Materials Journal of Composite

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Surface Texture and the Stress Concentration Factor for FRP Components with Holes D. Arola and M. L. McCain Journal of Composite Materials 2003; 37; 1439 DOI: 10.1177/0021998303034462

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D. AROLA* AND M. L. MCCAIN Department of Mechanical Engineering University of Maryland, Baltimore County 1000 Hilltop Circle, Baltimore, MD 21250, USA

(Received June 25, 2002) (Revised January 8, 2003)

ABSTRACT: The influence of hole quality on the mechanical behavior of fiber reinforced laminates was studied. Holes were introduced in tensile specimens of a graphite/epoxy (Gr/Ep) laminate using an abrasive waterjet and commercial drills (diamond coated tungsten carbide twist drills or tungsten carbide drill-reamers). The machined surfaces were characterized using contact profilometry and the surface texture was used in estimating the effective stress concentration factor (KKt). Utilizing the macroscopic stress concentration posed by the hole and KKt, the total stress concentration was estimated using the principle of superposition. The Gr/Ep specimens were then loaded in tension and acoustic emission was used to monitor the failure process. The apparent stress concentration factor (KtðappÞ) of the tensile specimens was determined from the ratio of tensile strengths of coupons without holes to that of specimens with holes. Based on results from tension tests the KtðappÞat first fiber failure ranged from 2.50 to 3.40. The stress concentration factors determined from experiments were within 6% of that predicted using superposition and KKt. Although the hole quality was dependent on the method of machining and drilling, results from this study confirm previous reports that there is no correlation between the surface texture and first fiber failure or ultimate tensile strength of Fiber Reinforced Plastics (FRPs) with open holes. Holes introduced using worn diamond coated twist drills exhibited the lowest surface roughness but resulted in a significant reduction in first fiber failure strength. Results from this study indicate that surface texture and KKt cannot be used for a reliable estimate of hole quality in FRPs, especially for holes produced with worn cutting tools.

KEY WORDS: drilling, fiber reinforced plastics, hole, stress concentration, surface texture.

*Author to whom correspondence should be addressed. E-mail: [email protected]

Journal of COMPOSITE MATERIALS, Vol. 37, No. 16/2003 1439

0021-9983/03/16 1439–22 $10.00/0 DOI: 10.1177/002199803034462 ß 2003 Sage Publications

INTRODUCTION

IBER REINFORCED PLASTICS (FRPs) are used in the design of primary and secondary Fstructural components for a variety of applications. Component parts are molded to near net-shape and often require finish machining and secondary features to facilitate assembly. Adhesives, rivets, and bolts are used in joining FRP components but mechanical fasteners generally offer sustained reliability [1–4]. Therefore, postmold drilling is by far the most common machining process used in the development of FRP structures [2,5,6–9]. Several problems have been reported in drilling FRPs including entrance and exit ply delamination, matrix depletion, tool wear, and fiber pullout [5,7,9–13]. Inferior hole quality accounts for an estimated 60% of all part rejections and represents a costly manufacturing concern [14]. In addition to the economic burden, drilling damage may reduce the component’s strength. Delamination, waviness and roughness of the holes interior, axial curvature, and roundness errors may influence the mechanical behavior of laminates with holes [6–8,14–16]. While considerable effort has been placed on establishing drilling parameters that minimize damage, the influence of hole quality on part performance has received less attention. According to experimental results Wood [17] postulated that hole quality does not affect the ultimate strength of FRPs. Similarly, Tagliaferri et al. [18] concluded that the tensile strength of polymer composites with open holes was not affected by the hole quality whereas the bearing strength was. Persson et al. [19] found that while the compressive strength of carbon/epoxy laminates was not dependent on the hole quality the monotonic and fatigue strength was significantly reduced under pin loading. Though experimental studies have concluded that there is no effect of surface texture on the tension or compression behavior of FRP laminates with open holes, the effects of hole quality on the strength of FRPs cannot be disregarded. The influence of surface texture and surface integrity of holes on the mechanical behavior of component parts must be considered in a thorough design evaluation. The primary objectives of this study were to quantify the influence of hole quality on the tensile properties of FRP components with open holes and to evaluate an analytical approach to account for hole quality in design. A simple methodology is presented to account for hole quality on the strength of FRP components with drilled and machined holes. The approach is evaluated through a comparison of predictions with experimental results.

