, excellent design flexibility and high erosion/corrosion resistance which are much better than the use of each individual material [START_REF] Xu | Wear characteristics of polycrystalline diamond tools in orthogonal cutting of CFRP/Ti stacks[END_REF][START_REF] Xu | Experimental studies on the cutting characteristics of hybrid CFRP/Ti stacks[END_REF]. Carbon fiber reinforced polymer (CFRP) and Ti6Al4V alloy are two representative materials widely used in the structural stacks which are often assembled by rivets or bolts. Drilling of the bi-material sandwich in one-shot time is the effective method applied in the actual production to produce boreholes with high assembly precision and high productivity. However, drilling of composite/metal stacks with desirable quality is rather challenging due to the disparate machinability of each constituted material [START_REF] Kim | Drilling process optimization for graphite/bismaleimide-titanium alloy stacks[END_REF]. In particular, the highly abrasive carbon fibers and the chemically-reactive titanium alloy during drilling result in harsh cutting environments with coupling effects of heavy mechanical and thermal loads to the drills. Problems of excessive tool wear and short tool life need to be carefully addressed.

Chip removal process in cutting CFRP/Ti6Al4V stacks is of fundamental importance for understanding the wear mechanisms of tools. As claimed by Xu and El Mansori [START_REF] Xu | Wear characteristics of polycrystalline diamond tools in orthogonal cutting of CFRP/Ti stacks[END_REF], due to the brittle-fracture dominated chip separation mode, "powder" like chips form in machining CFRP phase while continuous chips are generated because of the elastoplastic deformation of Ti6Al4V in the cutting process. The powdery chips of CFRP can be harmful not only from the view of human health but also from the environmental perspective. Additionally, in a typical process of drilling CFRP/Ti6Al4V stacks under constant feed rates, the continuous spiral chips generated in cutting the metallic phase cannot eject effectively. The residual titanium chips with inability to evacuate can get entangled at the interaction zones between the drill edges and the workpiece causing the excessive tool wear and the dramatic increase of cutting forces and temperatures. As experimentally studied by Brinksmeier et al. [START_REF] Brinksmeier | Drilling of composites and resulting surface integrity[END_REF], the conventional drilling (CD) brings more thermal loads caused by the higher friction between the drill clearance face and the borehole wall which results in the damage to surface integrity and the adhesive wear of tools. Xu and El Mansori [START_REF] Xu | Experimental study on drilling mechanisms and strategies of hybrid CFRP/Ti stacks[END_REF] pointed out that the titanium chip evacuation has certain influences on the drilling forces. Thus, in order to achieve the reduction of health hazard and the tool wear, the vibration-assisted machining (VAM) and cutting fluids are proposed from the aspects of reducing mechanical loads and of providing better chip removal ability.

As discussed in the review article published by Brehl and Dow [START_REF] Brehl | Review of vibration-assisted machining[END_REF], the fundamental difference between conventional machining and VAM is the periodic contact between the tool rake face and the workpiece material, which leads to the interrupted generation of chips. According to the range of the vibrational frequency, the VAM can be divided into ultrasonic machining (USM, usually frequency ≥ 20 kHz) [START_REF] Thoe | Review on ultrasonic machining[END_REF] and low frequency vibration assisted machining (usually frequency < 1000 Hz) [START_REF] Liu | A review of mechanical drilling for composite laminates[END_REF]. USM can be utilized as a non-thermal process with relatively small amplitudes of about 10-20 µm to deal with materials with hardness above 40 HRC [START_REF] Singh | Ultrasonic machining of titanium and its alloys: a review[END_REF]. However, the material removal rate of USM limited by its small amplitudes becomes the key drawback in the industry for the drilling of CFRP/Ti6Al4V stacks [START_REF] Pecat | Tool wear analyses in low frequency vibration assisted drilling of CFRP/Ti6Al4V stack material[END_REF]. On the contrary, as revealed by Kuo et al. [START_REF] Kuo | Multi-objective optimisation in vibration-assisted drilling of CFRP/Al stacks[END_REF], the LFVAD with amplitudes of about 0.05 mm is practical and feasible in both reducing drill flank wear and improving the processing efficiency. The cutting fluids are commonly used to serve as a medium to take away the chips and to provide cooling in the drilling processes [START_REF] Meena | Study of dry and minimum quantity lubrication drilling of novel austempered ductile iron (ADI) for automotive applications[END_REF]. Ishikawa et al. [START_REF] Ishikawa | A study on combined vibration drilling by ultrasonic and low-frequency vibrations for hard and brittle materials[END_REF] pointed out that the decreased drilling forces and tool wear could be attributed to the discharge of cutting chips taken away by the application of working fluids during the LFVAD. However, the cyclic utilization of cutting fluids in the machining of CFRP is uneconomic and environmentally unfriendly due to the heavy pollution of the powdery chips.

