Drilling analysis of thin woven glass-fiber reinforced epoxy‎ composites

Accepted Manuscript Title: Drilling analysis of thin woven glass-fiber reinforced epoxynull composites Authors: U.A. Khashaba, A.A. El-Keran PII: DOI: Reference:

S0924-0136(17)30239-X http://dx.doi.org/doi:10.1016/j.jmatprotec.2017.06.011 PROTEC 15264

To appear in:

Journal of Materials Processing Technology

Received date: Revised date: Accepted date:

3-12-2016 4-6-2017 8-6-2017

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Drilling analysis of thin woven glass-fiber reinforced epoxy composites U.A. Khashabaa,b,* , A.A. El-Keranc a Mechanical

Engineering Department, Faculty of Engineering, King Abdulaziz University, P.O. Box 80204, Jeddah 21589, Saudi Arabia

b

Mechanical Design and Production Engineering Department, Faculty of Engineering, Zagazig University, P.O. Box 44519, Zagazig, Egypt c Department

of Mechanical Power Engineering, Mansoura University, Mansoura, Egypt

ABSTRACT In this study, the influence of cutting speed (16.3, 32.7 & 65.3 m/min) and feed (0.025 to 0.1 mm/rev) on cutting temperature, thrust force, torque, delamination factor, surface roughness and bearing strength in drilling thin woven glass-fiber reinforced epoxy (GFRE) composites was investigated, for the first time, grouped in one article. The results of drilling tests showed that the thrust force (Ft ), torque, and delamination factor are increased with increasing feed and vice versa for the induced temperature. Drilling of GFRE composites at lower speed (16.3 m/min) result in lower temperature rise, higher thrust force, higher torque and accordingly, higher delaminatio n factor compared with that drilled at 65.3 m/min. Drilling at temperature close to glass transition temperature is useful for reducing Ft , delamination and thus obtaining better bearing strength compared with drilling at lower temperature levels. The critical thrust force at the onset of push -out delamination was predicted for different uncut layers, in front of the drill chisel edge, using some analytical models. For the first time, in drilling [0°]8 woven composites, the [ABD] stiffness matrices was effectively represented by [0°/90°]4s cross-ply laminate with the advantages that the elastic properties were determined by the rule of mixtures (no need to perform mechanical tests). Keywords: Drilling thin composite laminate; Flank surface temperature; Thrust force; Torque; Delamination factor; bearing strength 1. INTRODUCTION The growing use of high-performance fiber composite polymer (FRP) materials in automotive and aircraft structural applications makes drilling holes are almost unavoidable operations for assembly/joining and repair purposes (Khashaba et al. 2017). For this reason, the effect of machining parameters on pin-joint bearing strength of glass-fiber reinforced epoxy (GFRE) composite will be investigated in the present study as a machinabilit y parameter. Because of the geometrical parameters of twist drill were varied along the cutting edge in the static conditions and changed during the drilling process due to the fact that the drill is simultaneously rotating and feed axially, all the points on the cutting edges are moving in a helical path resulting in a complex machining process and analysis.

*

Corresponding author. Tel.: +966553507515; Fax: +96626952181 E–mail addresses: [email protected], [email protected] (U.A. Khashaba).

Abhishek et al. (2015) pointed that the delamination is a major problem associated with drilling of fiber-reinforced composite materials, which tends to reduce structural integrity of the said material. The anisotropy, heterogeneity and the abrasion resistance of fibers lead to increasing the complexity of the drilling processes in FRP composites, which often result in delamination, fiber pull-out, spalling, fuzzing, and thermal degradation. The defects in the drilled holes represent a critical problem that can leads to catastrophic failure of the composite structure. In the particular case of the aircraft industry, the economic impact of the damage induced by drilling is significant, especially when considering the added value associated with the component when it reaches the assembly stage (Wong et al., 1982 and Debnath et al., 2015). Delamination size depend on the interaction of many parameters including: drill-point geometry, thrust force, temperature rise, back support distance, laminate thickness and stiffness, glass transition temperature of the matrix and fiber/matrix interfacial bonding. Two delamination mechanisms are always observed in drilling FRP composites, which are: peel-up delamination at drill entry and push-out delamination at drill exit. The peel-up delamination depends on the tool geometry as reported by some investigators (Hocheng and Dharan, 1990 and Luo et al., 2016) while, the push-out delamination depends on the thrust force of the drill point. Hocheng and Dharan (1990) reported that the cutting force acting in the peripheral direction generates a peeling force in the axial direction through the slope of the drill flute and accordingly, it can be considered the driving force for peelup delamination. The peripheral cutting force is a function of tool geometry and friction between the tool and workpiece. Luo et al., (2016) observed in drilling thin CRFE laminates that when uncut thickness was too thin to hold the drill bit, springback of workpiece enlarged the actual feed rate. The thrust force was decreased, because drill bits pushed on the material instead of cutting through it. The push movement in exit stage increases the extension of delamination. Khashaba et al., (2010) pointed out that the thermal conductivity of GFRP composites is very low (0.59 W/mC) compared to steel (= 53 W/mC), brass (= 109 W/mC) and Aluminum (= 210 W/mC). Therefore, the generated heat during the drilling operation cannot be conducted to the outside easily and thus it plays contrary roles in drilling process. The different thermal properties of polymer matrix and glass fiber makes the temperature rise in drilling FRP composites is a complex problem. Softening and re-solidification of the machined surface in drillin g FRP composites increase the stress concentration that localizes crack initiation and propagation, which leads to decreasing the bearing capacity of the drilled holes. Rawat and Attia, (2009) reported that temperature rise during drilling process leads to red ucing the workpiece stiffness, thrust force and thus less damage. On the other hand, Hocheng et al. (1993) showed that when the temperature rise is much higher than the glass transition temperature (T g ) the polymers become more ductile and their static friction coefficient rises. The ductile material is smeared over the surface of the borehole wall by the minor cutting edges of the rotating drilling tool. This causes detectable unevenness at the surface. For the same context, Merino-Pérez et al. (2016) found that drilling at temperatures over Tg of the matrix decreases the elastic modulus of the composite, accordingly the maximu m thrust force was decreased. Rawat and Attia, (2009) pointed that as the cutting temperature exceeds the melting temperature of the polymer, it will cause burning to the mat rix at hole edge and thus, the unreinforced fibers produce wavy rough surface.