BACKGROUND

The influence of machining and the resulting machined edge quality on the mechanical behavior of component parts is a concern that accompanies the development of all new structural materials. Machined edge quality comprises the surface texture, surface integrity and process dependent defects that result from material removal. The surface texture describes external features of the machined surface geometry (e.g. lay and roughness) while the surface integrity encompasses subsurface qualities (e.g. heat affected zone, subsurface cracks, fiber pullout, etc).

Machined Edge Effects

The effects of machined edge quality on the strength and stiffness of FRPs without holes has been studied in detail. An early study of machined edge quality compared the effects of laser machining and abrasive waterjet (AWJ) machining on the strength of several FRP laminates [20]. Though laser machining was found to be detrimental to the tensile strength, there was no degradation in strength resulting from AWJ machining. Colligan and Ramulu [21] evaluated the ultimate compression strength of graphite/epoxy (Gr/Ep) laminates machined using diamond abrasive cutters and found that the strength decreased with increasing surface roughness. However, the reduction in strength resulting from edge ply delamination was more significant than the effects of surface roughness. The influence of machining defects and edge quality on the strength of FRPs has also been examined under quasi-static and dynamic flexure [22,23]. In an evaluation using Gr/Ep and graphite bismaleimide (Gr/Bmi) laminates the flexure strength was found to be significantly dependent on the manufacturing process, surface texture and surface integrity. Laminates trimmed with single-point cutting tools underwent the largest reduction in strength due to subsurface damage that was introduced in off-axis plies [23]. Net-shape machining, surface texture and their influence on the fatigue behavior of Gr/Bmi laminates have also been evaluated under fully reversed flexural fatigue loading [24]. At two different amplitudes of cyclic loading it was found that the reduction in stiffness increased with the machined edge surface roughness. A complete review of machined edge effects in composite materials is presented in [25]. Based on results from past studies on edge effects, the quality of drilled holes is expected to affect the mechanical behavior of FRPs with open holes. An analytical treatment of holes must consider the effects of edge quality and the change in stress distribution posed by the macroscopic notch.

Notches and Holes

The influence of notches and holes on the mechanical behavior of FRP materials has been approached using linear elastic fracture mechanics (LEFM) and stress concentration factors. In fact, several models have been proposed for estimating the strength of composite materials with holes [26–29]. A comprehensive review of notches in composite materials and their treatment is available in [30]. The primary objective of many existing formulations is to account for size effects posed by the gradient in stress distribution about the discontinuity. Nevertheless, these formulations do not account for damage resulting from the material removal process and the potential for machined edge effects on strength.

Treatment of Surface Texture

The surface texture resulting from net-shape machining or drilling may be viewed as a series of geometric irregularities that are introduced on the surface of a component. The influence of geometric features on the strength of engineering components is traditionally approached through the use of a stress concentration factor (Kt). Neuber [31] proposed a semiempirical relationship to describe the stress concentration posed by surface roughness according to sffiffiffiffiffiffiffiffiffi R K ¼ 1 þ n z ð1Þ t