The overall tool wear processes in drilling CFRP/Ti6Al4V stacks are a combination of coupling effects of heat and tribology when cutting the bi-material sandwich. Generation and transfer of heat in machining fiber reinforced composites concern not only the thermal defects but also the tool wear which should be carefully measured and studied [START_REF] Yashiro | Temperature measurement of cutting tool and machined surface layer in milling of CFRP[END_REF][START_REF] Xu | Recent advances in drilling hybrid FRP/Ti composite: a state-of-the-art review[END_REF]. In order to capture the in-situ cutting temperatures, applications of thermocouples embedded into the drills through the coolant holes, dynamic thermocouples between the tool/workpiece and infrared thermography camera are the three most-used measuring methods [START_REF] Yashiro | Temperature measurement of cutting tool and machined surface layer in milling of CFRP[END_REF][START_REF] Merino-Pérez | On the temperatures developed in CFRP drilling using uncoated WC-Co tools Part II: nanomechanical study of thermally aged CFRP composites[END_REF].

Tool wear processes of LFVAD are related to the interrupted chips generated by the periodic separation and contact between the drill rake face and the workpiece material. As reported by Pecat and Meyer [START_REF] Pecat | Low frequency vibration assisted drilling of aluminium alloys[END_REF], the shape and radian of undeformed chip segments depend on the cutting parameters, vibration frequency, vibration amplitude, etc. In the LFVAD of CFRP/Ti6Al4V stacks, although the continuous spiral chips are sheared into segments, the hot and sharp metallic chips can still be capable of clogging the drilled holes because of the improper chip shapes induced by inappropriate cutting parameters and the poor chip removal ability. Thus, the thermal damage to the drilled CFRP holes and the abrasive wear to tools may remain severe.

In order to deal with the tool wear issues by efficiently evacuating the residual titanium chips when drilling CFRP/Ti stacks under the low frequency vibration assistance, in this work, a novel integrated system of the LFVAD coupled with forced air-cooling was designed. The forced air-cooling system provides a strip of rectangular area with negative pressure close to the cutting position which can suck away the chip segments during drilling processes ensuring better chip removal ability and cooling effects. Drilling temperatures were in-situ measured by embedded thermocouples and infrared thermography camera in the LFVAD subjected to varying conditions, and the results were compared with those gained under the conventional drilling. Special attention was paid to the issues of tool wear in the LFVAD of CFRP/Ti6Al4V stacks using uncoated twist drills with/without forced air-cooling. This paper makes efforts to improve the performance of the LFVAD of CFRP/Ti6Al4V stacks which is achieved through a modified cooling equipment and systematic experimental investigations taking drilling forces, cutting temperatures and tool wear into account.

Experimental procedures

Composite/metal stacks and drill bits

The material sandwich investigated in this work consists of the CFRP and Ti6Al4V boards. The specimen has a total size of 300 mm (length) × 200 mm (width) × 17.16 mm (thickness, 9.16 mm for the CFRP and 8 mm for the titanium alloy) as depicted in Fig. 1. In the study, the aeronautical grade CFRP laminate was fabricated by carbon/ epoxy prepregs containing T800 reinforced fibers with volume fraction of 65% and the matrix of X850 epoxy resin. The detailed mechanical/ physical properties of the stacked Ti6Al4V alloy and T800/X850 composites are summarized in Table 1.

