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ear mechanisms and tool life management of WC–Co drills during dry highpeed drilling of woven carbon fibre composites
anjay Rawata, Helmi Attiaa,b,∗
Mechanical Engineering Department, McGill University, Montreal, CanadaAerospace Manufacturing Technology Centre, Institute for Aerospace Research, National Research Council of Canada,145 Ave Decelles, Campus of University of Montreal, Montreal, Quebec, Canada H3T 2B2
r t i c l e i n f o
rticle history:eceived 24 September 2008eceived in revised form 25 January 2009ccepted 29 January 2009
Woven carbon fibre composites are extensively used in aerospace, automotive and civil applications.Optimization of high speed drilling of composites is an issue of great economical impact. Tool wear andpart quality are central to the definition of the objective function and constraints of the optimizationscheme. This paper presents an experimental investigation of the wear mechanisms of tungsten carbide(WC) drills during dry high speed drilling of quasi-isotropic woven graphite fibre epoxy composites. Toolwear was evaluated at spindle speeds of up to 15,000 rpm using a standard two flute drill. The paperexamines the nonlinear behavior of this tribo-system and the interdependence of the wear process andcutting forces in relation to surface damage of the system components. It was found that chipping andabrasion were the main mechanisms controlling the deterioration of WC drill. The two friction regimes,the lightly and heavily loaded, were found to dictate the increase in forces, delamination of compositeand surface roughness. The aggressive rubbing by fractured graphite fibres and WC grains against the soft
epoxy matrix caused high temperature rise and consequently enhanced flank wear. During the primaryand secondary wear stages, wear on the flank face of main cutting edges was found to be dominant,while adhesion of carbon was found to occur along with abrasion in the tertiary zone. Tool life resultsrevealed the increase in the delamination and surface roughness with transition from the primary totertiary wear regime. The correlation between tool wear, delamination damage and surface roughness
was c
was established. Finally itthe cutting forces.
. Introduction
Woven graphite fibre epoxy composites are being extensivelysed in aerospace, automotive and civil applications, due to theirnique mechanical properties. In comparison to unidirectionalomposites, this class of materials possess higher strength-to-eight ratio, higher fracture toughness and excellent corrosion
esistance properties [1–5]. Ever since its inception in 1970s, theomposite material design and manufacturing technologies haveatured to a level that the Boeing company is using compos-
te material for 50% of the primary structure in its 787 program.achining of composites, particularly drilling, is extensively used
or producing riveted and bolted joints during assembly operations.ny defect arising in the structure during machining has a signifi-
∗ Corresponding author at: Aerospace Manufacturing Technology Centre, Instituteor Aerospace Research, National Research Council of Canada, 5145 Ave Decelles,ampus of University of Montreal, Montreal, Quebec, Canada H3T 2B2.el.: +1 514 283 9002; fax: +1 514 283 9662.
cant economical impact [6]. From this perspective, tool wear playsa critical role in selecting the optimum cutting conditions for min-imum production cost or maximum productivity. Tool wear alsoresults in the damage of the laminate. This process is further com-pounded by the characteristic attributes of woven composites, e.g.,the non-homogeneous structure, anisotropy and high abrasivenessof fibres [7].
For high speed/high performance machining of composites, theprocess has to be optimized, monitored and controlled to meet thedesign tolerances and achieve defect-free components for betterin-service fatigue life. The delamination model developed by Ho-Cheng and Dharan [8] for unidirectional composites, which wasbased on linear elastic fracture mechanics and drilling mechanics[9], established the threshold force Fth required to cause delami-nation at the hole entry and exit. This force depends on materialproperties and the uncut laminate depth under the tool. Jain et al.
extended Ho-Cheng et al.’s model and developed a variable feedrate strategy to avoid delamination [10]. Other models were thendeveloped to incorporate the bi-directional behavior of multidirec-tional laminates. Most of these studies highlighted the importanceof understanding and minimizing tool wear in order to maintain the
S. Rawat, H. Attia / Wear 267 (2009) 1022–1030 1023
wear i
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Fig. 1. System approach for assessment of tool
esired hole quality. However, there is still limited knowledge andnderstanding of performance of both the composite material andool for woven graphite epoxy composites, in spite of its extensivese in the industry.
