micromachines
Article Serrated Chips Formation in Micro Orthogonal Cutting of Ti6Al4V Alloys with Equiaxial and Martensitic Microstructures
ZeJia Zhao, Suet To * and ZhuoXuan Zhuang State Key Laboratory of Ultra-Precision Machining Technology, Department of Industrial and Systems Engineering, The Hong Kong Polytechnic University, Hung Hom, Kowloon, Hong Kong SAR, China; [email protected] (Z.J.Z.); [email protected] (Z.X.Z.) * Correspondence: [email protected]; Tel.: +852-2766-6587
Received: 27 February 2019; Accepted: 15 March 2019; Published: 20 March 2019
Abstract: The formation of serrated chips is an important feature during machining of difficult-to-cut materials, such as titanium alloy, nickel based alloy, and some steels. In this study, Ti6Al4V alloys with equiaxial and acicular martensitic microstructures were adopted to analyze the effects of material structures on the formation of serrated chips in straight line micro orthogonal machining. The martensitic alloy was obtained using highly efficient electropulsing treatment (EPT) followed by water quenching. The results showed that serrated chips could be formed on both Ti6Al4V alloys, however the chip features varied with material microstructures. The number of chip segments per unit length of the alloy with martensite was more than that of the equiaxial alloy due to poor ductility. Besides, the average cutting and thrust forces were about 8.41 and 4.53 N, respectively, for the equiaxed Ti6Al4V alloys, which were consistently lower than those with a martensitic structure. The high cutting force of martensitic alloy is because of the large yield stress required to overcome plastic deformation, and this force is also significantly affected by the orientations of the martensite. Power spectral density (PSD) analyses indicated that the characteristic frequency of cutting force variation of the equiaxed alloy ranged from 100 to 200 Hz, while it ranged from 200 to 400 Hz for workpieces with martensites, which was supposedly due to the formation of serrated chips during the machining process.
Keywords: serrated chips; Ti6Al4V alloy; micro orthogonal machining; martensite; cutting and thrust forces
1. Introduction To date, Ti6Al4V alloy is increasingly applied within the aerospace field due to its low weight, superior strength, and corrosion resistance [1,2]. Generally, the mechanical properties of the Ti6Al4V alloy are determined by its material microstructures, mainly including phase composition and grain size. Equiaxial and martensitic microstructures are two important types of structures for Ti6Al4V alloy [3]. An alloy with martensite can be obtained through high-temperature annealing (above β phase transformation temperature) followed by water quenching. The two-phase alloy with equiaxial grains usually has a balance of strength, ductility, and corrosion resistance, while alloy with martensite generally possesses high strength and corrosion resistance but low ductility [4]. Ti6Al4V alloy is regarded as one of the difficult-to-machine materials due to the low thermal conductivity and elastic modulus in traditional and ultraprecision machining [5–7]. Serrated chip (or saw tooth chip) formation is an important feature during the machining of Ti6Al4V alloys. The generation of serrated chips is totally different from the formation of continuous chips that are usually obtained in cutting most easy-to-machine materials, such as copper, aluminum, and most
Micromachines 2019, 10, 197; doi:10.3390/mi10030197 www.mdpi.com/journal/micromachines Micromachines 2019, 10, 197 2 of 11 of their alloys [8–10]. The chips can take away the majority of generated heat because of the superior thermal conductivity of easy-to-cut materials. However, heat is not prone to being dispersed during the machining hard-to-cut alloys because of their poor thermal conductivity or fracture properties, resulting in serrated chip formation according to the thermal softening effect or periodical crack generation [11–13]. Two theories, that is, adiabatic shear band theory and periodical cracks formation theory, are the most widely accepted mechanisms for analyzing the generation process of serrated chips. Each theory works in a certain condition, and no consensus has been achieved on serrated chip formation in the machining of Ti6Al4V alloys. Komanduri and Von Turkovich [14] proposed that the generation of an adiabatic shear band gives rise to the formation of serrated chips in the high-speed machining of titanium alloys. If thermal softening predominates over strain hardening, shear localization can occur in the primary cutting zone. This contributes to a reduction of the required stress to overcome plastic deformation of a workpiece [15]. Sun et al. [16] proposed that severe plastic deformation in the narrow shear band results in the formation of serrated chips, and also found the formation of serrated chips could induce a periodical cutting force evolution. The periodical crack initiation and propagation between two segments of one serrated chip is another theory to investigate the chip-forming mechanism. Vysa and Shaw [17] reported the detailed crack generation process in the orthogonal cutting of difficult-to-cut AISI 1045 steel. Jiang and Rajiv [18] discussed that the serrated chip morphology was primarily caused by crack initiation and propagation during the machining of Ti6Al4V alloy. In addition, the formation of serrated chips was successfully simulated by using various models, including the Johnson–Cook (JC) law [19,20], modified JC law [21,22] and tangent hyperbolic (TANH) model [23]. Besides, the JC damage law and ductile fracture criterion [24,25] were adopted as criteria to analyze the fractural characteristic in discussing the saw tooth chip formation of titanium and its alloys. Material microstructures also significantly affect the machinability of a workpiece during the machining process. Nouari and Makich [26,27] reported that tool wear was greatly influenced by the microstructures of titanium alloys. Ahmadi et al. [28] proposed that small grain size and low content of β phase could result in high cutting forces in the micro milling of Ti6Al4V alloys. Though some research has been carried out on studying the effects of microstructures on tool wear and cutting forces in the traditional machining of Ti6Al4V alloys, little research has been conducted on comparing the chip features from the perspective of material microstructures of the Ti6Al4V alloys in micro cutting. As strength and ductility vary between microstructures, the machinability of the Ti6Al4V alloys should also show different characteristics that can be indirectly reflected from the features of cutting chips. Hence, this research aims to investigate the effects of equiaxial and martensitic microstructures on the serrated chip formation of Ti6Al4V alloys via micro orthogonal machining.