Downloaded from http://jcm.sagepub.com at UNIV OF MARYLAND BALTIMORE CO on March 19, 2007 © 2003 SAGE Publications. All rights reserved. Not for commercial use or unauthorized distribution. 1442 D. AROLA AND M. L. MCCAIN where Rz and are the ten-point roughness and notch radius, respectively. The stress state is represented by the empirical factor n (n ¼ 1 for shear and n ¼ 2 for tension) and refers to the ratio between spacing and height of surface irregularities. For surfaces resulting from mechanical processes ¼ 1 has been suggested [32]. Arola and Ramulu [33] proposed an alternative expression to the Neuber rule that quantifies the effects of machined surface texture on the strength of FRPs in terms of an effective stress concentration factor. The effective stress concentration factor (KKt)is defined according to Ra Ry KKt ¼ 1 þ n ð2Þ Rz

where Ra, Ry, Rz, and are the average roughness, peak-to-valley height roughness, ten-point roughness, and effective notch root radius, respectively. Important factors including the material and load type are accounted for through the empirical constant (n). In general, n ¼ 2 is recommended for uniform tension and n ¼ 1 for shear loads; the factor may be modified for the stress state and material as necessary. The model has been used successfully in evaluating the effects of surface texture on the strength of FRPs under static and dynamic loads [33], and in estimating the effective fatigue stress concentration factor (KKf ) for the machined surface of metals and FRP materials [34,35].

Superposition of Stress Concentration Factors

Damage at the boundary of a hole could promote premature failure of FRP components with drilled holes and must be considered in addition to the macroscopic stress concentration. Therefore, it is necessary to consider the superposition of stress concentrations posed by the drilled hole and process damage. Paul and Faucett [36] examined the near-field stress distribution resulting from the superposition of semi- circular edge notches using photoelasticity. Previous investigations by Mowbray [37] and James [38] suggested that when a small notch of stress concentration factor Kt2 was placed in the region of maximum stress of a second larger notch (Kt1), the total stress concentration factor is simply the product of the two individual factors. Mathematically this can be expressed as

Kt ¼ Kt1ÁKt2 ð3Þ

where Kt1 and Kt2 are the stress concentration factors for the individual notches (Figure 1(a)). Paul and Faucett [36] found that this closely agreed with their experimental results. Mitchell [39] found that validity of the superposition principle requires the second notch to be smaller in size than the primary notch and in close proximity. Using the principle of superposition for the individual stress concentration factors posed by a drilled hole with surface texture (Figure 1(b)) the total stress concentration can be described by

Kt ¼ KtðholeÞÁKK t ð4Þ

σ (a)

K t2 K t1

(b) σ

K t2 K t1

Figure 1. Superposition of the stress concentration factors: (a) hole with second smaller notch; (b) drilled hole with damage.

The macroscopic stress concentration for the drilled hole (KtðholeÞ) can be defined according to design charts [40] or an alternative model which accounts for size effects [26]. The effective stress concentration factor (KKt) for the hole should be estimated in terms of the surface texture according to Equation (2) based on an evaluation of the plies with the most significant machining damage.

MATERIALS AND METHODS

An experimental investigation was conducted to determine the influence of hole quality on the tensile strength of FRP laminates and to evaluate use of the superposition principle for the design of FRP components with holes. Gr/Ep laminates with a stacking sequence of [(0/45/90/À45)3]S were vacuum bag molded using Fiberite Hy-E 3034K prepreg comprised of standard modulus IMIS graphite fibers and Fiberite 934 epoxy resin. Each ply was 125 mm in thickness. Mechanical properties of the prepreg are listed in Table 1. The laminates consisted of 24 plies and had a nominal thickness of 3.2 mm.

Drilling, Machining and Hole Quality

Straight-sided tensile specimens with dimensions of 38 Â 350 mm were machined from the Gr/Ep laminates using a numerical slicer/grinder1 and #220 mesh diamond abrasive

1K.O. Lee S3818EL Surface Grinder and Slicer.

Table 1. Mechanical properties of the Hy-E 3034 K prepreg.

E11 (GPa) E22(GPa) G12(GPa) n12 Xt(GPa) Yt (MPa) S (MPa) 138 10.3 4.8 0.28 1.9 61 117

Table 2. Process conditions used for drilling and AWJ machining.