Over the years, different tool materials like high-speed steelHSS), carbides, coated HSS, coated carbides and diamond coatedarbides have been tested to understand the wear mechanisms foranaging and enhancing tool life. Most of the research work car-
ied out to assess the performance of these materials has identifiedbrasive wear as one of the major wear mechanism [11–14]. In ref.11], Malhotra reported the occurrence of both chisel edge and flankear during drilling of carbon epoxy composite laminate with HSS
nd Carbide drills at a relatively low spindle speed of 1250 rpm andeed rate of 60 �m/rev. Carbide drills were found to perform bet-er than HSS tools. Lin and Chen [12] compared the performancef standard twist drill and multifaceted drills in ultra high speedrilling (up to 38,000 rpm) of CFRP composites and found aggres-ive tool wear to be a major problem at such speeds. They observedhat the increase in tool wear is associated with the increase inhrust and tangential forces. Inoue et al. [13] investigated the influ-nce of tool wear on the internal damage in small diameter drillingf glass fibre epoxy composites. They found higher flank wear toccur at lower feed rates and higher cutting speed. Velayudham etl. [14] evaluated the performance of carbide drills in the drilling ofigh volume fraction glass fibre polymeric composites. They iden-ified wear mechanisms similar to those reported in ref. [11] andound that flank wear increases steadily up to about 300 holes andhen accelerates rapidly. In ref. [15], Kim and Ramulu performed anvaluation of HSS and carbide tools in drilling graphite fibre bismal-mide (Gr/Bi) composite and titanium alloy stacks. Carbides, havingigher hot hardness, were found to outperform HSS drills. The flankear was found to increase both with spindle speed and feed rate.
amkumar et al. [16] evaluated the performance of coated (TiN andiC) and uncoated HSS tools during drilling of glass fibre epoxy atpindle speed of 1700 rpm and feed rate of 250 �m/rev. It was con-luded that coated drills wear faster beyond a certain number ofoles. This was mainly attributed to the peeling and chipping off of
ig. 2. SEM images of the damages observed at high speed: (a) matrix burning (n = 1500n = 5000, f = 20 �m/rev).
n drilling of woven graphite epoxy composites.
the coating from the base material. In ref. [17], Murphy et al. testedthe performance of coated and uncoated tungsten carbide drills indrilling carbon fibre epoxy composites at spindle speed of 3000 rpmand feed rate of 50 �m/rev. The effect of the increase in the flankwear on the cutting force, torque and the damage to the carbonfibre reinforced polymer (CFRP) composite material was studied. Acomparison of the studies reported in [16] and [17] clearly revealedthat uncoated carbides perform much better than HSS tools for bothglass and carbon fibre epoxy composites. Most of these investiga-tions focused on low speed drilling, except [12], where ultra highspeeds were investigated.
In the present work, a system-based approach for high speeddrilling (10,000–15,000 rpm) of woven graphite epoxy compositesis used to investigate various tool wear mechanisms of carbidetools and their affect on the hole quality. The following sectionspresent the experimental setup, the system approach to tool wearassessment, and the experimental results that reveal the interactionbetween tool wear, cutting forces, delamination and hole qualityattributes.
2. System approach to tool wear during the machining ofCFRP
A system approach to tool wear process reveals the non-linear nature of the tool–workpiece tribological interaction, asschematically shown in Fig. 1. The system consists of three ele-ments, namely, the independent input variables (materials andcutting conditions), the dependent intermediate variables (cut-ting forces, temperature and friction) and finally the outputs(desirable quality attributes of the hole). The nonlinear behaviorand closed-loop interactions of this tribo-system are demon-strated by the fact that the tool wear process is both affecting
and being affected by the intermediate process variables. Whilecutting forces and temperature control the mechanism and thekinetics of the wear process, the latter in turn alters the toolgeometry, deteriorates the cutting capability of the tool and thusultimately dictates the hole quality of the laminate. The results of
rpm, f = 20 �m/rev), (b) fibre pullout (n = 1500, f = 800 �m/rev) and (c) fibre pullout
1024 S. Rawat, H. Attia / Wear 267 (2009) 1022–1030
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ig. 3. Experimental setup: (a) the composite drilling arrangement with the fixtureEM image of the composite laminate showing the fibers along the fill and wrap dirn drilling dynamometer.
he experimental investigation presented in Section 4 will exam-ne various wear mechanisms, the wear rates, the interaction of
ear process with thrust and cutting forces, and finally the out-ut parameters; namely, delamination damage, hole circularityrror, hole diameter error and surface roughness during high speedrilling.