Threaded shank drills were specially designed in this work to fit the adapter of the tool holder used in the LFVAD. As shown in Fig. 1, the uncoated solid tungsten carbide drill bits with coolant holes provided by Fengqi Machinery Technology Co., Ltd. were utilized. The main geometrical features of the used drills include a 6.35 mm diameter, a 140°point angle, a 30°helix angle and two cutting edges.

Experimental setup of LFVAD with forced air-cooling

As depicted in Fig. 2, all the experiments were carried out on a HURCO VMX42 machining center with a maximum spindle speed of 12,000 rpm and a positioning accuracy of 0.01 mm. With a speciallydesigned fixture, the CFRP/Ti stacks and the forced air-cooling system were clamped on a dynamometer (Kistler 9272) connected with a multichannel charge amplifier (Kistler 5070A) during the drilling operation. In addition, the cutting forces were recorded at an acquisition frequency of 20 kHz. The forced air-cooling system consists of a chip collection unit equipped with a strip of rectangular suction nozzle connected to a tube and a vacuum pump which can provide the negative pressure to finalize the in-situ collection of resected chips.

The axial low frequency oscillations were generated by a mechanical structure (equipment reference: PP3020) consisting of the roller, flat ring and sine ring provided by MITIS Engineering Company. Amplitudes ranging from 0.10 to 0.35 mm and oscillation containing 1.5 and 2.5 osc/rev can be obtained by changing the sine rings. In the present study, a fixed frequency of 1.5 osc/rev (N v = 1.5 osc/rev) and an amplitude of 0.20 mm (A = 0.20 mm) were applied. The tool holder was modified to reduce the effects of the assembly errors induced by the threaded connection structure, and the radial run-out of the drill bits was maintained within 8 µm.

In order to identify how the LFVAD equipped with the forced aircooling affects the tool wear in the drilling of CFRP/Ti6Al4V stacks, three groups of experimental tests were performed in this work as summarized in Table 2. Additionally, the drilling temperatures (T) were measured by both embedded thermocouples (T e ) and infrared thermography camera (T f ) in test group 1 to identify the functional correlations of these two temperature measuring methods guiding the rest of the two test groups. All the drilling experiments were carried out in dry machining conditions.

In-situ measurement of the drilling temperatures

Due to the structural constraints of the modified tool holder, conventional embedded thermocouples inside the drill bits are not suitable for the in-situ measurement of the drilling temperatures in the LFVAD. Thus, an infrared thermography camera was used to realize the in-situ measurement of the drilling temperatures under various cutting conditions. As revealed in the relevant research [START_REF] Yashiro | Temperature measurement of cutting tool and machined surface layer in milling of CFRP[END_REF], the temperatures measured by the infrared thermography camera tend to be lower than those captured by the embedded thermocouples due to the hidden cutting points. To obtain the realistic development tendency of the drilling temperatures, the measuring method of embedding thermocouples inside the drill bits was applied to serve as the calibration and confirmation in order to improve the measuring accuracy.

The overview of the temperature measurement is given in Fig. 3. All the holes were drilled at a distance of 2 mm close to the edge of the CFRP/Ti stack to make the temperature measurements more reliable. The standard OMEGA K-type thermocouples were embedded into the drill bits through the internal coolant holes. The FLIR A615 infrared thermography camera (IFTC) which has a working temperature ranging from -20 to 2000 °C and an image acquisition frequency ranging from 50 Hz to 200 Hz was applied to in-situ measure the drilling temperatures. The temperature resolution of the equipment is less than 0.05 °C which ensures the accuracy and comparability of the monitored data.