2. Experimental Procedures Ti6Al4V alloy with a diameter of 3 mm was utilized as the raw material. The alloy was first cut into six pieces with a length of about 10 mm and half of these were then treated for 2 min using an electropulsing device (THDMIII). The voltage and frequency of the electropulsing were 50 V and 400 Hz, respectively. After the electropulsing treatment (EPT), samples were quenched immediately in water (25 ◦C). Subsequently, all EPT-treated and original samples were cut into rectangular shapes using wire electrical discharge machining (WEDM) to then carry out micro orthogonal machining. The dimensions of the workpiece are shown in Figure1c. An ultraprecision machine (Moore Nanotech 350FG, Nanotechnology Inc, Swanzey, NH, USA) was used for the micro orthogonal cutting of the Ti6Al4V workpiece and the setup is schematically shown in Figure1a,b. The head width, rake angle, and front clearance angle of the fresh flat diamond tool were 2 mm, 0◦, and 10◦, respectively. Rough machining with another diamond tool (nose radius of 1.103 mm, rake angle of 0◦, and front clearance angle of 12.5◦) was first carried out on the workpiece to obtain a flat surface on the ultraprecision machine. Then, micro orthogonal machining with a Micromachines 2018, 9, x FOR PEER REVIEW 3 of 11
Micromachines 2019, 10, 197 3 of 11 was carried out on the workpiece, and machining was stopped at the cutting length of 6 mm. Figure 1d briefly illustrates the details of the micro orthogonal machining process. A Kistler force sensor
(9256C1cuttingMicromachines width, Kistler of 2018 1 Corporation, mm,, 9, x FOR cutting PEER speedREVIEW Winterthur, of 60 mm/min, Switzerland and) cutting was installed depth of 3 underµm was the carried cutting3 out of tool11 on the to measureworkpiece, variation and machining in cutting was and stopped thrust at theforces. cutting The length measured of 6 mm. frequency Figure 1 wasd briefly set as illustrates 100 kHz, the so frequenciesdetailswas of carried the lower micro out orthogonalthanon the 50workpiece kHz machining could, and machining be process. accurately was A Kistler stopped calculated force at the sensor incutting the (9256C1, powerlength Kistlerof spectrum 6 mm. Corporation, Figure density Winterthur,analysis1d briefly by Switzerland)consideration illustrates the was ofdetails the installed Nyquistof the undermicro limit. orthogonal the An cutting optical machining tool microscope to measureprocess. (LEICAA variation Kistler DFCforce in cutting 450sensor, Leica, and Wizlar,thrust(9256C1 forces. Germany, TheKistler) measured was Corporation, utili frequencyzed for Winterthur, metallographic was set Switzerland as 100 observation. kHz,) was so frequencies installed The electronic under lower the than microstructures cutting 50 kHz tool could to beof measure variation in cutting and thrust forces. The measured frequency was set as 100 kHz, so theaccurately workpiece calculated and cutting in the chips power were spectrum measured density by a analysisTescan VEGA3 by consideration scanning electron of the Nyquist microscope limit. frequencies lower than 50 kHz could be accurately calculated in the power spectrum density (SEM, Tescan, Brno, Czech Republic). The compressive stress–strain test was conducted on a MTS An opticalanalysis microscope by consideration (LEICA of DFC the Nyquist 450, Leica, limit. Wizlar,An optical Germany) microscope was (LEICA utilized DFC for metallographic 450, Leica, 810observation. materialWizlar, Germanytesting The electronic system) was utili ( microstructuresMTSzed systems for metallographic corporation, of the workpiece observation. Eden Prairie, and The cutting electronicMN, chipsUSA microstructures) werewith measureda strain ofrate by of a -1 0.01Tescan the s VEGA3, workpiece and scanning the and nano cutting electron-hardness chips microscope were was measured measured (SEM, by Tescan,a Tescanusing Brno, VEGA3a NanoCzech scanning Republic). Indenter electron TheG200 microscope compressive (Keysight Technologies,Wokingham,stress–strain(SEM, Tescan, test was Brno, conducted Czech UK Republic). on a MTS). The 810 compressive material testing stress system–strain test (MTS was systems conducted corporation, on a MTS Eden Prairie,810 MN, material USA) testing with system a strain (MTS rate systems of 0.01 corporation, s−1, and the Eden nano-hardness Prairie, MN, wasUSA) measured with a strain using rate aof Nano Indenter0.01 G200 s-1, a (Keysightnd the nano Technologies,-hardness wasWokingham, measured UK). using a Nano Indenter G200 (Keysight Technologies,Wokingham, UK).