The tool wear mechanisms encountered in machining of com-osites are characterized by some unique features, due to thehermo-mechanical interactions of the tool–workpiece system:
(a) The thermal aspect: As observed in this study, the temperaturerise on the tool cutting edge during the machining of carbonfibre reinforced polymer (CFRP) may exceed a threshold level of300 ◦C [18,19] that causes a reduction of the fiber-matrix inter-facial shear strength [20] and ultimately the matrix burnout andthe acceleration of the fibre pullout as shown in Fig. 2. The tem-perature rise of the matrix is demonstrated by the reduction inthe thrust and cutting forces at high speeds due to the matrixsoftening.
b) The mechanical fracture aspect: The cutting process in CFRP isentirely based on fracture, as opposed to shearing phenomenonin metals [20] due to the presence of the fibres that impairuniform plastic deformation. Due to the brittleness of the ther-moset matrix, chips tend to fracture at earlier stages in theform of powder-like chips [19,21], particularly at high machin-ing speeds due to the increase in strain rate and the reducedchain sliding [22]. Also, owing to the inhomogeneous structureand the brittleness of carbon fibres, the indenting chisel edge ofthe drill causes the fracture of hard fibres inside the soft epoxymatrix [5].
These two aspects make the abrasive tool wear more dominanthrough the following two modes:
(a) Hard abrasion by fractured WC grains: Dynamic stresses aregenerated on the WC hard grains due to the impacts from thereinforcement, the broken fibers and the powder-like chips [5].These stresses result in crack initiation and propagation insidethe WC grains, which eventually cause the WC grains to fall
away partly or wholly by brittle fracture [23]. These particlescontribute to the three-body abrasion wear of the tool as theyslide over the rake and clearance faces. This process is demon-strated by the scratches and the fractured surface of the tool,shown in Fig. 4 and reported by others [5].
ounted on the force dynamometer (2) and a vacuum-assisted dust collector (3); (b); (c) the side view of the setup highlighting the fixture laminate assembly mounted
(b) Soft abrasion mode: Due to the relatively low hardness of car-bon fibre (CF) relative to tungsten carbide grains (WC), as willbe shown in Fig. 4, the WC grains cannot be abraded by the car-bon fibres pulled out of the matrix. These fibres can, however,damage the relatively soft Co binder [24–26], through a three-body abrasion process. As the binders wear deeper, the exposedarea of the WC grains increases and the fracture of these grainsby fatigue is accelerated. The removal of a small amount of thecobalt binder by the abrasive particles results in a rapid pro-cess of crack nucleation. Under the cyclic loading imposed bythe abrasive particles, these cracks propagate into the subsur-face, following a tungsten carbide grain intergranular path, untilmicroscopic spalling takes place. The formulation of this theorywas carried out in [27,28]. It is evident that this wear mecha-nism is promoted by the machining-induced matrix damage;namely, matrix burnout, fibre pullout and fibre fracture.
Another wear mechanism that operates in parallel is attributedto the anisotropy of the mechanical properties of WC, which resultsin material removal by the shearing of very thin platelets parallel toits prismatic plane. It was demonstrated in [29] that the prismaticplane hardness (H = 11 GPa) is much smaller than that of the per-pendicular plane (H = 22 GPa). This explains why the relatively softsilicon nitride (H = 18 GPa) does not wear, but the harder WC–Comaterial is transferred to it [29].
Intensive adhesion wear of the WC tools in contact with thechip and the machined surface of fibre glass reinforced polymericcomposite has also been observed under some cutting conditions[30].
3. Experimental setup
A quasi-isotropic laminate comprising of 28 plies ofwoven graphite epoxy prepreg having a final lay-up of[(0◦/−45◦/+45◦/90◦)3/0◦/90◦]s was the test material used (Fig. 3(b)).The laminate was manufactured by autoclave molding with a curetime of 60 min at 260 ◦F under 516.75 kPa autoclave pressure toproduce a final cured thickness of about 5.90 ± 0.02 mm. The fibrevolume fraction of the composite is Vf = 60%. A 2-flute, 5 mm diam-
eter drill with a 30◦ helix angle, 118◦ point angle, a fluted lengthof 44.5 mm and a total length of the drill of 76.2 mm was usedin the tests. The carbide grade of the drill used was ISO K10–K20with approximately 7% cobalt as binder. This drill is commonlyused for hole drilling in aerospace applications. The laminate was
S. Rawat, H. Attia / Wear 26
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AdFtGaFetamKladamacsSficoaf
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ig. 4. Micro-hardness of the WC single crystal, the cemented tungsten carbide cWC,he cobalt binder and the carbon fibre CF.
andwiched inside as special fixture (Fig. 3(a) and (c)) for multipleole drilling in a single setup.