Characterization methods

In order to investigate the tool wear issue in the LFVAD of CFRP/ Ti6Al4V stacks with the forced air-cooling, several indicators including the drilling forces, cutting temperatures and tool wear were analyzed. Processes of the worn surface morphologies of the used drills in the drilling tests were observed by a high-precision digital microscope (Keyence VHX-500FE) and further inspected using both the scanning electron microscope (SEM) and the energy dispersive spectroscopy (EDS). All the measurements were repeated three times in order to get reliable results.

Results and discussion

Cutting temperatures and chip removal processes

The process temperatures are of great importance to the tool wear in dry drilling CFRP/Ti stacks. However, the temperature measurement in the drilling process especially in the LFVAD is quite difficult because of the hidden cutting points. To realize the in-situ measurement of drilling temperatures in the LFVAD with high precision, both the infrared thermography camera and the embedded thermocouples into the drill bits were employed in the conventional drilling (CD) to serve as the references to the individual measuring method of the infrared thermography camera used in the LFVAD. The comparative results of process temperatures obtained by the thermocouples and IFTC in the conventional drilling (V = 20 m/min and f = 0.015 mm/rev) are plotted in Fig. 4.

It is noted that for the same drilling process, temperatures of the cutting zone measured by IFTC are lower than those obtained by the thermocouples at the stable stage which is mainly caused by the poor thermal conductivity of both the CFRP and titanium alloy. Although there exists a time delay and interferential temperature points caused by the evacuation of the hot chips as depicted in Fig. 4(a), the temperatures recorded by the IFTC show the same varying trend as that obtained by the thermocouples, which means that the temperature measurement via IFTC is reliable and convincing.

Fig. 5 presents the thermal images of the drilling temperatures when using different drilling methods (CD, LFVAD with and without the forced air-cooling) subjected to the cutting conditions of V = 20 m/ min, f = 0.015 mm/rev, A = 0.20 mm and N v = 1.5 osc/rev. In the conventional drilling, the continuous spiral chips can be clearly observed which lead to the highest temperature of 316.65 °C. By contrast, the use of the LFVAD brings a significant decrease to the drilling temperatures of about 47.9% while the LFVAD with the forced aircooling leads to a dramatic temperature decrease of about 51.3%. Thanks to the periodical movement of the cutting edge, continuous spiral Ti chips were broken into chip segments and free motion of chip segments can be clearly observed during the vibration assisted drilling in Fig. 5. The proper shapes of chip segments do favor the evacuation of heat generated by drilling. Furthermore, it is shown that the forced aircooling system can absorb the chip segments timely during the vibration assisted drilling which enhances the ejection process of chips leading to a sharp decrease of the temperatures.

The results of in-situ measurement of the drilling temperatures obtained in the conventional drilling, LFVAD with and without the forced air-cooling under the cutting conditions of V = 20 m/min, f ranging from 0.015 to 0.060 mm/rev, A = 0.20 mm and N v = 1.5 osc/rev are given in Fig. 4(b). It can be noted that the vibration assisted drilling can bring reductions of cutting temperatures efficiently under the low feed rates of 0.015 and 0.030 mm/rev while the temperatures rise significantly being higher than those gained in the conventional drilling under the high feed rates of 0.045 and 0.060 mm/rev. This is caused by the large and thick chip segments produced by the high feed rates coupled with the amplitude of cutting edge. Thus, it can be assumed that there exists a critical size of the chip segments beyond which the effectiveness of the chip evacuation would be worse than the conventional drilling. Additionally, the forced air-cooling system can make contributions to the efficient ejection of the chip segments and then lead to the reduction of abrasive wear caused by the Ti adhesion under the high cutting temperatures. It can be concluded that, to take advantages of the vibration assisted drilling, the coupling effects of the vibration frequency, feed rate and amplitude should be carefully considered.