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The following question remain open and the answer to it required better understanding of composite degradation owing to thermal and mechanical damages in drilling composite materials: what is the effect of matrix softening on reducing the cutting forces and thus the mechanical damages (e.g. delamination)?. Therefore, considerable attention has been paid in the last years by many researchers to analysis the cutting temperatu re in drilling FRP composites, which will be considered in the present study. To the author’s knowledge, the influence of cutting parameters (feed and speed) on cutting temperature, thrust force, torque, delamination, surface roughness and bearing strengt h in drilling thin woven composites was not investigated grouped in one article. Therefore, the main objective of the present work is to investigate the effect of feed (0.025, 0.05, 0.075 and 0.1 mm/rev), cutting speed (16.3, 32.7 & 65.3 m/min) on the mach inabilit y parameters in drilling thin woven glass -fiber reinforced epoxy (GFRE) composites. The drilling processes were performed using a CNC milling machine with a standard 6.5 mm coated twist drill. The investigated machinabilit y parameters included: flank surface temperature, thrust force, torque, delamination factor, bearing strength and surface roughness. Flank surface temperature was online monitored using instrumented drill with K-thermocouple. The thrust force and torque were online monitored and recorded using Kistler dynamometer model 9272. The peelup and push-out surface delaminations were measured using the AutoCAD method. Two pin -bearing joints were assembled in double-shear test fixtures to measure bearing strength of the drilled specimens. Th e critical thrust force at the onset of push-out delamination was predicted using some published theories. 2. EXPERIMENTAL WORK 2.1. Fabrication and characterization of woven GFRE composites 2.1.1. Fabrication of woven GFRE composites Drilling processes were conducted on thin woven glass -fiber reinforced epoxy (GFRE) composite laminates with 2.7±0.1 mm thickness and fiber volume fraction (Vf ) of 35%. The laminate was locally fabricated using hand layup technique. Details about the constituent materials of the fabricated laminate are illustrated in Table 1. 2.1.2. Tensile properties of woven GFRE composites The tensile properties of the fabricated woven GFRE laminate are determined using computer controlled universal testing machine model CMT5205/5305 MTS SYSTEMS (300 kN) at constant cross –head speed of 1 mm/ min . Five specimens are tested in tension in accordance with ASTM D 3039. The average value and the standard deviation of tensile strength (t ) are presented in Table 2. The Young’s moduli of GFRE specimens are estimated from the slope of the initial linear elastic portion of stress -strain curves, which always give small variation compared to the tensile results. This was attributed to the fact that the Young’s modulus is a material property that does not depend on the fracture mechanisms of the composites. Therefore, the Young’s modulus and Passion’s ratio are measured using two specimens equipped with longitudinal and transverse strain gauges. The strain gauges are connected to PC via 4-channel data acquisition model 9237 NI. The average values are presented in Table 2. The measured mechanical properties, Table 2, will be used to predict the critical thrust force at the onset of delamination as shown later.

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2.2. Drilling operations Drilling processes are implemented on CNC milling machine model EMCO FB-4 using cobalt coated HSS drill with 6.5 mm diameter. Experimental setup for online measuring thrust, torque and temperature is shown in Fig. 1. The drilling processes are performed at cutting speeds of 800 RPM (16.3 m/min), 1600 RPM (32.7 m/min), and 3150 RPM (65.3m/min). Four levels of feed ranging from 0.025–0.1 mm/rev with increment of 0.025 mm/rev have been selected based on the literature review (Luo et al., 2016, Chen, 1997 and Sadek et al., 2013). Two specimens (4-holes) were drilled for each cutting condition and the average values of the machinability parameters are considered for plotting the different relationships. 2.3. Measurements of the Machinability parameters 2.3.1. Thrust force and torque The thrust force and torque were online monitored and recorded using Kistler dynamometer model 9272, which is connected with PC through multichannel charge amplifier type 5070A and data acquisition (DAQ) type 5697A . The dynamometer was fixed on the machine table. The instrumented drill was mounted on three -jaws selfcentering chuck, which is fixed on the dynamometer using four independent jaws chuck. Special fixture was manufactured to hold the specimen in the machine spindle as shown in Fig. 1. The fixture with center hole of 15 mm was mounted into the machine spindle. The workpiece is clamped firmly in the fixture using metallic strip, with 15-mm hole at its center. The workpiece is fitted into the strip via U-slots of 2-mm depth and 20-mm width that was machined at the center of the strip. The workpiece and strip are assembled with the fixture through two bolts as shown in Fig. 1. These procedures confirm that the holes will be drilled on the longitudinal center of t he specimen, which is essential for bearing tests. 2.3.2. Temperature Flank surface temperature was online monitored using instrumented drill with K-thermocouple. The thermocouple can measures temperature up to 370 °C and its response time is 10 μs (Rames h et al., 2016). The thermocouple was embedded in a groove of 0.75 mm wide on the flank surface ahead of the cutting edge. The groove was machined using a dental high-speed carbide drills. The thermocouple was bonded using cold -hardening epoxy resin. The wires of the thermocouple were wounded on the spiral flute of the drill body and connected to a digital thermometer type 2809. The analog signal (mV=1o C) of the digital thermometer was recorded in the PC through USB DAQ type NI 6289. The temperature measuring setup was calibrated using Fluke thermometer. 2.3.3. Delamination The peel-up and push-out surface delaminations have been measured using the AutoCAD technique that was developed earlier by Khashaba (2004). The drilled specimen was scanned using high resolution flatbed color scanner. The measuring technique is suitable only for transparent or quasi-transparent materials in which the transmitted light from the scanner into the delaminated or damaged zone make it brighter, with different degrees, than the virgin material. The contrast, brightness and focusing utilities were applied to the acquired image to distinguish between the shadow (delaminated) zones and the virgin material. Consequently, the image appears to be darker. The image was analyzed using CorelDraw software, which can accurately determine the delaminatio n

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within 10-3 mm. The delamination factor was defined as the ratio of the maximu m delaminated diameter to the hole nominal diameter (6.5 mm). 2.3.4. Bearing strength The effect of cutting conditions on the bearing strength of the drilled woven GFRE composite was investigated using two-pins-bearing joints as shown in Fig. 2. The bearing test configuration was developed earlier by Turvey and Wang (2007). Each test specimen was drilled by two holes with the same cutting conditions at its longitudinal center and away 20 mm from each edge. The specimen was installed into two double -shear test fixtures model Instron T1058-65 via two pins ( = 6.4 mm) with clearance of 100 μm, which is recommended for bolted joints in aerospace structures (DiNicola and Fantle, 1993). The specimen was loaded in tension using computer controlled servohydraulic testing machine model Instron 8872 (10 kN) at a rate of 1.0 mm/min. The bearing strength of the drilled specimens was calculated from the following equation:

b 

P d .t

(1)