Figure 1. (a) Schematic set up of the micro orthogonal cutting and (b) the corresponding magnified imageFigureFigure; ( 1.c) (dimensiona )1. Schematic (a) Schematics of set the set up workpiece up of of the the micro micro, and orthogonal orthogonal(d) serrated cutting cutting chip formationand and (b ()b the) the incorresponding correspondingcutting. magnif magnifiedied image;image (c) dimensions; (c) dimension ofs theof the workpiece, workpiece and, and ( d(d)) serratedserrated chip chip formation formation in incutting. cutting. 3. Results and Discussion 3. Results3. Results and and Discussion Discussion 3.1. Microstructures, Stress–Strain Curves, and Nano-Hardness 3.1. Microstructures,3.1. Microstructures, Stress–Strain Stress–Strain Curves, Curves and, andNano-Hardness Nano-Hardness FigureFigure 2 2shows shows2 shows the the material the material material microstructure microstructure microstructure of the original of of the the original and original EPT-treated and and EPT EPT- Ti6Al4Vtreated-treated Ti6Al4V workpieces. Ti6Al4V workpieceThe microstructureworkpieces. Thes. Themicrostructure of microstructure the original of alloyof the the consistedoriginal original alloyalloy of equiaxial consistedconsistedβ of grainsof equiaxial equiaxial (bright β grainsβ white grains (bright particles) (bright white andwhite α particles)phase,particles) and and the andα average phase α phase, sizeand, and ofthe the average average grains size was size of aboutof thethe grains 600 nm. was was The about about EPT 600 treated600 nm n.m T alloyshe. T heEPT EPT were treated treated composed alloys alloys of wereorthorhombic werecomposed composed acicular of orthorhombic of martensite orthorhombicα acicular and acicular the martensite width martensite of the αα martensiteaandnd the the width width ranged of of the between the martensite martensite 100 andranged 500ranged nm. betweenbetween 100 and100 and500 500 nm. nm.
Figure 2. SEM microstructures of the (a) original alloy, and (b) electropulsing treatment Figure(EPT 2. 2.)SEM -treatedSEM microstructures Ti6Al4V microstructures alloy. of the of (a the) original (a) original alloy, and alloy (b) electropulsing, and (b) electropulsing treatment (EPT)-treated treatment (Ti6Al4V EPT)-treated alloy. Ti6Al4V alloy.
Micromachines 2019, 10, 197 4 of 11 MicromachinesMicromachines 20182018,, 99,, xx FORFOR PEERPEER REVIEWREVIEW 44 ofof 1111
The stress–strain curves and nano-hardness of the alloy with equiaxial β grains and martensitic The stress stress–strain–strain curves and nano nano-hardness-hardness of the alloy with equiaxial β grains and martensitic structures are shown in Figure 3. It can be clearly seen from Figure 3a that the yield stress of the EPT structures are shown in Figure 3 3.. ItIt cancan bebe clearlyclearly seenseen fromfrom FigureFigure3 3aa that that the the yield yield stress stress of of the the EPT EPT treated alloy was higher than that of the original alloy, reaching about 1400 MPa. However, the ttreatedreated alloy alloy was was higher higher than than that that of the of originalthe original alloy, alloy, reaching reaching about 1400about MPa. 1400 However, MPa. However, the ductility the ductility of the alloy after electropulsing was much lower in comparison with the original. As the ductilityof the alloy of the after alloy electropulsing after electropulsing was much wa lowers much in comparisonlower in comparison with the original.with the original. As the ultimate As the ultimate strength of the treated alloy was also smaller than the alloy without treatment, the stress ultimatestrength ofstrength the treated of the alloy treated was alloy also smallerwas also than smaller the alloy than without the alloy treatment, without thetreatment stress required, the stress to required to fracture should be less for the treated alloy. The nano-hardness of the workpiece is requiredfracture should to fracture be less should for the be treated less alloy.for the The treated nano-hardness alloy. The of nano the workpiece-hardness is of shown the workpiece in Figure3 b. is shown in Figure 3b. The average nano-hardness of the treated alloy increased by 26.3% compared shownThe average in Figure nano-hardness 3b. The average of the nano treated-hardness alloy increased of the treated by 26.3% alloy compared increased with by the26.3% original compared alloy with the original alloy with an average nano-hardness of about 4.384 GPa. The increase in with an the average original nano-hardness alloy with ofan about average 4.384 nano GPa.-hardness The increase of aboutin nano-hardness 4.384 GPa. was The primarily increase due in nano-hardness was primarily due to the increase of the yield stress, which required high stress to nanoto the-hardness increase ofwa thes primarily yield stress, due which to the required increase high of the stress yield to overcomestress, which plastic require deformation.d high stress to overcomeovercome plasticplastic deformation.deformation. (a) 88 (b) (a) 00 V,V, 00 HzHz (b) 22502250 66 50 V, 400 Hz 15001500 50 V, 400 Hz
4
4
Stress σ/MPa 0 V, 0 Hz Stress σ/MPa 750750 0 V, 0 Hz 2 50 V, 400 Hz Nano-hardness/GPa 2 50 V, 400 Hz Nano-hardness/GPa 00 00 0 5 10 0 2 4 6 8 10 0 5 10 0 Distance2 from4 workpiece6 edge/mm8 10 StrainStrain ε/%ε/% Distance from workpiece edge/mm Figure 3. (a) Stress–strain curves, and (b) nano-hardness of the alloys without EPT treatment and with FigureFigure 33.. ((aa)) SStresstress––strainstrain curvescurves,, andand ((bb)) nanonano--hardnesshardness ofof thethe alloysalloys withoutwithout EPTEPT treatmenttreatment andand EPT of 50 V, 400 Hz. withwith EPTEPT ofof 5050 V,V, 400400 Hz.Hz. 3.2. Cutting Chips 3.2.3.2. CuttingCutting CChipships Figure4 schematically illustrates the serrated chip formation process in the micro orthogonal Figure 4 schematically illustrates the serrated chip formation process in the micro orthogonal machining.Figure Chip4 schematically thickness t 0illustrates, distance the of chip serrated segment chipd sformation, serrated chipprocess angle in θthe, and micro segment orthogonal length machining. Chip thickness t0, distance of chip segment ds, serrated chip angle θ, and segment length machining.Ls are the main Chip features thickness of onet0, distance continuous of chip serrated segment chip. ds, serrated chip angle θ, and segment length s LLs areare thethe mainmain featuresfeatures ofof oneone continuouscontinuous serratedserrated chip.chip.