Drilling Diameter (mm) Cutting Speed (rpm) Feed Rate(mm/min) Description Drill A 6.35 3500 267 CVD diamond coated 9.53 2350 179 WC twist drill (new) Drill B 6.35 3500 267 CVD diamond coated 9.53 2350 179 WC twist drill (worn) Drill C 6.35 1500 76 WC drill-reamer 9.53 1000 51 (new) Machining Diameter Traverse Speed Garnet Flow Rate Abrasive Size (mm) (mm/min) (g/s) (Mesh #) AWJ A 6.35 226 4.5 80 9.53 343 4.5 80 AWJ B 6.35 130 4.5 80 9.53 175 4.5 80

WC ¼ Tungsten Carbide; CVD ¼ Chemical Vapor Deposition.

saw with continuous coolant; the average surface roughness (Ra) of the machined edges was 0.2 mm. Holes were either drilled or machined in the center of the tensile specimens with diameter of 6.35 or 9.53 mm. All conventional drilling was completed on a vertical milling center2. A fixture was used for backing, which minimized deflection of the specimen and exit-ply delamination. Three different drills were used for both hole sizes including two diamond coated tungsten carbide (WC) drills (Drills A and B) and one tungsten carbide drill-reamer (Drill C). Both diamond coated drills were standard 2-flute spiral twist drills with a 118 four-facet point angle. One of the diamond coated drills of each size (Drill B) was used in repeated drilling of a glass/polyester laminate to introduce cutting edge wear. The wear was considered sufficient when visible to the naked eye and was achieved after approximately 0.75 m of drilled hole length. Drilling of the Gr/Ep laminate was conducted according to the manufacturer’s recommendations without coolant (Table 2). An AWJ3 was also used to introduce holes in selected tensile specimens to expand the range of surface texture available from use of conventional twist drills. Two qualities were obtained for both hole sizes and were designated ‘‘AWJ A’’ and ‘‘AWJ B’’ (Table 2). The cutting parameters were chosen according to those reported for AWJ machining of similar FRPs [34]. Five tensile specimens were prepared with each hole quality and hole size resulting in a total of 50 tensile specimens (5 Â 5 Â 2). Additional specimens with each hole quality and size were prepared for evaluating surface texture; these specimens were sectioned and examined but were not subjected to uniaxial loading. Two straight-sided tensile specimens without holes were also prepared and used to estimate the unnotched tensile strength of the laminate.

2Fadal VMC20 with 88HS CNC control. 3OMAX Model 2652 JetMachiningTM Center, Auburn WA.

(a)

Pitting

(b)

Axial Circumferential

Figure 2. Assessment of the hole quality: (a) typical features of a drilled hole; (b) surface profile orientations.

The hole quality obtained from each process was quantified in terms of standard surface roughness parameters and the effective stress concentration factor (KKt). Contact profilometry4 was used to analyze the surface texture using a skidless probe and 10 mm diameter diamond stylus. A representative specimen resulting from each method of preparation was sectioned prior to uniaxial loading; the damage and surface texture wereattributedsolelytothemethodofmachiningordrilling.Axialprofileswereobtainedfirst (parallel to the thrust axis) to identify the plies with the most drilling or machining damage. Multiple profiles were then obtained along damaged plies about the hole’s circumference. An example hole with ply specific damage is shown in Figure 2(a) and the definitions for axial and circumferential profiles are illustrated in Figure 2(b). In a single rotation twist drills encounter fiber orientations from 0 to 180 twice (Figure 3(a) and (b)). Fiber orientations from 90 to 180 are also commonly referred to as negative fiber orientations. Based on differences in material removal in cutting FRP components with positive and negative fiber orientations the machined surface characteristics are often described specifically with regard to fiber orientation [6,10,25,41]. The þ45 and –45 fiber orientations are important in discussing hole quality and are illustrated for the reader in Figure 3(c).

4A Hommel America T8000, Connecticut.