where P is the first peak load on the load-displacement curve, d is the hole diameter and t is the specimen thickness. 2.3.5. Surface roughness The surface roughness of the drilled hole wall (Ra ) was measured using Taylor Hobson form Talysurf series 2 surface profilometer with miniature bore stylus probe (112/2623) of 2μm radius, Fig. 3. The cut -off value was taken 0.8 mm. Four readings were taken at 90° to each other along the hole and the average value is reported. 3. RESULTS AND DISCUSSIONS 3.1. Mechanical Properties of Woven GFRE Composites The output of the universal testing machine is load (stress) vs time (displacement) curves. The output of DAQ is strain vs time. The true tensile modulus was calculated by dividing the slope of the initial linear portions of stress time curve by the slope of strain-time curve. Fig. 4 was constructed by substituting the curve fitting equation of stress-time relationship of testing machine into the strain-time data (of DAQ) the true stress -strain relationships. This figure also shows the apparent stress -strain curve of the testing machine. The apparent strain was estimated by dividing the machine displacement by the specimen gauge length. It is interested to note that the true Young’s modulus of woven GFRE composite is about 1.5 times higher than that of the apparent modulus. This result was attributed to the displacements caused by: the clearances between the moving elements and jo ints of the testing machine, deformations of the frame columns, loading heads and spindles of the testing machine itself. These displacements were added to the specimen elongation leading to decrease the apparent modulus of elasticity compared to the true value. The average tensile properties of five tests are illustrated Table 2. The experimen t al results in this table will be used later to predict the critical thrust force for delamination onset. 3.2. Temperature The GFRE composites have a very poor thermal conductivity owing to the lower thermal conductivity of the constituent materials. The thermal conductivity of epoxy (Km ) and E-glass fiber (Kf ) are 0.21 and 1.3 W/(mC)

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respectively (Gibson, 1994). The estimated thermal conductivity of woven GFRE composite using rule of mixtures equation is (= 0.59 W/mC), very low compared to steel (= 53 W/mC), brass (= 109 W/mC) and aluminum (= 210 W/mC). The accumulated heat around tool edge owing to poor thermal conductivity of woven GFRE composites is not easily conducted to the outside and accordingly, leads to softening the epoxy matrix, which can play a key role in reducing the workpiece stiffness, thrust force and mechanical damage as reported by Rawat and Attia (2009) and Chen (1997). On the other hand, the excessive heat can leads to burning the matrix at hole edge and thus, the unreinforced fibers produce wavy rough surface. Figs. 5 to 7 show the flank surface temperature over a complete cycle in drilling woven GFRE composite at cutting speeds of 16.3, 32.7 and 65.3 m/min respectively and different feeds. Because of the temperature was measured on the flank surface, that has clearance with the machined surface, the cutting edge temperature may has higher value especially for high cutting speed (short drilling time). The highest temperature is observed for the lowest feed (0.025 mm/rev) as shown in Figs. 5 to 7. This result is due to the fact that drilling at lower feed increases the machining time and thus, the higher flank surface temperature. The s lope of temperature-time curves (rate of temperature rise) was increased with increasing feed. Fig. 8 shows enlarge view of flank surface temperature over a complete cycle in drilling woven GFRE composite at 16.3 m/min and 0.025 mm/rev. The rate of temperature increase can be divided into four different stages as shown in Fig. 8. In the first stage (I), the chisel edge and part of the cutting edge that precedes the thermocouple entry the workpiece with lower rate of temperature rise. This behavior because o f the thermocouple not entry the machined hole and hence, the temperature of the chisel edge and part of the cutting edges transferred to it by conduction. The rate of temperature rise was increased as the cutting edges are fully engaged with the workpiece as shown in stage (II). In stage (III) the temperature rise up to the ultimate value but with lower rate compared to stage (II). This behavior is observed when chisel edge exit the machined hole (i.e. cooled in air) while, the cutting edges are still engaged with the workpiece. In stage (IV) the flank surface temperature was gradually decreased as the cutting edges exit the hole and slowly cooled in air. The four stages were clearly observed in drilling woven GFRE composite with the following cutting conditions: 16.3 m/min (for all feeds), 32.6 m/min (for 0.025 and 0.05 mm/rev feeds) and 65.3 m/min (for 0.025 mm/rev) as shown in Figs. 5 to 7 respectively. For these cutting conditions, the drilling time was  3.5 s, which is calculated from the following equation:

Drilling time 

Specimen thickness  length of drill point f .N

Length of drill point 

D/2 tan( / 2)

(2)

(3)

where f is the feed (mm/rev), N is the speed (RPM), D is the drill diameter (= 6.5 mm) and  is the drill point angle (=118o ). The drilled specimens at higher cutting conditions (i.e. lower drilling time < 3.5 s) showed sharp increase in the flank surface temperature up to the peak value at which the cutting edges start exit the hole and hence, the temperature was gradually decreased. This behavior was clearly observed for specimens that were drilled with the

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following cutting conditions: 32.6 m/min (0.075 mm/rev and 0.1 mm/rev) and 65.3 m/min (0.05 to 0.1 mm/rev ) as shown in Figs. 6 and 7 respectively. Fig. 9 shows the variation of peak flank surface temperature vs feed at different cutting speeds. The results in this figure showed that for the investigated cutting speeds the peak flank surface temperature was decreased with increasing feed. This behavior was due to the fact that the larger the feed value, the shorter the drilling time, Eq. (2), and accordingly, the lower the rise of flank surface temperature. The higher and lower flank surface temperatures were clearly observed for the drilled specimens at 32.7 m/min and 16.3 m/min respectively. The flank surface temperature in drilling woven composite at 65.3 m/min has approximately intermediate value between that observed at 16.3 and 32.7 m/min. This result was attributed to drilling at the highest cutting speed (= 65.3 m/min) result in lowest machining time and thus accumulate heat on the drill point. This behavior may be changed for thick/high-stiffness laminate, which is the subject of the following study. For the investigated feeds and speeds the maximu m rise in flank surface temperature was observed for specimen that was drilled at lower feed (0.025) and 32.7 m/min. 3.3. Cutting forces 3.3.1. Thrust force Fig. 10 shows enlarge view for the thrust force-time (hole depth) curve in drilling woven GFRE composite at 16.3 m/min and 0.025 mm/rev feed. Four different stages are observed for the thrust force-time (hole depth) relationship that is described as follows: In the first stage, a sharp linear increase in the thrust force (up to about 60% of its maximu m peak value) was observed at the entry of chisel edge into the workpiece as shown in Fig. 10. The sharply linear increase of thrust force represent the elastic loading of the workpiece owing to the negative rake angle of the chisel edge accompanied with its zero center speed. Therefore, the chisel edge will push-out (extrude) the materials instead of cutting it. The first peak thrust force or the load at the deviation from the linearity of the thrust -time curve represents the equivalent chisel edge force that needed to penetrates the workpiece (Fch  60% Ft ). This value is closed to that reported by Tsao and Hocheng (2003) in drilling woven carbon fiber composites. They measured the chisel edge thrust force using high speed steel (HSS) twist drill with 10 mm diameter. The specimens were pre-drilled with a pilot hole equal to the chisel edge width of the HSS drill. They found that the use of pilot hole can significantly reduce the thrust force by about 25–50%. The second stage starts after the chisel edge penetrates the workpiece surface layer. At this instant, the peak thrust force was followed by a valley due to stiffness reduction of the workpiece. The load is then further gradually increased in the form of peaks and valleys respectively due to the cutting edge enter new layer (uncut-chip area increases) and the remaining uncut layers in front of chisel edge is decreased by one. These peaks and valleys of the thrust force were repeated several times up to the maximu m peak value, which are observed at about 1.5 mm drilling depth ( 83% of drill point length and  53% specimen thickness). Decreasing the workpiece stiffness owing to cutting 53% of its thickness accompanied with the induced temperature result in sharp drop of the thrust force up to about 60% of its maximu m peak value as shown in the third stage of Fig. 10. At this instance, the chiseling edge start to exit the workpiece with a clear “knee” on the