t0 t0 CuttingCutting directiondirection
ds ds Diamond θθ Diamond UncutUncut surfacesurface tooltool tc φ tc φ CutCut surfacesurface
WorkpieceWorkpiece
Figure 4. A schematic diagram of the serrated chip formation process. FigureFigure 4.4. AA schematicschematic diagramdiagram ofof thethe serratedserrated chipchip formationformation process.process. Figure5 compares the chip morphologies of the alloys after micro orthogonal cutting. Continuous Figure 5 compares the chip morphologies of the alloys after micro orthogonal cutting. serratedFigure chips 5 consisting compares of the a large chip number morphologies of consecutive of the chip alloys segments after were micro formed orthogonal at an extremely cutting. Continuous serrated chips consisting of a large number of consecutive chip segments were formed Continuouslow cutting speedserrated (60 mm/min),chips consisting which of is differenta large number from a previousof consecutive study that chip proposed segments that were continuous formed at an extremely low cutting speed (60 mm/min), which is different from a previous study that atserrated an extremely chips could low be cutting observed speed at cutting (60 mm/ speedsmin), higher which than is different 75 m/min from in the a traditionalprevious study machining that proposed that continuous serrated chips could be observed at cutting speeds higher than 75 m/min proposedof Ti6Al4V that alloys continuous [16]. This serrated may be chips because could the cuttingbe observed depth at in cutting ultraprecision speeds higher micro cuttingthan 75 is m much/min in the traditional machining of Ti6Al4V alloys [16]. This may be because the cutting depth in insmaller the traditional in comparison machining to traditional of Ti6Al4V machining, alloys which [16] usually. This may has a be cutting because depth the of more cutting than depth several in ultraprecision micro cutting is much smaller in comparison to traditional machining, which usually ultraprecisionhundred micrometers. micro cutting is much smaller in comparison to traditional machining, which usually hashas aa cuttingcutting depthdepth ofof moremore thanthan severalseveral hundredhundred micrometers.micrometers.
Micromachines 2019, 10, 197 5 of 11 Micromachines 2018, 9, x FOR PEER REVIEW 5 of 11
Shear angle angle is is the the angle angle between between the the shear shear direction direction in the in primary the primary deformation deformation zone and zone cutting and cuttinspeedg direction, speed direction, which iswhich an important is an important parameter parameter in evaluating in evaluating material material removal removal and cuttingand cutting chip chipthickness thickness [12,29 [12,29]. In]. Figure In Figure5a,b, 5a the,b, the measured measured shear shear angles angles were were almost almost equal equal for for the the alloy alloy with with equiaxial and and martensitic martensitic microstructures, microstructures, reaching reaching values values of 47.1 of ◦ 47.1and° 47.5and◦ , 47.5 respectively.°, respectively. This means This meansthat microstructures that microstructures had slight had effectsslight effects on the on shear the shear angle angle in micro in micro orthogonal orthogonal cutting, cutting, which which was wasalso also proposed proposed by Sun by etSun al. et [16 al.]. [16] However,. However, the number the number of chip of segmentschip segment pers unit per lengthunit length varied varie withd withthe microstructure. the microstructure. The The continuous continuous serrated serrated chips chip fors for the the EPT-treated EPT-treated alloy alloy with with thethe martensitic microstructure had more segments per unit length in comparison to the alloy without treatment, as compared in Figure 55cc,d.,d. Furthermore,Furthermore, thethe lengthlength ofof oneone chipchip segmentsegment ofof thethe martensiticmartensitic alloyalloy waswas shorter than the original and the serrated chip angle also variedvaried between microstructures.microstructures. As microcracks microcracks throughout throughout two two neighboring neighboring chip chip segments segments were were obviously obviously found found for both for alloys, both alloys,the serrated the serrated cutting chips cutting were chips supposedly were supposedly a result of a periodical result of crackperiodical initiation crack and initiation propagation and propagationbetween two between adjacent two segments adjacent during segments the micro during orthogonal the micro machining orthogonal of machining Ti6Al4V alloys. of Ti6Al4V Chip alloys.formation Chip was formation affected was by the affected cutting by conditions, the cutting cutting conditions, tools, andcutting materials tools, [and30– 32materials]. As machining [30–32]. Asconditions machining and conditions cutting tools and were cutting all the tools same were during all the micro same orthogonal during micro cutting, orthogonal the different cutting, serrated the differentchip features serrated can be chip attributed features to can the be various attributed material to the microstructures. various material The microstructures. ductility of the alloy The withductility acicular of the martensite alloy with was acicularmuch lower martensite than the was alloy much with equiaxial lower than microstructures, the alloy with according equiaxial to microstructuresthe stress–strain, curveaccording shown to the in Figurestress–3strain. Therefore, curve theshown martensitic in Figure alloy 3. Therefore, showed a the more martensitic “brittle” alloyfeature show thaned the a equiaxialmore “brittle alloy.” Hence,feature fracturethan the failure equiaxial is prone alloy. to occurHence, for fracture an alloy failure with a is martensitic prone to occurmicrostructure, for an alloy which with contributes a martensitic to more microstructure, chip segments which per contributes unit length. to more chip segments per unit length.