(a) fiber direction

(b) x fiber cutting edge orientation cutting direction

(c)

Tool Tool

+45˚ +135˚ y

-45˚ x

Figure 3. Drilling, fiber orientation and associated terminology: (a) axial view of drill penetrating a laminae. The drill’s rotation is indicated by the directions of arrows. Both of the individual cutting edges at the hole’s periphery represents a single point cutting tool; (b) magnified view of cutting edge and definition of fiber orientation. The fiber orientation is defined positive clockwise in this figure from 0 to 90; (c) þ45 and À45 fiber orientations.

A traverse length of 2.4 mm and cutoff length of 0.4 mm were used for all measurements. The average surface roughness (Ra), peak-to-valley height (Ry), and ten-point roughness (Rz) were calculated according to ANSI B46.1. A graphical radius gage was used to identify the radii of dominant profile valleys from profiles taken parallel to the circumference; dominant valleys were those with the maximum profile height variation and smallest valley radii (). The effective profile valley radius () for each hole was defined from an average of at least six valley radii. The surface roughness parameters and for each hole quality were used to estimate KKt according to Equation (2). An examination of the machined surfaces was also conducted using a Scanning Electron Microscope (SEM).5

Tension Tests and Stress Concentration Factors

The tensile specimens were prepared for loading by bonding G10 fiberglass end tabs to the grip surfaces using Hysol 9309 adhesive according to procedures outlined by Carlsson

5Jeol JSM, Model 5600.

P,d

MTS ε controller

extensometer V

AE computer hardware Pre-amplifier Digital wave conditioner

Figure 4. Schematic diagram of the hardware and data acquisition system. The axial force and displacement (P,d), strain (") and acoustic emission ( V) were recorded during each test. and Pipes [42]. Tensile tests were conducted in accordance with ASTM Standard D5766M- 95 using a universal test center6. The specimens were loaded to failure in load control actuator displacement at a rate of 445 N sÀ1 and the axial load, displacement and strain were acquired at 10 Hz. The acoustic emission corresponding to discrete failure events was monitored during tensile loading using a transducer and accompanying hardware7 at a rate of 500 Hz. Vacuum grease was used to couple the transducer and tensile specimens. A schematic diagram of the hardware and data acquisition system is shown in Figure 4. The load corresponding to first fiber failure of the specimens without holes was used to determine the unnotched tensile strength (o) of the Gr/Ep laminate. Similarly, the load and acoustic history of the specimens with holes were used to determine the notched tensile strength (N). The apparent stress concentration (KtðappÞ) of each specimen with drilled or machined hole was determined according to

o KtðappÞ ¼ ð5Þ N

The average first fiber failure strength (o) of the two straight-sided specimens without holes was 636.5 MPa. Using KtðappÞ and the stress concentration factor for a hole in a finite plate (Kt(hole)), the stress concentration for the drilled hole quality Kt(q) was extracted using the superposition principle according to

KtðappÞ KtðqÞ ¼ ð6Þ KtðholeÞ

6Material Test System (MTS) Model 810 load frame with Model 647 hydraulic wedge grips. 7Digital Wave F4012 Fracture Wave Detector, B1025 Broadband transducer, PA2040 G/A Broadband preamplifier.

Similarly, the ultimate tensile strength of the Gr/Ep laminate was determined from the maximum load experienced by the unnotched specimens prior to failure. The ultimate tensile strength of the specimens was determined in the same manner and the apparent stress concentration for the ultimate strength of the laminate with each hole quality was determined according to Equation (6). Note that Kt(hole) was estimated using design charts for holes in finite panels [40] and from a finite element model of the specimens according to the stress distribution in each ply. The panels were treated using classical lamination theory and first fiber failure was estimated using the Tsai–Hill criterion. Both the finite element analysis and design charts identified a stress concentration factor of 2.55 and 2.42 for the 6.35 and 9.53 mm holes, respectively.