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thrust-time (hole depth) curve. Increasing the thrust force up to 60% of the maximu m peak value (Ft ) during the initial penetration of the chisel edge into the workpiece and decreasing the thrust force a similar value during exit the chisel edge from the workpiece, demonstrate that the chisel edge force is equal to about 60% Ft . Therefore, it can be concluded that the chisel edge force in drilling thin (low stiffness) GFRE laminate, that is about 60% of Ft in the present study, can be accurately determined from the thrust force-time relationship (no needs to perform pilot hole technique). In the fourth stage, the two cutting edges exit the workpiece layer by layer and accordingly, the thrust force is gradually decreased over the drill point length (1.8 mm length) as shown of Fig. 10. Fig. 9 shows the effect of feed on the maximu m peak temperature and thrust force in d rilling woven GFRE composite at different cutting speeds. In general, it is evident that the thrust force increase with increasing feed as a result of increasing the undeformed chip cross -sectional area (= D.f/4) and thus the resistance of chip formation. In addition, increasing the temperature rise as a result of drilling at lower feed values, Fig. 9, leads to decreasing the laminate stiffness and thus thrust force. For the investigated feed values, the minimum thrust force was observed at 32.7 m/min at which the highest temperature (lower specimen stiffness) obtained as shown in Fig. 9. On the other hand, the maximu m thrust force was observed in drilling GFRE composite at 16.3 m/min, at which the minimum temperature (higher workpiece stiffness) obtained, compared to the other REMs. The thrust forces in drilling woven composite at 65.3 m/min have approximately intermediate values between that observed at 16.3 m/min (maximu m Ft ) and 32.7 m/min (minimu m Ft ). This behavior is agreed with the lower induced temperature (higher stiffness) at the former speed and the higher temperature (lower stiffness) at the latter speed as shown in Fig. 9. It is evident from Fig. 9 that for all the investigated feeds and speeds the maximum thrust force was observed for specimen that was drilled at maximu m feed (0.1 mm/rev) and minimum speed (16.3 m/min). 3.3.2. Torque Fig. 11 shows the effect of feed on the torque and cutting temperature at different cutting speeds. It is evident that the torque was decreased with decreasing feed as a result of increasing the temperature rise. This behavior because of the accumulated heat around tool edge owing to poor thermal conductivity of woven GFRE composites is (65 90°C) enough to softening the epoxy matrix. The softer matrix acts as a lubricant and hence, reduces the friction forces and moment of friction force on the margins. 3.4. Delamination 3.4.1. Peel-up delamination at drill entry The negative rake angle of chisel edge assisted by its zero speed at drill center leads to pushing (extrude) the GFRE specimen and accordingly, it was elastically deformed by about 0.1 mm as can be seen in Fig. 10. As a result, the top and bottom layers of GFRE composite have higher interlaminar shear stress and thus, the upper layer spirals (peel) up along the flute before it is completely machined. With the progress of the drilling process the thickness of the peel-up layers increases and thus the resistance to bending. Increasing the bending resistance of the peeled up layers accompanied with the high interlaminar shear stress leads to progressive peeling -up the composite upper

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layers around the hole boundary as shown in Fig. 12a and c. The images in Fig. 12 showed that the peel-up delamination in drilling thin woven GFRE composites is more pronounced compared to the push -out delamination. Fig. 13a shows the variation of peel-up delamination vs feed in drilling thin woven GFRE composites. Although the peel-up delamination depends on the tool geometry, it is evident from Fig. 13a that the measured values were increased with increasing feed. In addition, the peel-up delaminations are higher than push-out delamination. This behavior is contrary to the earlier result of Khashaba et al. (2010) for drilling thick (8.3 mm) woven GFRE composites. This result was attributed to the induced elastic deformation at the start of the drilling the thin (low stiffness) laminate, Fig. 10, which results in higher interlaminar shear stress compared with drilling of thick laminates. In addition, the elastic "springback" in drilling low stiffness GFRP composites increases the actual feed rate of the cutting edges that result, in combination with the helix angle of the tool, in ten sile forces superimposed to the feed force of the chisel edge. These tensile forces at the cutting edges can play a vital role in increasing peelup delamination in drilling processes. 3.4.2. Push-out delamination at drill exit As the drill approaches the exit, the number of uncut-layers is reduces and thus the resistance to bending. At a critical thickness, the bending stress becomes greater than the interlaminar strength and thus interlaminar crack is initiated before the laminate is completely penetrated by the drill. Further pushing down the uncut-layers by the drill chisel edge causes the crack to propagate. Therefore, the supporting span of the uncut-layers is increased, which leads to decreasing their flexural rigidity as reported by Ghasemi et al. (2011). Therefore, the uncut-layers start to fracture as the chisel edge proceeds to exit the laminate. As the cutting edges exit from the workpiece, the damages are extended in the form of push-out delamination as shown in Fig. 12b and d. Fig. 13b shows the variation of push-out delamination vs feed in drilling woven GFRE composites. The results in this figure showed that the push-out delamination is increased with increasing feed as a result of increasing thrust force. The results in Fig. 13b also s howed that for the investigated feed values, drilling of GFRE composites at 16.3 m/min result in lower temperature rise, higher thrust force, higher torque and accordingly, higher delamination factor compared with that drilled at 32.7 m/min. It is evident from Fig. 13b that drilling at both 32.7 m/min and 65.3 m/min has a marginal effect on the delamination factor. 3.5. Critical Thrust Force The thrust force Ft in drilling process is the sum of the vertical components on the cutting edges (2FZ1 ) and the chisel edge thrust force (FZ2 ) as shown in Fig. 14. At the start of drilling process, the thrust force increased sharply to about 60% Ft due to the zero speed at drill center accompanied with the negative rake angle of the chisel edge. Therefore, the chiseling edge has push-out (extrude) the materials instead of cutting it. The rest of the thrust force (40% Ft ) is the vertical components of the cutting edges (2FZ1 ), which can be measured using predrilled pilot hole with diameter equal the chisel edge width. The critical thrust force at the onset of push-out delamination can be predicted using some published models that are based on linear elastic fracture mechanics and classic plate bending theory as follows:

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Hocheng and Dharan (1990) showed that the critical load at the onset of crack propagation can be calculated using Timoshenko’s classical plate bending theory for a circular isotropic plate with clamped ends and concentrated loads as: 1/ 2

FHD

 8G E h 3     IC 1 2   3(1   ) 

(4)

where GIC (=211 J/m2 ) is the interlaminar fracture toughness in Mode I (Blake et al., 2012), E1 = elastic modulus,

 = Poisson’s ratio and h is the uncut depth under the tool. In the present work, h equal the last ply thickness = specimen thickness 2.8 mm/8 layers = 0.35 mm). The mechanical properties of the woven composites are illustrated in Table 2. Jain and Yang (1993) developed a model considering the delaminated area takes an elliptical shape with a and b being half the delamination sizes in the fiber and transverse directions respectively. The critical thrust force is given by;

FJY 

3



2GIC D*

(5)

where  = a/b, and D∗ is an expression of the stiffness of the laminate and is given by

D*  D11 

2 D12  2D66  2  D22 4 3

(6)

The critical thrust force is a function of ellipticity ratio a/b. Minimizing the critical thrust force with respect to a/b results in a/b = (D11 /D22 )1/4 . where Dij represents the bending stiffness which were calculated as follows (Tsao and Chen, 1997);

D11 

E11h3 12(1  12 21 )

D22 

E22h3 12(1  12 21 )

D12 

12E22h3 12(1  12 21 )

D66 

G12h3 12

(7)

where E11 , E22 are the longitudinal and transverse Young's moduli, respectively,  12 is the major Poisson’s ratio ( 21 =  12 E11 /E11 ), and G12 is the in-plane shear modulus of woven GFRE composite. The values of E11 , 12 , G12 and of the UD-layer are calculated from the mechanical properties of the constituent materials, Table 3, using rule of mixture Eqs. (8), (9) and (10) respectively, while E22 is estimated using Halpin-Tasi Eq. (11) (Barbero, 2010). The elastic properties of the constituent materials as well as the estimated values of E11 , E22 , 12 and G12 are illustrated in Table 3 (Khashaba, 2016).

E11  E f V f  Em (1  V f )

(8)

11   f V f  m (1  V f )

(9)

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V f (1  V f ) GmG f 1   ; or G12  G12 G f Dm (1  V f )G f  V f Gm

(10)

where E,  and G with subscripts f and m are Young’s modulus, Possion’s ratio and shear modulus of fiber and matrix respectively.

E22  Em

1   ..V f

where  

(11)

1  .V f E f / Em  1 E f / Em  

;   2 for circular fiber

Lachaud, Piquet, Collombet and Surcin

(2001) considered the uncut plate under the drill as a thin circular

orthotropic plate clamped on the laminate surface and subjected to either a uniformly distributed load or concentrated load as follows:

Distributed load model:

FLPCS  D

GIC D GIC D* / 8  8  24 (1 / 3)  ( D ' / 8D) 1  D ' / D*

Concentrated load model FLPCS C  8

2GIC D 1  ( D ' / 8D )

 24

GIC D* / 4 3  D ' / D*

(12)

(13)

Where D and D’ are calculated using relations of laminated plate theory

1 3 D  (3D11  2 D12  4 D66  3D22 )  D* 8 8 D' 

(14)

D11  D22 D12  D66  2 3

(15)

Zhang, Wang and Liu (2001) developed a general closed-form mechanical model for predicting the critical thrust force at which delamination is initiated at different ply locations in the form of ellipse with  = a/b. The developed model is given by

FZWL 

GIC

(16)

 (C3  K )

where C3 and K are a function of the uncut laminate properties (constituent material properties and lay -up) and ellipticity ratio. The values of C3 and K can be calculated for orthotropic unidirectional materials using the following equations:

C3 

  , K * 6D 9D*

(17)

Substituting Eq. (17) into (16) yields the same equations as derived by Jain and Yang, Eq. (5).

11

Based on the displacements, constitutive equations and strain energy calculations of Zhang et al. (2001), Gururaja and Ramulu (2009) proposed a modified exit-ply delamination model to predict the critical thrust force in drillin g multidirectional FRP composites as:

GIC

FGR 

(18)

c   3  k  3 

For orthotropic materials, the values of C3 and K can be calculated as follows:

c3 

  , k * 144D* 24D

(19)

Substituting Eq. (19) into Eq. (18), the critical thrust force expression becomes:

FGR 

12



GIC D*

(20)

The above models were used in the present work to predict the critical thrust force at the onset of push -out delamination. Based on the delamination shape in Fig. 12 and the previous observation of Khashaba et al. (2010) for drilling thick woven GFRE, the delamination can be approximated to a circular shape with  = a/b = 1. This assumption is the best for the used orthogonal woven GFRE composite, which has the same mechanical properties for both longitudinal and transverse directions. The layer elastic properties of the composite laminate, which were used to estimate the [ABD] stiffness matrices of the uncut thickness (layers), were determined by two different methods: The first method: the layer properties of [0°]8 woven laminate were measured experimentally as illustrated in Table. 3. The thickness of each layer is 2.8/8 = 0.35 mm. The second method: each layer of the [0°]8 woven laminate is considered as two unidirectional plies crossing at 90° angles with each other (two cross plies of 0° and 90° orientation). Accordingly the [0°] 8 woven laminate was represented as [0°/90°]4s symmetrical balanced orthotropic cross -ply laminate in which the extension bending coupling [B] matrix is zero along with A 16 , A 26 , D16 , D26 = 0. The thickness of each layer is 2.8/16 = 0.175 mm. The elastic properties of the unidirectional layer were determined from the properties of the constituent materials by rule of mixtures as indicated in Table 2. The critical thrust forces were calculated for different number of uncut layers (thickness) in front of chisel edge. The thickness of the uncut layers are 0.35, 0.7, 1.05 and 1.4 mm, which are equivalent respectively to 1, 2, 3 and 4 layers of woven composites and 2, 4, 6 and 8 layers of cross -ply laminate. The bending stiffness D11 , D12 , D22 and D66 as well as the values of D’ and D* are estimated using classical lamination theory for the different uncut layers of woven and cross -ply laminates as shown in Table 4. This table also shows the predicted critical thrust force at the onset of delamination using the above models. The predicted critical thrust forces of the last uncut layer (0.35 mm) using Hocheng-Dharan model, Jain and Yang model and Zhang et al. model are (63-73 N) in the range of the measured thrust forces (30-85 N) as shown in Table