FigureFigure 5. SerratedSerrated chip chip morphologies morphologies for ( a,,c)) original alloys and ( b,,d) EPT EPT-treated-treated alloys. 3.3. Cutting and Thrust Forces 3.3. Cutting and Thrust Forces Figure6 compares the cutting and thrust forces variation for the two types of alloy with respect to Figure 6 compares the cutting and thrust forces variation for the two types of alloy with respect microstructures. The average cutting and thrust forces were about 8.41 N and 4.53 N, respectively, for all to microstructures. The average cutting and thrust forces were about 8.41 N and 4.53 N, respectively, the equiaxial two-phase alloys, which was consistently lower than in those with martensitic structure. for all the equiaxial two-phase alloys, which was consistently lower than in those with martensitic As the yield stress of plastic deformation of the alloy with martensite was higher than the original, the structure. As the yield stress of plastic deformation of the alloy with martensite was higher than the cutting force required to overcome flow stress should also be larger for the martensitic alloy. original, the cutting force required to overcome flow stress should also be larger for the martensitic Furthermore, the evolution of cutting and thrust forces of the equiaxial alloy showed a stable alloy. trend during the whole machining process, while obvious force fluctuations could be clearly seen Furthermore, the evolution of cutting and thrust forces of the equiaxial alloy showed a stable in the cutting and thrust force curves. As the martensitic distribution of the three workpieces was trend during the whole machining process, while obvious force fluctuations could be clearly seen in different, the forces also showed various variation trends. Figure6a,c illustrate the cutting and thrust the cutting and thrust force curves. As the martensitic distribution of the three workpieces was force of one workpiece with microstructures shown in Figure7. Obvious force fluctuations were different, the forces also showed various variation trends. Figure 6a,c illustrate the cutting and thrust observed in the curves, especially in the region marked by the black dash rectangle in Figure6a, as the force of one workpiece with microstructures shown in Figure 7. Obvious force fluctuations were average cutting force showed a sudden increase from 9.18 to 10.45 N. A similar thrust force, increasing observed in the curves, especially in the region marked by the black dash rectangle in Figure 6a, as the average cutting force showed a sudden increase from 9.18 to 10.45 N. A similar thrust force,
Micromachines 2019, 10, 197 6 of 11
Micromachines 2018, 9, x FOR PEER REVIEW 6 of 11 from 5.92 to 6.23 N, was also found during machining, however the increase was not obvious when increasingcompared with from the 5.92 cutting to 6.23 force. N, was also found during machining, however the increase was not obvious when compared with the cutting force.
Figure 6. 6. (a(a)) Cutting Cutting force force variation, variation and, and (b) corresponding(b) corresponding magnified magnif image;ied (imagec) thrust; (c force) thrust variation, force variationand (d) corresponding, and (d) corresponding magnified magnif image.ied image. Micromachines 2018, 9, x FOR PEER REVIEW 7 of 11 The cutting and thrust force variation was primarily due to the effects of different martensitic orientations, as shown Figure 7. The microstructure of the region where the cutting and thrust force experienced a sudden increase is shown in Figure 7a. Initial β grain boundary was observed between two orientations of martensite and the martensitic density of the left region was slightly more intensive than that of the right region. Figure 7b compares the two indentation marks of the Vickers hardness and shows that measured hardness in the left region was slightly higher than the right region, which indicates that the required stress for plastic deformation was larger than the right region. The martensite density in left region was more intensive than that in the right region, which increased the number of phase boundaries. Hence, the required flow stress for plastic deformation should also be higher in the left region, which agrees with previous studies showing that thin and small martensites usually give rise to high yield and ultimate strength [3,33]. In addition, the primary martensitic orientation of the two sites was totally different at the two sites, which is schematically shown by different colors in Figure 7c. The equiaxial Ti6Al4V alloy was composed of hexagonal closest packing (HCP) α phase and body centered cubic (BCC) β phase at room temperature. However, the crystallographic structure of the Ti6Al4V alloy with acicular martensite mainly consisted of HCP α phase, and only 0.4% of BCC β phase was obtained based on the previousFigure 7. study(a()a )M Microstructuresicrostructures [34]. Hence, of plastic of the the martensitic martensitic deformation Ti6Al4V Ti6Al4V of alloy the alloy alloy at the at was theregion regionsignificantly with with force force increase affected increase,, (b by) the martenscomparison(b)i comparisontic α phase of indentation instead of indentation of markers β phase. markers and Vickers and Vickershardness, hardness, and (c) a andsketch (c) of a martensitic sketch of martensiticorientation change.orientation change.
ForThe deforming cutting and HCP thrust titanium, force variation slipping wasmodes primarily mainly dueinclude to thed prismatic effects of slip different of {10 martensitic1 0} plane, basalorientations, slip of { as0001 shown} plane, Figure and7. Thepyramidal microstructure slip of {10 of1 the1} regionplane. The where slipping the cutting direction and thrustof the forcethree systems was all in the direction of < 112 0 >. The first-order and second-order pyramidal
Figure 8. Schematics of crystalline structure of hexagonal closest packing (HCP) titanium: (a) {101 0} and {0001} planes; (b) {101 1} and {112 2} planes.