Cutting forces

Analysis of drilling forces is the fundamental step in understanding the effects of tool kinematics on the wear of drills. Signals of thrust forces in the time domain with respect to different cutting conditions are shown in Fig. 6 for the case corresponding to V = 20 m/min, f = 0.015 mm/rev, A = 0.20 mm and N v = 1.5 osc/rev. The results show that the local thrust force fluctuations in drilling CFRP are bigger than those gained in drilling Ti6Al4V due to the different machining characteristics of the two materials. For the anisotropic machinability of CFRP, the unstable process of chip formation results from the dynamic changes of the angle between the cutting edge and fibers leading to the random friction on the tools which is externally expressed by the local force fluctuations. In the stable cutting stage of drilling Ti6Al4V, the thrust forces of the conventional drilling maintain around 150 N except for a sudden fluctuation caused by the clogging spiral Ti chips. In contrast, for the LFVAD with and without the forced air-cooling, no significant alterations of thrust force are observed, which means the efficient ejection of the Ti chips. Furthermore, the drilling process of the LFVAD with the forced air-cooling generates thrust forces of about 25 N lower than the LFVAD without the forced air-cooling. Since the residual chips can lead to higher cutting resistance per revolution, the reduction of thrust forces represents the assistance of the forced air-cooling to remove the residual chips, which makes contributions to the reduction of tool wear caused by the heavy mechanical loads.

The thrust force in the drilling process commonly signifies the formation of chips and the tribological interaction at the tool-work interface. The nearly constant thrust force of the conventional drilling at the stable cutting stage represents the formation of continuous spiral chips with nearly the same chip thickness, while for the LFVAD, the thrust force fluctuates in a certain degree ranging from the minimum (0 N) to the maximum (~280 N). Although the LFVAD produces a much higher mechanical load, the actual cutting time is reduced, and the average cutting forces are lower than those of the conventional drilling. According to the change of the thrust force, the formation processes of chip segments can be analyzed: every two adjacent wave troughs represent the two pieces of separated chips. In fact, the vibration frequency (F v ) of the thrust force can be calculated using the following equation, where n is the spindle speed in rpm and N v is the oscillation frequency in osc/rev:

= F n N 60 v v (1)
The regularly broken Ti chips produced by the low frequency vibration (25 Hz) assisted drilling cause less contacts between the cutting edge and the workpiece compared with the conventional drilling. As a result, the tool wear caused by the durable contacts with long distance chips can be reduced remarkably. Additionally, it appears that the forced air-cooling leads to much lower thrust forces in drilling Ti6Al4V assisted with the low frequency vibration. However, the effects of reducing thrust forces are insignificant in the drilling of CFRP. This is attributed to the disparate mechanisms of chip formation in the conventional drilling of CFRP and Ti6Al4V. Cutting chips of CFRP exist in the form of dust while continuous spiral chips are the dominant chip type in the conventional drilling of Ti alloys. The forced air-cooling system can effectively promote the evacuation of chip segments and the continuous chips lead to both the significant reduction of the thrust forces and the less local force fluctuation.

Tool wear

Tool wear occurring in drilling is induced by the coupling effects of thermal and mechanical loads which is further enhanced by the combination of abrasive fibers and the difficult-to-machine titanium alloy in the CFRP/Ti stacks. Localized dynamic loads in the cutting zone are modified during the drilling assisted by low frequency vibration and forced air-cooling, which leads to the changes of the cutting and wear mechanisms. In order to understand the tool wear process in the LFVAD with the forced air-cooling, SEM inspections and EDS analyses were conducted, and the results are shown in Figs. 7 and8 respectively. All of the three test groups of drilling methods were investigated after drilling an identical number of 45 holes under the fixed cutting conditions of V = 20 m/min, f = 0.015 mm/rev, A = 0.20 mm and N v = 1.5 osc/rev.

It is noted that the adhesive wear is the main wear pattern and the edge fracture operates as the key drill failure mode in the conventional drilling. Large titanium adhesions can be observed along the cutting edge with a large width. This can be attributed to the generation of the clogging Ti chips derived from the continuous process of chip formation in the conventional drilling corresponding to the sudden alteration of thrust force signals shown in Fig. 6. The constant contact between the drill flank face and the workpiece leads to the heavy thermal loads and high temperatures due to the low thermal conductivity of the titanium alloy. In contrast, the edge chipping which leads to the most severe reduction of tool life is reduced by the use of the LFVAD with the suitable cutting parameters.