12

4. Therefore, the delamination onset was clearly observed as shown in Fig. 13a and b. Because of the used models neglects the effect of temperature rise on the induced thrust force, the delamination onset was observed for cutting forces lower than the predicted values. For the unidirectional composites, Jain and Yang model is a special case of Zhang et al. model and accordingly, the two models give the same results. The predicted thrust forces using the concentrated load model of Lachaud et al. (2001) is better than that of the distributed load model. The results in Table 4 also showed that the predicted thrust forces using Gururaja and Ramulu model are highest compared with both the experimental results and the predicted values using the other models. Although Gururaja and Ramulu (2009) mentioned that for unidirectional composites, the developed model is a special case of Lachaud et al. model, Eq. (12), the results in Table 4 showed a clear difference for the predicted critical thrust forces using the two models. This because Eq. (25) in Gururaja and Ramulu (2009) considered that D’ = D*/2, which can’t be verified from the above Eqs. (6) and (15). The results in Table 4 indicated that the [0°]8 woven composite can be effectively represented by [0°/90°] 4s crossply laminate with the advantages that the elastic properties were calculated from the properties of the constituent materials using rule of mixtures (no need to perform mechanical tests). The relative variation between the predicted critical thrust forces of [0°]8 woven laminate and [0°/90°]4s cross-ply laminate, for the last uncut layer, is 1.7%. This variation was attributed to the effect of fiber undulation in woven composites, which leads to the obvious differences for D11 , D12 , D22 and D66 as shown in Table 4. The transverse displacement of clamped circular plate with radius “a” can be calculated for a uniformly distributed loading and concentrated loading using Eqs. (21) and (22) respectively, Lachaud et al. (2001).



w( r ) D 

FLPCS  D a 2  r 2 64 D a 2

w( r )C 

FLPCS C 16D



2

(21)





 2 r 2 2  2r ln a  a  r   

(22)

where w(r)D and w(r)C are the transverse displacements of the plate subjected to a uniformly distributed load and concentrated load respectively, and r is ranged from 0 (specimen center) to a (plate radius = 15 mm). The initial elastic deformation at the center of the workpiece owing to pushing (extrude) the GFRE specimen by chisel edge can be calculated by substituting r = 0 into Eqs. (21) and (22) as:

w(0) D 

FLPCS  D a 2 64D

(23)

w(0)C 

FLPCS C a 2 16D

(24)

The values of FLPCS-D (= 4028.5 N), FLPCS-C (= 2849.6 N), and D (= 27.084 Nm) are calculated for the woven laminate with 8-layers (no cutting at the initial elastic deformation) using Eqs. (12), (13) and (14) respectively. Substituting the calculated values of FLPCS-D , FLPCS-C, and D into Eqs. (23) and (24) yields; w(0)D = 0.166 mm and w(0)C = 0.471 mm. It is obvious that the estimated value of the initial elastic deformation, w(0)D , at the center of

13

circular clamped plate (r = 0) under a uniformly distributed load has a reasonable agreement with that observed (≈ 0.1 mm) in drilling of thin woven GFRE composite, as shown in Fig. 10. 3.6. Bearing Fig. 15 shows bearing strength vs feed in drilling thin woven GFRE composites at 16.3 and 32.7 m/min at which the minimum and maximu m temperature levels are observed respectively. The results in this figure showed that the feed values have insignificant effect on the bearing strength of the drilled specimens at 16.3 m/min. On the other hand, feed values have a clear effect on the bearing strength of woven GFRE composites at 32.7 m/min. In addition, bearing strength of the drilled woven GFRE composites at 32.7 m/min is higher than that drilled at 16.3 m/min. This behavior was attributed to the temperature induced in drilling thin woven composites at 16.3 m/min is lower than the glass-transition temperature (Tg = 72 °C, Khashaba, 2016). Therefore, the lower bearing strength of woven GFRE composites that were drilled at 16.3 m/min is almo st due to the mechanical damages caused by the lower induced temperature, higher thrust force and thus delaminations, as shown in Figs. 9 and 13, compared with drilling at 32.7 m/min. The higher induced temperature in drilling GFRE composite at 32.7 m/min, Fig. 9, reduces the thrust force and thus delamination factor, which lead to higher bearing strength than that drilled at 16.3 m/min as shown in Fig. 15. The induced temperature in drilling GFRE composite at 32.7 m/min and 0.25 mm/rev feed is (= 90°C, Fig . 9) higher than the glass-transition temperature (Tg = 72 °C), after which the mechanical properties is irreversible. Accordingly, the damage level is a trade-offs between the highest thermal damage (Temperature >Tg ) and the lowest mechanical damage (lowest thrust force). For these reasons, the bearing strength of the drilled GFRE composite at 32.7 m/min and 0.25 mm/rev feed is lower than that drilled at 0.5 and 0.75 mm/rev at which the induced temperature is 75 and 74°C respectively (closed to Tg ). The lower bearing strength of the drilled specimens at 32.7 m/min and 0.1 mm/rev was attributed to decreasing drilling temperature to about 68°C and thus increasing the thrust force and delamination factor as shown respectively in Figs. 9 and 13. 3.7. Surface roughness Fig. 16 shows surface roughness (Ra ) vs feed in drilling woven GFRE composites at 16.3 and 32.7 m/min at which the minimum and maximu m temperature levels are observed respectively. In general, the results in Fig. 16 showed that surface roughness of the drilled specimens at 16.3 m/min is lower than that of 32.7 m/min. The surface roughness of the drilled specimens at 16.3 m/min was slightly increased with increasing feed as a result of increasing the thrust force and thus the mechanical damages. In addition, drilling at high feed values lead to partially shearing the fibers and accordingly, higher rougher surface was obtained compared to drilling at lower feed.

The higher surface roughness of the drilled specimens at 32.7 m/min was attributed increasing the temperature rise than the Tg of GFRE composite especially at lower feed values (0.025 to 0.075 mm/rev). The highest surface roughness was observed at minimum feed, where the induced temperature is maximu m (= 90°C >Tg ). At this