The prismatic slip of {101 0} < 112 0 > was the most prone to be activated in comparison to other systems during deformation HCP titanium due to the relatively low critical resolved shear stress (CRSS). However, it was not the only slip model because five independent slip systems are required to accommodate the plastic deformation, therefore the basal slip, pyramidal slip, pyramidal
Micromachines 2018, 9, x FOR PEER REVIEW 7 of 11
Micromachines 2019, 10, 197 7 of 11 experienced a sudden increase is shown in Figure7a. Initial β grain boundary was observed between two orientations of martensite and the martensitic density of the left region was slightly more intensive than that of the right region. Figure7b compares the two indentation marks of the Vickers hardness and shows that measured hardness in the left region was slightly higher than the right region, which indicates that the required stress for plastic deformation was larger than the right region. The martensite density in left region was more intensive than that in the right region, which increased the number of phase boundaries. Hence, the required flow stress for plastic deformation should also be higher in the left region, which agrees with previous studies showing that thin and small martensites usually give rise to high yield and ultimate strength [3,33]. In addition, the primary martensitic orientation of the two sites was totally different at the two sites, which is schematically shown by different colors in Figure7c. The equiaxial Ti6Al4V alloy was composed of hexagonal closest packing (HCP) α phase and body centered cubic (BCC) β phase at room temperature. However, the crystallographic structure of the Ti6Al4V alloy with acicular martensite mainlyFigure consisted 7. (a) M oficrostructures HCP α phase, of andthe martensitic only 0.4% ofTi6Al4V BCC β alloyphase at the was region obtained with based force increase on the previous, (b) studycomparison [34]. Hence, of plasticindentation deformation markers and of the Vickers alloy hardness, was significantly and (c) a affectedsketch of bymartensitic the martensitic orientationα phase change. instead of β phase. For deforming HCP titanium, slipping modes mainly included prismatic slip of 1010 plane, For deforming HCP titanium, slipping modes mainly included prismatic slip of {101 0} plane, basal slip of {0001} plane, and pyramidal slip of 1011 plane. The slipping direction of the basal slip of {0001} plane, and pyramidal slip of {101 1} plane. The slipping direction of the three three systems was all in the direction of < 1120 >. The first-order and second-order pyramidal systems was all in the direction of < 112 0 >. The first-order and second-order pyramidal
Figure 8. Schematics of crystalline structure of hexagonal closest packing (HCP) titanium: (a) 1010 andFigure{0001 8. Schematics} planes; (b of) crystalline1011 and structure1122 planes. of hexagonal closest packing (HCP) titanium: (a) {101 0} and {0001} planes; (b) {101 1} and {112 2} planes. The prismatic slip of 1010 < 1120 > was the most prone to be activated in comparison to otherThe systems prismatic during deformation slip of {101 HCP0} < titanium1120 > duewas tothe the most relatively prone lowto be critical activated resolved in comparison shear stress to (CRSS).other systems However, during it was deformation not the only HCP slip modeltitanium because due to five the independent relatively low slip critical systems resolved are required shear tostress accommodate (CRSS). Howe the plasticver, it deformation,was not the only therefore slip model the basal because slip, five pyramidal independent slip slip, systems pyramidal are
3.4. Power Spectral Density (PSD) of Cutting Forces PSD analysis via transforming time domain to frequency domain by the fast Fourier transform (FFT)Micromachines was adopted2019, 10, 197 to investigate the cutting force signals of the Ti6Al4V alloys with two types8 of of 11 microstructure, as shown in Figure 9. Three high frequencies at about 16.05, 25.75, and 29.10 kHz could be clearly observed in the air cutting conditions when the diamond tool did not make contact 3.4. Power Spectral Density (PSD) of Cutting Forces with either the original or EPT-treated alloy, and the corresponding PSD values were fairly small, as shownPSD in analysisFigure 9a via,b. transformingHence, the three time high domain frequencies to frequency should domain be intrinsic by the background fast Fourier frequencies transform (FFT)generate wasd by adopted the machining to investigate system. the cutting force signals of the Ti6Al4V alloys with two types of microstructure,During the as machining shown in Figureprocess9. when Three the high diamond frequencies tool at was about consistently 16.05, 25.75, cutting and 29.10 the workpiece, kHz could thebe clearlycharacteristic observed frequency in the air of cutting the cutting conditions force signals when thewas diamond mainly in tool a lo didw frequency not make contactscale with with a largeeither PSD theoriginal value in or comparison EPT-treated to alloy, the andthree the high corresponding intrinsic background PSD values frequencies were fairly small,for both as shownof the workpiecein Figure9a,b.s, as Hence, shown the inthree Figure high 9c– frequenciesf. The characteristic should be frequency intrinsic background of cutting frequenciesforce variation generated of the equiaxialby the machining two phase system. alloy ranged from 100 to 200 Hz, while it ranged from 200 to 400 Hz for the workpiece with martensites.
Figure 9. PPowerower spectral spectral density density (PSD) (PSD) analyses analyses for for air air cutting cutting of ofthe the alloy alloy with with (a) (equiaxeda) equiaxed and and (b) (martensiticb) martensitic microstructure; microstructure; (c–f ()c – PSDf) PSD analyses analyses under under different different frequencies frequencies for for the the equiaxed equiaxed and martensitic microstructures, respectively.