RESULTS

A comparison of the specimens with drilled and AWJ machined holes indicated that the hole quality was process dependent. Typical axial profiles from a drilled and AWJ machined hole are shown in Figure 5(a) and (b), respectively. Damage resulting from fiber pullout and intra-ply delamination was clearly visible on the interior surface of the holes obtained with Drill A (Figure 5(a)). The damage appeared in four specific quadrants oriented 90 from one another and in 0.5 mm increments of hole depth. The periodicity in damage corresponded to the distance between plies with À45 fiber orientation relative to the tool path (Figure 3(c)) and the laminate’s stacking sequence. All of the 6.35 and 9.53 mm holes introduced using Drills A and C exhibited damage in locations with À45 fiber orientation relative to the tool path. In contrast, axial profiles from holes introduced

Axial (a) 50

-25

Surface Height, z (µm) -50 0.0 0.5 1.0 1.5 2.0 2.5 Traverse Length (mm)

Axial (b) 50

-25

Surface Height, z (µm) -50 0.0 0.5 1.0 1.5 2.0 2.5 Traverse Length (mm)

Figure 5. Typical axial surface profiles of 9.53 mm holes resulting from drilling and AWJ machining; (a) Drill A; (b) AWJ A.

600 600

400 400

200 200

0 0 CD A B -200 -200 A -400 B CD -400 Surface Height, z (µm) Surface Height, z (µm) -600 -600 0.0 0.4 0.8 1.2 1.6 2.0 2.4 0.0 0.4 0.8 1.2 1.6 2.0 2.4 (a)Traverse Length (mm) (b) Traverse Length (mm)

-200 100

-250 ρ=38 µm 50

-300 0 ρ=45 µm

-350 -50 Surface Height, z (µm) Surface Height, z (µm) -400 -100 1350 1400 1450 1500 1550 900 950 1000 1050 1100 (c)Traverse Length (µm) (d) Traverse Length (µm)

Figure 6. Circumferential surface profiles and selected valley radii of 9.53 mm holes resulting from drilling and AWJ machining: (a) Surface profile resulting from Drill A; (b) Surface profile resulting from AWJ A; (c) Valley (C) from profile in (a); (d) Valley (B) from profile in (b).

Table 3. Surface roughness parameters of the drilled and machined holes.

Average Roughness Peak to Valley Roughness Ten Point Height

Ra (mm) Ry (mm) Rz (mm) Method ¼ 6.35 ¼ 9.53 ¼ 6.35 ¼ 9.53 ¼ 6.35 ¼ 9.53 Drill A 2.0 4.1 16.1 28.1 12.9 24.3 Drill B 1.7 0.6 19.3 2.3 12.6 1.5 Drill C 1.6 0.9 14.5 14.4 8.8 5.2 AWJ A 5.9 7.0 39.3 44.4 31.9 33.3 AWJ B 3.6 3.4 27.6 23.7 20.0 19.1

Table 4. Effective profile valley radii and stress concentration factors of the holes.

Method ¼ 6.35 mm ¼ 9.53 mm (mm) KKt (mm) KKt Drill A 28 Æ 17 1.09 32 Æ 16 1.15 Drill B 22 Æ 7 1.12 1 1.00 Drill C 19 Æ 8 1.14 36 Æ 18 1.07 AWJ A 34 Æ 15 1.21 48 Æ 25 1.19 AWJ B 28 Æ 18 1.17 28 Æ 22 1.15