14

temperature level the ductile matrix is smeared over the hole wall surface that causes detectable unevenness at the surface as reported by Weinert et al. (2007). The surface roughness of the drilled specimens at 32.7 m/min was decreased with increasing feed as a result of decreasing the induced temperature and thus the thermal damage. 5. CONCLUS ION The present work was devoted to investigate, experimentally and analytically, the effect of machining parameters on the machinability parameters in drilling thin woven glass -fiber reinforced epoxy laminates using a standard 6.5 mm coated twist drill. The machining parameters include: cutting speed (16.3, 32.7 & 65.3 m/min) and feed (0.025 to 0.1 mm/rev with increment of 0.025 mm/rev). The machinability parameters include: flank surface temperature, thrust force, torque, delamination factor, surface roughness and bearing strength. The following concluding remarks can be drawn from the present study: For the investigated cutting speeds the thrust force (Ft ), torque, and delamination factor are increased with increasing feed and vice versa for the induced temperature. For the investigated feed values, drilling of GFRE composites at 16.3 m/min result in lower temperature rise, higher thrust force, higher torque and accordin gly, higher delamination factor compared to those drilled at 32.7 m/min. The critical thrust force at the onset of push out delamination was predicted for different uncut layers in front of chisel edge using some analytical models, which are based on linear elastic fracture mechanics and classic plate bending theory. The stiffness matrices of the uncut composite layers were calculated by two different methods: in the first method, the elastic properties of [0°] 8 woven laminate were measured experimentally. The thickness of each layer is 2.8/8 = 0.35 mm. In the second method, the [0°]8 woven laminate was effectively represented by [0°/90°]4s cross-ply laminate with the advantages that the elastic properties are calculated from the properties of the constituen t materials using rule of mixtures (no need to perform mechanical tests). The predicted critical thrust forces of the last uncut layer (0.35 mm) using Hocheng-Dharan model, Jain and Yang model and Zhang et al. model are in the range of the measured thrust forces (30-85 N) and accordingly the delamination onset was clearly observed in drilling woven GFRE composites. Because of the used models neglects the effect of temperature rise on the induced thrust force, the delamination onset was observed for cutting forces lower than the predicted values. The lower thermal conductivity of woven GFRE composite (= 0.59 W/mC) compared to steel (= 53 W/mC), brass (= 109 W/mC) and aluminum (= 210 W/mC), makes the temperature measurement in drilling process is essential for the analysis of the machinability parameters. The accumulated heat around tool edge can play conflicting roles, either increasing the thermal damage, by softening the epoxy matrix, or decreasing the mechanical damage via reducing the workpiece stiffness, thrust force and moment of friction force on the margins. Drilling at temperature closed to glass transition temperature (72°C) is useful for reducing thrust force, delamination and thus obtaining better bearing strength compared with drilling at lo wer temperature levels. Therefore, the lower bearing strength of woven GFRE composites that were drilled at 16.3 m/min was attributed to the highest thrust force and thus delaminations, where the induced temperature is (62-71°C) lower than Tg of GFRE composite. For the investigated feed values, the surface roughness of the drilled specimens at 32.7 m/min (temperature > Tg ) is higher than that of 16.3 m/min (temperature < Tg ). Therefore, it can be concluded that in drilling thin woven

15

GFRE composite the thermal damage has more pronounced effects on surface roughness compared to the mechanical damage. REFERENCES Abhishek K, Datta S, Mahapatra S. Optimization of thrust, torque, entry, and exist delamination factor during drilling of CFRP composites. Int J Adv Manuf Tech 76(2015) 401-416. Blake SP, Berube KA, Lopez-Anido RA. Interlaminar fracture toughness of woven E-glass fabric composites. J Compos Mater, 46(2012) 1583–1592 Barbero EJ. Introduction to Composite Materials Design. 2nd Edition, CRC Press (2010) Chen WC. Some experimental investigations in the drilling of carbon fiber reinforced plastic (CFRP) composite laminates. Int J Mach Tool Manuf 37-38(1997)1097–108. Debnath K, Singh I, Dvivedi A. Rotary mode ultrasonic drilling of glass fiber-reinforced epoxy laminates. J composite materials 49 (2015) 949-963. DiNicola AJ and Fantle SC. Bearing strength behavior of clearance-fit fastener holes in toughened graphite/epoxy laminates. In: Camponeschi ET (ed.) Compos Mater: testing and design (ASTM STP 1206), vol. 11. Philadelphia, PA: ASTM, 1993, pp.220–237. Ghasemi FA, Hyvadi A, Payganeh G, Arab NBM. Effects of Drilling Parameters on Delamination of Glass Epoxy Composites. Aust J Basic Appl Sci 5(2011) 1433-1440. Gibson RF. Principles of composite material mechanics. Mc-Graw-Hill, Inc; 1994. Gururaja S, Ramulu M. Modified Exit-ply Delamination Model for Drilling FRPs. J Compos Mater 43(2009)483-500 Hocheng H, Dharan CKH. Delamination during drilling in composite laminates. J Eng Ind, ASME 112(1990)236–239. Hocheng H, Pwu HY, Yao KC. Machinability of some fiber reinforced thermoset and thermoplastics in drilling. Mater Manufact Process 8(1993)653–82. Jain S, Yang DCH. Effects of feedrate and chisel edge on delamination in composites drilling. Transactions of the ASME, J Eng Ind 115(1993): 398-405. Khashaba UA, El-Sonbaty IA, Selmy AI, Megahed AA. Machinability Analysis in Drilling Woven GFR/Epoxy Composites: Part I- Effect of Machining Parameters. Compos Part A–Appl S 41(2010)391-400 Khashaba UA, Khdair AI. Open hole compressive elastic and strength analysis of CFRE composites for aerospace applications. Aerosp Sci Technol, 60 (2017) 96-107. Khashaba UA. Delamination in drilling GFR-thermoset composites. Compos Struct 63 (2004) 313–327 Khashaba UA. Drilling of polymer matrix composites: A Review. J. Compos Mater, 47 (2013) 1817–1832. Khashaba UA. Improvement of toughness and shear properties of multiwalled carbon nanotubes/epoxy composites. In-press, Polymer composites (2016), DOI: http://dx.doi.org/10.1002/pc.24003. Lachaud F, Piquet R, Collombet F, Surcin L. Drilling of Composite Structures. Comput Struct 52(2001)511-516. Luo B, Li Y, Zhang K, Cheng H, Liu S. Effect of workpiece stiffness on thrust force and delamination in drilling thin composite laminates. J Compos Mater 50 (2016) 617-625 Merino-Pérez JL, Royer R, Merson E, Lockwood A, Ayvar-Soberanis S, Marshall MB. Influence of workpiece constituents and cutting speed on the cutting forces developed in the conventional drilling of CFRP composites. Comput Struct 140 (2016) 621–629.

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Ramesh B, Elayaperumal A, Satishkumar S, Kumar A, Jayakumar T, Dinakaran D. Influence of cooling on the performance of the drilling process of glass fibre reinforced epoxy composites. Arch Civ Me ch Eng 16(2016)135–146. Rawat S, Attia H. Characterization of the dry high speed drilling process of woven composites using Machinability Maps approach. CIRP Ann-Manuf Techn 58 (2009) 105–108 Sadek A, Attia MH, Meshreki M, Shi B. Characterization and optimization of vibration-assisted drilling of fibre reinforced epoxy laminates. CIRP Ann-Manuf Techn 62 (2013) 91–94. Tsao CC, Chen WC. Prediction of the location of delamination in the drilling of composite laminates, J Mater Process Tech 70(1997)185-189. Tsao CC, Hocheng H. The effect of chisel length and associated pilot hole on delamination when drilling composite materials. Int J Mach Tool Manu 43(2003)1087–1092. Turvey GJ. Wang P. Failure of pultruded GRP single-bolt tension joints under hot–wet conditions. Comput Struct 77 (2007) 514–520. Weinert K, Brinkel F, Kempmann C, Pantke K. The dependency of material properties and process conditions on the cutting temperatures when drilling polymers. Prod. Eng. Res. Devel. (2007) 1:381–387 Wong TL, Wu SM, Croy GM. An analysis of delamination in drilling composite materials. 14th National SAMPE Technology. Conf., p. 471. Atlanta, GA. (1982). Zhang LB, Wang LJ, Liu XY. A Mechanical Model for Predicting Critical Thrust Forces in Drilling Composite Laminates. P I Mech Eng B-J Eng 215(2001)135-146.