The high speed drilling tests were conducted on a 5-axis Makino88� machining center having a 50 kW spindle, maximum spin-le speed of 18,000 rpm and maximum feed rate of 50 m/min.lank wear measurements were carried out in accordance withhe International Standard ISO 8688-1, using an Olympus ModelZX 12 optical microscope. The tests were discontinued when anverage flank wear criterion “Vb” of 225 �m was reached (referig. 6(b)). Although Vb is set at 300 �m by ISO 8688-1, it was low-red to 225 �m in this investigation due to the extensive damageo laminate. Machining tests were conducted under dry conditionsnd the fine chips were collected using a special vacuum arrange-ent (Fig. 3(a)). Cutting forces were recorded using a 9272-type
istler dynamometer (Fig. 3(a) and (c)). Delamination was ana-yzed using the Olympus Model GZX 12 optical microscope. It wasssessed in terms of the ratio of maximum damaged area to theesired hole diameter (refer Fig. 10(b)). The hole circularity toler-nce is defined by a pair of concentric circles that must contain theaximum and minimum radius points of a circle. Hole circularity
nd diameter error were measured using a “Mitutoyo-Mach 806”oordinate measuring machine (CMM). Surface roughness mea-urements (CLA) were done in accordance with the Internationaltandard IS0 4288:1996 using a Form Talysurf series 2 surface pro-lometer from Taylor Hobsons. The high speed drilling tests werearried out at 12,000 and 15,000 rpm and relatively low feed rate
f 100 �m/rev. The basis for selecting this low feed was based onpreliminary investigation, which was carried out to identify the
eed rate that minimizes delamination damage.Micro-hardness tests were carried out on a Struers Duramin
300 machine. The hardness of the cemented tungsten carbide
ig. 5. Wear mechanisms of the WC drill at high speed drilling of 15,000 rpm and low feedf 10 holes; (b) chipping on the rake face after drilling of 100 holes; (c) chipping on rake farilling of 50 holes.
7 (2009) 1022–1030 1025
(cWC) tool material at room temperature is HcWC = 1900 ±150 HV.At 300 ◦C, HcWC is to below 1710 HV [9], and then to one-sixth ofits RT value at 700 ◦C [31]. It should be noted that high-resolutionnano-indentation mapping was used in [32] to measure the hard-ness of the different phases of cWC. Across a single WC crystal,HsWC = 26–48 GPa, i.e., 2650–4895 HV. This exceeds significantlythe values reported in the literature using micro-hardness inden-ter: 12–24 GPa (1225–2500 HV). Such variation is attributed to thesize of the indenter and the dependence of hardness on the crystal-lographic orientation of the crystal [32,33]. It was also shown thatthe hardness of the cobalt binder varies between 410 and 1225 HV[34], as opposed to the bulk hardness of pure Co of 200 HV [34].
The overall hardness of the composite material HCFRP is 80 Bar-col, which is equivalent to 418 HV. The measured hardness Hepoxy
of the epoxy matrix is 73 HV. With fibre volume fraction of Vf = 60%,and by applying the rule of mixture, the hardness of the carbon fibreis estimated as HCF = 648 HV. Attempts made to measure the micro-hardness of the carbon fibre in the traverse, longitudinal and normaldirections produced inaccurate under-estimated values (<200 HV)due to the architecture of the composite, the relatively large sizeof the indenter and the elasticity of the matrix. Fig. 4 shows theminimum and maximum measured and estimated values of thecemented carbide cWC, the single WC crystal, the Co-binder, andthe carbon fibre.
The tool temperature was measured by inserting two K-typethermocouples into the through-coolant-holes inside the drill. Thehot junctions were welded on the flank face of the drill at 0.95 mmfrom the cutting edges. To obtain appreciable amount of tempera-ture rise, four composite plates were stacked together to achieve alarger depth of cut of 23.6 mm. At spindle speed of 15,000 rpm andfeed rate of 200 �m/rev, the rate of temperature rise was found to be350 ◦C/s. A steady state could not be reached except for the lowestspindle speed of 1500 rpm and the lowest feed rate of 20 �m/rev.