During the machining process when the diamond tool was consistently cutting the workpiece, the characteristic frequency of the cutting force signals was mainly in a low frequency scale with a large PSD value in comparison to the three high intrinsic background frequencies for both of the workpieces, as shown in Figure9c–f. The characteristic frequency of cutting force variation of the equiaxial two phase alloy ranged from 100 to 200 Hz, while it ranged from 200 to 400 Hz for the workpiece with martensites. Micromachines 2019, 10, 197 9 of 11
The cutting force variation was estimated from the serrated chip feature in machining Ti6Al4V alloys according to Sun et al. [16]. The frequency f of cutting force variation could be calculated by the segment length and cutting speed (f = v/ Ls). Since segment lengths measured from Figure5 ranged from 4.34 to 8.26 µm and 2.17 to 3.91 µm, the calculated vibration frequencies were in ranges of 121.07–230.41 Hz and 255.75–460.83 Hz for the original and EPT-treated alloy, respectively.
4. Conclusions The serrated chip formation of Ti6Al4V alloys was experimentally investigated in this study, which provides useful instructions for analyzing chip features with respect to equiaxed and martensitic microstructures during the micro machining of Ti6Al4V alloys. Conclusions could be drawn as follows: (1) The needlelike martensitic microstructure of the Ti6Al4V alloy with widths ranging from 100 to 500 nm could be obtained by electropulsing treatment for 2 min followed by water quenching. (2) The shear angle was slightly affected by the microstructures of the Ti6Al4V alloy. However, the chip features varied significantly between alloy microstructures. The number of chip segments per unit length of the martensitic alloy were more than the original alloy due to poor ductility, however the segment length of the former was shorter than the latter. (3) The average cutting and thrust forces remained stable at about 8.41 and 4.53 N for the equiaxial Ti6Al4V alloy, which were both lower than those with martensitic microstructures that demonstrated fluctuating cutting force variation. Besides, the sudden increase of cutting force from 9.18 to 10.45 N for the martensitic Ti6Al4V alloy was due to the change of martensitic orientations. (4) The characteristic frequency of cutting force signals for the equiaxed alloy ranged from 100 to 200 Hz, while it ranged from 200 to 400 Hz for the alloy with martensites, according to the power spectral density (PSD) analyses. The varying frequencies were supposedly due to the different serrated chip features in the machining process.
Author Contributions: Conceptualization, S.T.; Data curation, Z.J.Z.; Funding acquisition, S.T.; Investigation, Z.J.Z.; Methodology, Z.J.Z.; Resources, S.T.; Software, Z.X.Z.; Writing—original draft, Z.J.Z.; Writing—review & editing, S.T. Funding: The work described in this paper was jointly supported by the Research Committee (Project No. G-YBLE) and Partner State Key Laboratory of Ultra-precision Machining Technology (Project No. RUWB) of the Hong Kong Polytechnic University. Conflicts of Interest: The authors declare no conflict of interest.
References
1. Sha, W.; Malinov, S. Titanium Alloys: Modelling of Microstructure, Properties and Applications; Elsevier: Amsterdam, The Netherlands, 2009. 2. Sachdev, A.K.; Kulkarni, K.; Fang, Z.Z.; Yang, R.; Girshov, V. Titanium for automotive applications: Challenges and opportunities in materials and processing. JOM 2012, 64, 553–565. [CrossRef] 3. Lütjering, G.; Williams, J.C. Titanium; Springer Science & Business Media: Berlin, Germany, 2007. 4. Matsumoto, H.; Yoneda, H.; Sato, K.; Kurosu, S.; Maire, E.; Fabregue, D.; Konno, T.J.; Chiba, A. Room-temperature ductility of Ti–6Al–4V alloy with α0 martensite microstructure. Mater. Sci. Eng. A 2011, 528, 1512–1520. [CrossRef] 5. Zareena, A.; Veldhuis, S. Tool wear mechanisms and tool life enhancement in ultra-precision machining of titanium. J. Mater. Process. Technol. 2012, 212, 560–570. [CrossRef] 6. Lu, M.; Zhou, J.; Lin, J.; Gu, Y.; Han, J.; Zhao, D. Study on Ti-6Al-4V alloy machining applying the non-resonant three-dimensional elliptical vibration cutting. Micromachines 2017, 8, 306. [CrossRef][PubMed] 7. Balaji, M.; Rao, K.V.; Rao, N.M.; Murthy, B. Optimization of drilling parameters for drilling of TI-6Al-4V based on surface roughness, flank wear and drill vibration. Measurement 2018, 114, 332–339. [CrossRef] 8. Ng, C.K.; Melkote, S.N.; Rahman, M.; Kumar, A.S. Experimental study of micro-and nano-scale cutting of aluminum 7075-T6. Int. J. Mach. Tools Manuf. 2006, 46, 929–936. [CrossRef] Micromachines 2019, 10, 197 10 of 11