17

U-slot machined in the clamping strip Strip

Workpiece

Drilled 1st hole

Bolt

CNC milling machine

Machine spindle Workpiece fixture

Workpiece

Drilling 2nd hole Kistler DAQ 5697A for thrust & torque

Instrumented drill with thermocouple 3-Jaw chuck

PC for online monitoring thrust, torque & temperature.

Digital thermometer type 2809

M ultichannel charge amplifier 5070A

NI USB DAQ 6289 for temperature Fig. 1. Experimental setup for online measuring thrust, torque and temperature

18

4-Jaw chuck

Kistler dynamometer 9272

20

P

40

E

W=20

P

Fig. 2. Dimensions of test specimen.

Fig. 3. Talysurf series 2 surface profilometer

19

Fig. 4. Stress-strain curves of GFRE composites

20

Fig. 5. Temperature over drilling cycle of GFRE composite at 16.3 m/min and different feeds.

Fig. 6. Temperature over drilling cycle of GFRE composite at 32.7 m/min and different feeds.

Fig. 7. Temperature over drilling cycle of GFRE composite at 65.3 m/min and different feeds.

21

Chisel edge start exit the hole

Cutting edges exit the hole

III Cutting edges engaged with the workpiece

II

IV

Drill point exit the hole and gradually cooled in air

Thermocouple junction

I Heat transferred from chisel edge to thermocouple junction

Length of Drill point

Laminate thickness

(D/2)/tan(59)

= 2.8 mm

= 1.8 mm

Hole depth (mm)

Fig. 8. Temperature over drilling cycle of GFRE composite at 16.3 m/min and 0.025 feed.

22

120 Temperature 16.3 m/min 32.7 m/min 65.3 m/min

Temperature (oC)

80

100

60

80

40

60

Thrust force 16.3 m/min 32.7 m/min 65.3 m/min

20

0

40

20 0

0.025

0.05 0.075 feed (mm/rev)

0.1

0.125

Fig. 9. Peak thrust force and temperature vs feed at different cutting speeds

23

Thrust force (N)

100

Cutting edge enter new layers

III

 60% of maximum Ft

Elastic loading of workpiece

II

Workpiece elastic deformation

“Knee” at chisel edge exits the workpiece

I

Thrust force dropped about 60% Ft at chisel edge exit

Maximum Ft

Thrust force gradually decreased while the cutting edges exit the workpiece layer by layer

Work stiffness lost layer by layer

IV

Laminate thickness

Cutting edges exit the workpiece

(D/2)/tan(118/2) Hole depth (mm)

2.8 mm = 1.8 mm Fig. 10. Thrust force over drilling cycle of GFRE composite at 16.3 m/min and 0.025 feed.

24

Fig. 11. Peak torque and temperature vs feed at different cutting speeds

(a)

(c)

Peel-up delamination at drill entry the hole

Peel-up delamination at drill entry the hole

(b)

(d) Push-out delamination at drill exit the hole

Push-out delamination at drill exit the hole

Fig. 12. Delaminations in drilling woven GFRE composites at feed of 0.1 mm/rev: (a) and (b) are respectively the peel-up and push-out delaminations at 32.7 m/min. (c) and (d) are respectively the peel-up and push-out delaminations at 65.3 m/min.

25

(a)

(b)

Fig. 13. Delamination vs feed at different cutting speeds: (a) Peel-up delamination, (b) push-out delamination

Ft

Fig.14. Force analysis in drilling (Khashaba, 2013).

26

Fig. 15. Bearing strength vs feed at different cutting speeds

Fig. 16. Surface roughness vs feed at different cutting speeds

27

Table 1. The constituent materials of the composite laminates Material

Type Plain E-woven roving glass-fiber, 3.24 g/cm2

Reinforcement

Density of glass fiber = 2.58 g/cm3 Yarn count: 3.5 yarns/cm for the warp and weft fibers. Matrix

Epocast 50-A1 resin(100 parts by weight ) Hardener 946 epoxy (15 parts by weight). Huntsman Advanced Materials Americas Inc.

Table 2. The mechanical properties of woven GFRE composites Young’s modulus (GPa)

Poisson's ratio

Tensile strength (MPa)

 12 =  21

Standard deviation

E11 =E22

Standard deviation

t

Standard deviation

0.294

0.0149

15.22

0.117

175.02

3.908

Table 3. The mechanical properties of the constituent materials and the estimated properties of the UD-layer Properties of the constituent materials

Estimated properties of UD-Layer

Materials E (GPa)

E-glass

72.4

G (GPa)

Poisson’s ratio, 

28.84

0.22

E11

E22

G12

(N/m2 )

(N/m2 )

(N/m2 )

27.57 x109 Epocast 50-A1

3.43

1.45

7.94 x109 2.17x109

0.32

Table 4. Predicted critical thrust force at the onset of delamination for different uncut layers. [0°]8 plane woven laminate

[0°/90°]4s cross-ply laminate

1

2

3

4

1

2

3

4

Thickness (mm)

0.35

0.7

1.05

1.4

0.35

0.7

1.05

1.4

D11 x10-2 (Nm)

5.95

47.62

160.72

380.96

10.09

73.56

222.21

502.22

D12 x10-2 (Nm)

1.75

14.00

47.25

112.00

0.83

6.62

22.35

52.97

No. of uncut layers

28

 12

0.285

D22 x10-2 (Nm)

5.95

47.62

160.72

380.96

2.90

30.42

128.74

329.65

D66 x10-2 (Nm)

0.78

6.21

20.95

49.67

0.78

6.20

20.93

49.61

D* x10-2 (Nm)

14.11

112.85

380.87

902.81

14.58

116.67

393.75

933.32

D' x10-2 (Nm)

6.79

37.55

126.72

300.37

7.03

56.27

189.90

450.13

FHD (N)

63.0

178.1

327.3

503.9

FZW = FJY (N)

72.7

205.7

377.8

581.7

73.9

209.1

384.2

591.5

FLPCS-C (N)

129.6

356.2

654.4

1007.5

131.8

372.8

684.8

1054.4

FLPCS-D (N)

202.0

503.6

925.1

1424.3

205.5

581.3

1067.9

1644.1

FGR (N)

205.7

581.7

1068.7

1645.4

209.1

591.5

1086.6

1673.0

29