To analytically predict the steady state cutting temperature atother spindle speeds and feed rates, the tool was treated as a lumpedmass that behaves as a first order linear system. The actual tem-perature response from the experiments was used to evaluate thetime constant ‘�s’ and the steady state temperature ‘Ts’. By con-ducting FE analysis using DEFORM software [35], the temperaturedifference between the tool edge and the point of temperature mea-surement was estimated to exceed 150 ◦C. From this analysis, it
is concluded that the local steady state temperature may exceed500 ◦C at 15,000 rpm and 200 �m/rev, considering the intensivemicro-contact between the cutting edge and the reinforcement.It is established that the hardness of uncoated WC–Co tools andthe abrasive wear resistance are strongly degraded at temperatures
rate of 100 �m/rev: (a) chipping on the rake face along the chisel edge after drillingce (magnified image of (b)); (d) chipping at the corner and margin of the drill after
1026 S. Rawat, H. Attia / Wear 267 (2009) 1022–1030
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ig. 6. SEM images of wear damage observed during drilling at spindle speed of 15f WC drill cutting edge showing aggressive abrasive wear, (b) flank wear on the pbrasive wear, (c) rounding of corner of drill, and (d) adhesion of carbon on the flan
igher than about 500 ◦C, due to severe surface oxidation of thisaterial [31].
. Results and discussion
.1. Tool wear mechanisms and progression
As discussed in Section 2, the twist drill is subjected to aggressivebrasive action in the presence of hard graphite fibers embeddednside the soft epoxy matrix as well as fractured WC grains. Theuctuating chip load acting on the cutting edge leads to completelyifferent wear characteristics of the WC drills when used for drillingFRP composites, as compared to metals. Fracture (chipping), andbrasion were the main mechanisms observed in high speed drilling12,000 and 15,000 rpm) and feed rate of 100 �m/rev. At the start ofhe drilling process, chipping was observed at the following loca-ions on the drill: (i) the rake face relieved along the chisel edge
Fig. 5(a)), (ii) the rake face of the cutting edge (Fig. 5(b) and (c)) andiii) the corner of the drill and secondary cutting edges (Fig. 5(d)).he chipping at these areas occurred during the start of drilling,hen the edges were sharp and the stresses were high. Tungsten
ig. 7. Wear progression at the flank face of the carbide drill at spindle speed of2,000 and 15,000 rpm and feed rate of 100 �m/rev. (OA and OB represent the endf primary wear region and secondary wear region, respectively for spindle speed of5,000 rpm. O′A′ and O′B′ represent the end of primary wear region and secondaryear region, respectively for spindle speed of 12,000 rpm).
(
(
pm and feed rate of 100 �m/rev: (a) chisel edge wear with enlarged image of pointy cutting edge with enlarged image of flank face of WC drill cutting edge showingand the corner of the drill, which was analyzed chemically (Fig. 8).
carbide, being brittle in nature, is unable to sustain these highstresses and thus undergoes chipping. The cracks created becauseof chipping were worn away with the wear progress during contin-uous drilling.
As discussed in Section 2, tool wear by abrasion occurs mainlyon the rake and flank face of the cutting edge of the WC drill asa result of: (a) hard abrasion by fracture of the WC grains and (b)soft abrasion mode. Fig. 6 illustrates the abrasive wear observed atthe chisel edge, the flank face of the cutting edge, and the secondarycutting edge. In drilling woven composites, abrasive wear was foundto be more severe on the flank face of the two primary cutting edges.The fractured WC grains and the powdery chips formed at low feedrates and high spindle speeds rub against the flank face, resultingin a “three-body” abrasive wear mechanism. Moreover, due to itspowdery form, the mobility of the chips increases as it moves alongthe drill flutes, and as a result the abrasive action on the rake faceis not as aggressive as on the flank face. Although adhesion wear ofthe WC tools was observed in the drilling of fibre glass reinforcedpolymeric composite [30], it was not found to be as dominant asabrasive wear in the present study. However, there was evidenceof the adhesion of carbon residues on the flank face, as shown inFig. 6(d). The progression in the flank wear is not uniform and canbe divided into three distinct regions, as shown in Fig. 7:
a) Initial (or Primary) Wear Region: Wear in this region is causedby chipping or micro-cracking as shown in Fig. 5. As indicatedin Section 3, other phenomenon, namely, surface oxidation cansignificantly reduce the hardness and the abrasive wear resis-tance of WC–Co in the cutting temperature range encounteredin high speed machining. At the start of drilling, the new cuttingedges, having sharp corner radius, carry cutting forces over rela-tively small chip contact area. Consequently, the extremely highcontact pressure results in bulk subsurface flow and cause thetool–workpiece system to behaves as a heavily loaded system[5], resulting in high wear rate.