9. Wang, Q.; Bai, Q.; Chen, J.; Sun, Y.; Guo, Y.; Liang, Y. Subsurface defects structural evolution in nano-cutting of single crystal copper. Appl. Surf. Sci. 2015, 344, 38–46. [CrossRef] 10. Rahman, M.A.; Rahman, M.; Kumar, A.S. Modelling of flow stress by correlating the material grain size and chip thickness in ultra-precision machining. Int. J. Mach. Tools Manuf. 2017, 123, 57–75. [CrossRef] 11. Shokrani, A.; Dhokia, V.; Newman, S.T. Environmentally conscious machining of difficult-to-machine materials with regard to cutting fluids. Int. J. Mach. Tools Manuf. 2012, 57, 83–101. [CrossRef] 12. Sutter, G.; List, G. Very high speed cutting of Ti–6Al–4V titanium alloy–change in morphology and mechanism of chip formation. Int. J. Mach. Tools Manuf. 2013, 66, 37–43. [CrossRef] 13. Childs, T.H.; Arrazola, P.-J.; Aristimuno, P.; Garay, A.; Sacristan, I. Ti6Al4V metal cutting chip formation experiments and modelling over a wide range of cutting speeds. J. Mater. Process. Technol. 2018, 255, 898–913. [CrossRef] 14. Komanduri, R.; von Turkovich, B. New observations on the mechanism of chip formation when machining titanium alloys. Wear 1981, 69, 179–188. [CrossRef] 15. Komanduri, R.; Hou, Z.-B. On thermoplastic shear instability in the machining of a titanium alloy (Ti-6Al-4V). Metall. Mater. Trans. A 2002, 33, 2995. [CrossRef] 16. Sun, S.; Brandt, M.; Dargusch, M. Characteristics of cutting forces and chip formation in machining of titanium alloys. Int. J. Mach. Tools Manuf. 2009, 49, 561–568. [CrossRef] 17. Vyas, A.; Shaw, M. Mechanics of saw-tooth chip formation in metal cutting. J. Manuf. Sci. Eng. 1999, 121, 163–172. [CrossRef] 18. Hua, J.; Shivpuri, R. Prediction of chip morphology and segmentation during the machining of titanium alloys. J. Mater. Process. Technol. 2004, 150, 124–133. [CrossRef] 19. Wang, B.; Liu, Z. Investigations on the chip formation mechanism and shear localization sensitivity of high-speed machining Ti6Al4V. Int. J. Adv. Manuf. Technol. 2014, 75, 1065–1076. [CrossRef] 20. Wu, H.; To, S. Serrated chip formation and their adiabatic analysis by using the constitutive model of titanium alloy in high speed cutting. J. Alloy. Compd. 2015, 629, 368–373. [CrossRef] 21. Sima, M.; Özel, T. Modified material constitutive models for serrated chip formation simulations and experimental validation in machining of titanium alloy Ti–6Al–4V. Int. J. Mach. Tools Manuf. 2010, 50, 943–960. [CrossRef] 22. Che, J.; Zhou, T.; Liang, Z.; Wu, J.; Wang, X. Serrated chip formation mechanism analysis using a modified model based on the material defect theory in machining Ti-6Al-4 V alloy. Int. J. Adv. Manuf. Technol. 2018, 96, 3575–3584. [CrossRef] 23. Calamaz, M.; Coupard, D.; Girot, F. A new material model for 2D numerical simulation of serrated chip formation when machining titanium alloy Ti–6Al–4V. Int. J. Mach. Tools Manuf. 2008, 48, 275–288. [CrossRef] 24. Liu, G.; Shah, S.; Özel, T. Material Ductile Failure Based Finite Element Simulations of Chip Serration in Orthogonal Cutting of Titanium Alloy Ti-6Al-4V. J. Manuf. Sci. Eng. 2019, 141, 041017. [CrossRef] 25. Barge, M.; Hamdi, H.; Rech, J.; Bergheau, J.-M. Numerical modelling of orthogonal cutting: Influence of numerical parameters. J. Mater. Process. Technol. 2005, 164, 1148–1153. [CrossRef] 26. Nouari, M.; Makich, H. On the physics of machining titanium alloys: Interactions between cutting parameters, microstructure and tool wear. Metals 2014, 4, 335–358. [CrossRef] 27. Nouari, M.; Makich, H. Experimental investigation on the effect of the material microstructure on tool wear when machining hard titanium alloys: Ti–6Al–4V and Ti-555. Int. J. Refract. Met. Hard Mater. 2013, 41, 259–269. [CrossRef] 28. Ahmadi, M.; Karpat, Y.; Acar, O.; Kalay, Y.E. Microstructure effects on process outputs in micro scale milling of heat treated Ti6Al4V titanium alloys. J. Mater. Process. Technol. 2018, 252, 333–347. [CrossRef] 29. Berezvai, S.; Molnar, T.G.; Bachrathy, D.; Stepan, G. Experimental investigation of the shear angle variation during orthogonal cutting. Mater. Today Proc. 2018, 5, 26495–26500. [CrossRef] 30. Wan, L.; Wang, D.; Gao, Y. The investigation of mechanism of serrated chip formation under different cutting speeds. Int. J. Adv. Manuf. Technol. 2016, 82, 951–959. [CrossRef] 31. Dargusch, M.S.; Sun, S.; Kim, J.W.; Li, T.; Trimby, P.; Cairney, J. Effect of tool wear evolution on chip formation during dry machining of Ti-6Al-4V alloy. Int. J. Mach. Tools Manuf. 2018, 126, 13–17. [CrossRef] 32. Liu, R.; Melkote, S.; Pucha, R.; Morehouse, J.; Man, X.; Marusich, T. An enhanced constitutive material model for machining of Ti–6Al–4V alloy. J. Mater. Process. Technol. 2013, 213, 2238–2246. [CrossRef] Micromachines 2019, 10, 197 11 of 11
33. Zhao, Z.; To, S.; Sun, Z.; Ji, R.; Yu, K.M. Microstructural effects of Ti6Al4V alloys modified by electropulsing treatment on ultraprecision diamond turning. J. Manuf. Process. 2019, 39, 58–68. [CrossRef] 34. Zhao, Z.; To, S. An investigation of resolved shear stress on activation of slip systems during ultraprecision rotary cutting of local anisotropic Ti-6Al-4V alloy: Models and experiments. Int. J. Mach. Tools Manuf. 2018, 134, 69–78. [CrossRef]
© 2019 by the authors. Licensee MDPI, Basel, Switzerland. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC BY) license (http://creativecommons.org/licenses/by/4.0/).