One of the advantages of the vibration assisted drilling is the intermittent cutting process in which the discontinuous titanium chips are periodically produced. The absence of continuous chips provides a better environment for the chip evacuation and a possible reduction in coefficient of the sliding friction which signifies the elimination of unstable cutting stages. Thus, less abrasive wear and edge fracture constitute the combined wear mode in the vibration assisted drilling of CFRP/Ti stacks compared with the conventional drilling.

Furthermore, a large extent of discolored area can be identified on the drill flank surface which signifies the enhancement of the chemical wear in the LFVAD with the forced air-cooling instead of the edge fracture. Meanwhile, a significant reduction of titanium adhesions can be observed. This can be evidenced by the presence of various oxidizing particles which are detected via the EDS analysis shown in Fig. 8. Compared with the conventional drilling, the LFVAD provides not only small titanium chip segments but also the cooling effects to the cutting zone by enhancing the local air motion via the help of the forced aircooling system. On the one hand, slower flank wear rates can be achieved by the reduction of the cutting temperatures while on the other hand, heavier oxidation wear becomes pronounced. This phenomenon can be attributed to the air exchange occurring at the cutting zone during drilling via the help of the forced air-cooling system. The flowing air provides the cooling effects to the drilling process resulting in a lower tool wear rate while the air with oxygen will degrade the contact surface under high drilling temperatures. Thus, the discolored area signifying the chemical wear can be observed from the results of SEM depicted in Fig. 7.

In this work, tool wear was measured as the average flank wear land width (VB), as a function of number of drilled holes. The results are shown in Fig. 9. From this figure, it is clear that the tool wear processes undergo the initial, normal and rapid wear stages for all the test groups. In the initial wear stage, the interrupted contact between the cutting edge and the workpiece provides the drill bit with higher wear resistance under the vibration assisted drilling. Furthermore, the forced air-cooling provides not only the enhancement of the chip evacuation ability but also the cooling effects to the drilling processes which leads to a longer stable normal wear stage. As expected, the results confirm the feasibility of the vibration assisted drilling especially combined with the use of the forced air-cooling in reducing the extents of drill flank wear. The rising tendency of the flank wear indicates that the use of the vibration assisted drilling with forced air-cooling could effectively slow down the growth rate of the flank wear when proper cutting parameters are adopted.

Conclusions

In this paper, an integrated system of the LFVAD coupled with forced air-cooling was designed and applied to investigate the tool wear issue in drilling of CFRP/Ti6Al4V stacks using uncoated twist drills. The work concerns the effects of the LFVAD with the forced air-cooling on the tool wear process through the analysis of the drilling-induced temperatures and forces at the cutting zone. The flank wear was quantified to estimate the tool life with respect to different drilling methods. The results gained in the present work reveal some implicit mechanisms of the temperature and tool wear issues in the vibration assisted drilling. Based on this study, the following conclusions are • Compared with the embedded thermocouples, there exists a time delay and overall reductions of temperatures measured by the infrared thermography camera. The use of the vibration assisted drilling brings a significant decrease to the drilling temperatures of about 47.9% while the dramatic temperature decrease of about 51.3% can be achieved via the help of the forced air-cooling at the feed rate of 0.015 mm/rev.

• The small chip segments produced by the periodical cutting of the LFVAD can be efficiently evacuated via the help of the forced aircooling and better cooling effects can be achieved at the meanwhile. The forced air-cooling can lead to much lower thrust forces in drilling Ti6Al4V by the LFVAD while its effects on reducing the thrust forces of the CFRP phase are insignificant due to the dissimilar chip formation mechanisms of the two materials. The separated chips of the carbon/epoxy composite exist mainly in the form of dust while continuous spiral chips are the dominant chip type in the conventional drilling of Ti alloys.

• In contrast with the conventional drilling, the LFVAD with the forced air-cooling facilitates the decrease of the titanium adhesion and the flank wear on the drills and then prolongs the tool life. Furthermore, the oxidation wear becomes more pronounced due to the activated local air movement which indicates that the dominating wear mechanisms are changed.