b) Steady Wear (Secondary) Region: After the initial wear or cut-ting edge rounding, the increase in the area of contact betweenthe tool and workpiece results in lower contact stresses. As thetribo-system becomes a lightly loaded system, the wear rate
is reduced and becomes nearly constant with the increase inthe sliding distance (or time). This is also accompanied withimprovement in the surface micro-roughness.
(c) Severe (Ultimate, Catastrophic or Tertiary) Wear Region: As theflank wear reaches a second critical value, the cutting force and
S. Rawat, H. Attia / Wear 267 (2009) 1022–1030 1027
Fss
Exit or push out delamination showed similar trend as entry
ig. 8. Chemical analysis of tool rake face (shown in Fig. 4 (d)) of the drill used atpindle speed of 15,000 rpm and feed rate of 100 �m/rev, using Energy Dispersivepectrometer during SEM imaging.
temperature increase rapidly. The combined effect of thermalsoftening of the workpiece and tool materials and the increasein applied contact pressure cause the tribo-system to behaveagain as a highly loaded system, exhibiting sharp increase inwear rate. These self-induced changes in the tribological systemunder consideration underline the very nature of this nonlineardynamic system. The main findings about the flank wear regionsat high drilling speeds of 12,000 and 15,000 rpm are:i. The initial primary tool wear region at spindle speed
n = 15,000 rpm lasted up to 30 s (up to OA in Fig. 7), whichis equivalent to about 100 holes drilled at 100 �m/rev or atotal cutting length of 0.6 m.
However, at a lower spindle speed of 12,000 rpm the pri-mary wear region lasted up to approximately 1 min (O′A′ inFig. 7), which is equivalent to 200 holes drilled, or a totallength of cut of 1.19 m. Owing to the fact that this wear regionrepresents a heavily loaded system, chipping of the primarycutting edges, secondary cutting edges, and corners wereobserved (refer to Fig. 5).
ii. At spindle speed of 15,000 rpm, the secondary steady wearregion lasted approximately 1.35 min (between OA to OB),which is equivalent to 275 drilled holes or 1.62 meters of totalcutting length. At a lower speed of 12,000 rpm, the steadywear region lasted approximately 2.5 min (between O′A′ toO′B′), which is equivalent to 512 drilled holes, or 3 m of cut-ting length. In this region, the average tool abrasive wear ratewas 38.60 × 10−6 m/s (depth of average flank wear per unittime) at n = 15,000 rpm, as compared to 26.30 × 10−6 m/sobserved at spindle speed of 12,000 rpm.
iii. In the tertiary severe wear region (beyond OB and O′B′), thewear rate increased rapidly as shown in Fig. 7. At 15,000 rpm,the end of tool life, i.e., Vb = 225 �m, occurred at about2.12 min, which is equivalent to a total of 544 drilled holes or3.2 m of cutting length. At lower spindle speed of 12,000 rpm,the end of tool life was reached at about 3.17 min, which isequivalent to about 650 drilled holes or 3.8 m of total cuttinglength. The SEM images in Fig. 6(a)–(c) illustrate the exten-sive abrasive wear at the end of tool life at the chisel edge,the primary cutting edge, the secondary cutting edge, andthe corners, respectively, when n = 15,000 rpm.
iv. At higher spindle speed of 15,000 rpm, a large quantity ofcarbon deposit was found on the flank face of the primarycutting edge and the secondary cutting edge. Fig. 6(d) illus-trates the adherence of carbon on the flank face and thecorner of the drill during the continuous drilling process. The
chemical analysis performed at points on the flank face con-firmed the adhesion of carbon, as shown in Fig. 8. This showsthat the high cutting temperature obtained at higher spin-dle speed of 15,000 rpm, coupled with high flank wear, was
Fig. 9. Effect of flank wear on thrust force, cutting force, entry delamination and exitdelamination at spindle speed of 15,000 rpm and feed rate of 100 �m/rev.
responsible for burning of epoxy, adhesion of carbon fibreparticles and aggressive wear in the tertiary wear region.
The results of these tool wear experiments were used to estab-lish the relationship between the cutting velocity V (in m/min) andthe tool life T (in min), in the form of the conventional Taylor’sequation [5]:
VT0.56 = 3.08 (1)
The expected inverse relationship between the cutting velocityand the tool life is in agreement with other studies [15], and can beused to estimate the tool life at different cutting speeds and fixedfeed rate of 100 �m/rev. Having established the understanding ofvarious mechanisms and the rate of the flank wear, it was necessaryto analyze the effect of tool wear on cutting forces and eventuallythe hole quality attributes.
4.2. Tribological interactions in the drilling process
4.2.1. Effect of tool wear on cutting and thrust forcesBoth the thrust force and cutting force were found to increase
with the increase in flank wear, as shown in Fig. 9. Thrust force wasfound to be higher than the cutting force in the primary and thesecondary wear regions. However, in the tertiary wear region, thecutting force increases beyond the thrust force. This is likely due tothe high temperature built up on the tool with continuous drillingat such high speeds. This results in matrix burn-out and sudden risein the cutting effort in the tertiary region.
4.2.2. Effect of tool wear on entry and exit delaminationThe delamination at the hole entry and exit are defined by the
parameter � in terms of the hole diameter Dh and the maximumdelamination damage diameter Dd (Fig. 10(b)): � = (Dd − Dh)/Dh.Fig. 9 indicated that � increases with the increase in flank wear.It is also evident from the figure that as the flank wear increases,the cutting force (Fc) increases. Fig. 10(a) and (b) shows the entrydelamination at the end of tool life (average Vb = 225 �m) for 15,000and 12,000 rpm, respectively. Lower delamination was observed at15,000 rpm as compared to 12,000 rpm, due to the reduction inthrust and cutting forces. It is also interesting to observe that theprogressive increase in the cutting force and delamination stronglyfollowed the three wear regions. However, slight lag was observedin the tertiary wear region as clearly seen in Fig. 9.
delamination. It is observed that the steady increase in wear rate inthe primary and the tertiary wear regions strongly influenced thethrust force, which in turn caused higher delamination. At spin-dle speed of 15,000 rpm, no delamination was found to occur up
1028 S. Rawat, H. Attia / Wear 267 (2009) 1022–1030
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ig. 10. Delamination at the end of tool life: (a) entry delamination, n = 15,000 rpmxit delamination at n = 12,000 rpm.
o a critical thrust force of 70 N, which was found to be close tohe threshold thrust force of 65 N predicted by Ho-Cheng et al.’s
odel [8]. This corresponds to flank wear of approximately 90 �m.eyond this wear width, the exit delamination increased rapidly
n the primary wear region. The rate of delamination was found toe steady, and followed the flank wear rate in the secondary wearegion. Furthermore, in the tertiary wear region the exit delami-ation increased rapidly reaching a maximum of 30% at the end ofool life. Fig. 10(c) and (d) shows the exit delamination observedt the end of tool life for 15,000 and 12,000 rpm, respectively. Exitelamination of about 30% and 27% were found to occur due tohrust forces of 200 and 250 N, respectively, due to the excessiveank wear.
.2.3. Effect of tool wear on hole circularity errorHole circularity error, which is defined as the ratio of the tol-
rance zone bounded by two concentric circles within which each
ircular element of the surface must lie and the hole diameter, wasound to increase with the increase in flank wear at spindle speedf 12,000 and 15,000 rpm. Fig. 11 illustrates the effect of increasingank wear on the hole circularity error at entrance and exit. It is
ig. 11. Effect of flank wear on hole circularity and diameter error at entry and exit,t spindle speed of 15,000 rpm and feed rate of 100 �m/rev.
ntry delamination, n = 12,000 rpm, (c) exit delamination at n = 15,000 rpm, and (d)
evident from the figure that the hole circularity deteriorated pro-gressively with increase in wear, which in turn causes an increasein the thrust force. At spindle speed of n = 15,000 rpm, a hole circu-larity error of 0.17% was obtained, while at n = 12,000 rpm an errorof 0.19% was found to occur at the end of tool life. Fig. 10(a) and(b) shows the poor hole circularity for spindle speeds of 15,000 and12,000 rpm, respectively.