• In sum, the LFVAD coupled with the forced air-cooling is confirmed capable of improving the machinability of CFRP/Ti6Al4V stacks from the aspect of reducing tool wear. However, to improve the chemical wear resistance, coated drills can be adopted to reduce the diffusion of chemical elements between the tool and the chips. Moreover, the coupling effects of the vibration frequency, feed rate and amplitude should be carefully

Fig. 1 .

 1 Fig. 1. Photographs showing (a) the utilized multilayer CFRP/Ti6Al4V stack and (b) the uncoated carbide threaded shank drill.

Fig. 2 .

 2 Fig. 2. Photographs showing the experimental setup. Right is the tool holder for the LFVAD and the components of the source of axial oscillations (A = 0.20 mm and N v = 1.5 osc/rev) and below is the designed forced air-cooling system.

Fig. 3 .

 3 Fig. 3. Schematic view showing the in-situ measurement of drilling temperature by (a) the infrared thermography camera and (b) the thermocouples.

Fig. 4 .

 4 Fig. 4. (a) Evolution of the drilling temperatures in the conventional drilling (CD) with the cutting time (V = 20 m/min and f = 0.015 mm/rev) and (b) comparison of the drilling temperatures when using different drilling methods (CD, LFVAD with and without the forced air-cooling) (V = 20 m/min, f = 0.015, 0.030, 0.045, 0.060 mm/rev, A = 0.20 mm and N v = 1.5 osc/rev).

Fig. 5 .

 5 Fig. 5. Thermal infrared images showing the drilling processes when using different drilling methods (V = 20 m/min, f = 0.015 mm/rev, A = 0.20 mm and N v = 1.5 osc/rev).

Fig. 6 .

 6 Fig. 6. Comparison of signals of the thrust forces gained when using different drilling methods (V = 20 m/min, f = 0.015 mm/rev, A = 0.20 mm and N v = 1.5 osc/ rev).

Fig. 7 .

 7 Fig. 7. SEM morphologies of the worn drill surfaces obtained under different test groups.

Fig. 8 .

 8 Fig. 8. Results of comparison of the EDS analysis on the drill flank surfaces after the completion of three groups of drilling tests.

Fig. 9 .

 9 Fig. 9. Progression of the drill flank wear with the number of drilled holes under different test groups (V = 20 m/min, f = 0.015 mm/rev, A = 0.20 mm and N v = 1.5 osc/rev).

Table 1

 1 Material properties of the studied T800/X850 CFRP and Ti6Al4V alloy.

	Properties of T800/X850 CFRP	Density [g/cm 3 ]	Fiber volume fraction	Tow
		1.60	65%	5 µm, 12 K
		Tensile modulus [GPa]	Compressive modulus [MPa]	Tensile strain [%]
		180	160	1.68
	Properties of Ti6Al4V alloy	Density [g/cm 3 ]	Tensile strength [MPa]
		4.43	90-1160
		Thermal conductivity [W/(m °C)]	Specific heat [J/(kg•°C)]
		7.0	546

Table 2

 2 Details of the utilized parameters in drilling CFRP/Ti6Al4V stacks.

	Test group 1 (general cutting parameters)	Test group 2	Test group 3
	Spindle	Cutting	Feed rate,	Additional	Additional
	speed, n	speed, V [m/	f [mm/	conditions	conditions
	[rpm]	min]	rev]
	1000	20	0.015	Amplitude	Forced air-
			0.030	0.20 mm	cooling
			0.045	Oscillation
			0.060	1.5 osc/rev

Acknowledgements

The authors gratefully acknowledge the financial support of National Natural Science Foundation of China (Grant no. 51705319), Shanghai Pujiang Program (Grant no. 17PJ1403800), and Shanghai Academy of Spaceflight Technology (Grant no. SAST2017-060). The work is also partly sponsored by the Research Project of State Key Laboratory of Mechanical System and Vibration (Grant no. MSVZD201704).