4.2.4. Effect of tool wear on hole diameter errorAs Fig. 11 shows, the flank wear was found to strongly influence
the hole diameter both at entry and exit. At the start of drilling,when the tool is sharp, the holes produced were oversized in therange of 0.08–0.12%, which was within the H8 hole tolerance stan-dard commonly used in the aerospace industry for 5 mm carbidedrill. However, the magnitude of the hole oversize decreased withthe progress in flank wear in the primary wear region. Beyond thiswear region, only undersized holes were produced due to increas-ing flank wear. Similar findings were reported in [13,16,17], whereinduring the assessment of coated HSS and tungsten carbide drills, outof tolerance holes were produced with the increase in flank wear. Itis evident from Fig. 11 that the decrease in the hole diameter closelyfollowed the progression of flank wear, with the hole undersize atentry increasing rapidly to about 0.20% and at exit to about 0.28%at the end of tool life in the tertiary wear region.
4.2.5. Effect of tool wear on hole surface roughnessHole surface roughness (Ra) was found to increase with the
increase in the number of drilled holes, in agreement with theresults reported in [11]. As discussed earlier, at the start of drilling,primary cutting edges, secondary cutting edges and chisel edge ofthe WC drill experienced chipping at the sharp corners. The initialchipping combined with increasing flank wear resulted in increasein both the thrust and the cutting forces. The increasing thrust forceresulted in micro-cracks at the ply interfaces, while the increase inthe cutting force resulted in extensive fiber full out, thus deteri-
orating the surface finish. Fig. 12 illustrates the increase in forceswith increase in flank wear, which in turn increased the hole surfaceroughness.
The results of this investigation show strong correlationbetween tool wear, cutting/thrust forces and the quality of the
S. Rawat, H. Attia / Wear 26
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ig. 12. Effect of flank wear on hole surface roughness at spindle speed of 15,000 rpmnd feed rate of 100 �m/rev.
rilled hole, in terms of geometric errors and surface finish. Theseesults have far reaching implications in industrial environment.hrough real-time monitoring of the cutting and thrust forces, toolanagement system and adaptive control strategies can be devised
o control and optimize the drilling process. Although the force sen-ors are expensive, they could be an order of magnitude cheaperhan the total cost of a single large aerospace component, consider-ng the cost of tooling, the composite material, inspection and otherirect and indirect cost components. To take the findings of this
nvestigation a step forward towards implementation in industrialnvironment, other sensors and monitoring technologies need toe considered. For example, acoustic emission sensors [36], vibra-ion of the tool and workpiece [37], spindle motor current [36], feed
otor current, and spindle mounted strain gage sensors [38], ultra-onic techniques [39] and vision sensors [40] have been used forn-line tool wear measurement. Any of these tool wear monitoringensors can be linked to artificial intelligence and neural networkraining methods to establish tool replacement strategies [41].
. Conclusions
In this investigation, a system approach was followed to investi-ate the wear mechanisms of WC drills, evaluate tool life and studyhe effect of tool wear on the quality of holes produced in wovenraphite epoxy composites at high speeds. Based on the experimen-al results, the following conclusions are drawn:
1. Fracture (chipping) at the start of drilling process, subsequentabrasion and possibly adhesion of carbon are the three domi-nant wear mechanisms observed in high speed drilling of wovengraphite epoxy composites. Abrasive wear on the flank face ofthe primary cutting edge, being more dominant than the wearon the rake face, is the main drive of the nonlinear behavior ofthis tribo-system.
. The transition of tribo-system from primary to secondary andfinally to the tertiary wear region governed the change in inter-mediate process variables, namely, thrust and cutting forces, andthus affect the hole quality, i.e., delamination, geometric errors,and surface roughness.
. The effect of cutting speed on tool wear was established in theform of Taylor’s equation.
. Strong correlations exist between the controllable process vari-ables (cutting and thrust forces) and tool wear as well as the
quality of the final hole (delamination, geometric errors and sur-face finish). The analysis has demonstrated the potential for useof force measurement for real-time assessment of tool wear, opti-mization of the drilling process and thus indirectly influencingthe product quality.
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7 (2009) 1022–1030 1029
Acknowledgements
This work was conducted under the partial financial sup-port of the Natural Sciences and Engineering Research Council ofCanada (NSERC), which the authors greatly appreciate. The authorsacknowledge the support of the Aerospace Manufacturing Technol-ogy Centre (AMTC), Institute for Aerospace Research (IAR), NationalResearch Council Canada (NRC), where the experiments were car-ried out. The technical inputs of Dr. Nejah Tounsi and Mr. DanenChellan, of the AMTC, are highly appreciated.
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