(Drill B) and was defined as infinity. In general, of the AWJ machined holes exhibited more variation than that resulting from drilling. Using the surface roughness parameters and profile valley radii for each hole quality, KKt was calculated according to Equation (2) (Table 4). An empirical factor of n ¼ 1 was used for all specimens for convenience. The highest KKt for both 6.35 and 9.53 mm holes resulted from AWJ machining (AWJ A), whereas the lowest KKt of the two hole sizes resulted from drilling. Holes obtained with Drill A resulted in the lowest KKt of the 6.35 mm diameter holes. And according to the minimal damage exhibited by the 9.53 mm holes introduced with Drill B, KKt of this group was the lowest of all holes (KKt ¼ 1). Due to the large range in , variation in KKt of the AWJ machined holes was greater than that of the drilled holes. Typical Acoustic Emission (AE) and load history records obtained while tensile testing of the Gr/Ep specimens with drilled (Drill B) and AWJ machined (AWJ B) holes are shown in Figure 7(a) and (b), respectively. For comparison, the load and acoustic emission records for a straight-sided specimen without notch is shown in Figure 7(c). Matrix cracking in the 90 plies, first fiber failure of 0 plies and ultimate failure are highlighted for clarity. Although the onset of matrix cracking could not be identified clearly using the AE records for most specimens, the onset of fiber failure was readily evident. The distribution in first fiber failure load with surface roughness for the 6.35 and 9.53 mm holes is shown in Figure 8(a) and (b), respectively. Specimens with 9.53 mm holes failed at lower loads than those with 6.35 mm holes, as expected, based on the larger net-section stress. The highest first fiber failure load resulted from testing of the AWJ machined specimens (AWJ B) whereas the specimens with the lowest load at fiber failure were those produced with the worn drill (Drill B). Utilizing the load at first fiber failure, the apparent stress concentration factor (KtðappÞ) for each group of specimens with holes was calculated according to Equation (5) and are listed in Table 5. The highest Kt(q) posed by the drilled

---- Force Acoustic (a) 40 1.5

first fiber (0¡ ply) 1.0 30 matrix cracking 0.5 (90¡ ply)

20 0.0

Force (kN) -0.5 10 ultimate failure -1.0 Acoustic Emission (V)

0 -1.5 0 0.001 0.002 0.003 0.004 0.005 0.006 0.007 Strain (m/m)

---- Force Acoustic (b) 40 1.5

1.0 30 matrix cracking 0.5 (90 ply) 20 0.0

Force (kN) -0.5 10 first fiber (0 ply)

-1.0 Acoustic Emission (V) ultimate failure 0 -1.5 0 0.001 0.002 0.003 0.004 0.005 0.006 0.007 Strain (m/m)

Figure 7. Typical force history and acoustic emission records of Gr/Ep specimens: (a) 9.53 mm drilled hole (Drill B); (b) 9.53 mm AWJ machined hole (AWJ B); (c) no hole.

Figure 8. Distribution in the first fiber failure load in terms of the average roughness of the hole: (a) 6.35 mm hole diameter; (b) 9.53 mm hole diameter.

Table 5. Apparent stress concentration factors for the holes estimated from experimental results of the tension tests.

First Fiber Failure Ultimate Failure

Method Kt(app) Kt(hole) Kt(q) Kt(app) ¼ 6.35 Drill A 2.81 2.55 1.10 1.75 Drill B 2.87 2.55 1.12 1.73 Drill C 2.81 2.55 1.10 1.74 AWJ A 2.82 2.55 1.10 1.73 AWJ B 2.57 2.55 1.01 1.71 Avg. ¼ 2.78 Avg. ¼ 1.73 ¼ 9.53 Drill A 2.62 2.42 1.08 1.69 Drill B 3.40 2.42 1.40 1.73 Drill C 2.61 2.42 1.08 1.72 AWJ A 2.64 2.42 1.09 1.76 AWJ B 2.50 2.42 1.03 1.74 Avg. ¼ 2.75 Avg. ¼ 1.73

Figure 9. Distribution in the load at ultimate failure in terms of the average roughness of the hole: (a) 6.35 mm hole diameter; (b) 9.53 mm hole diameter. hole surface texture resulted from holes introduced using the worn diamond coated drill (Drill B). The lowest Kt(q) for both hole sizes was obtained from AWJ machined specimens (AWJ B). Although hole quality appeared to influence the open hole tensile strength of the Gr/Ep laminate, there was no clear trend between the surface texture and first fiber failure load. Damage evolution continued after the onset of fiber failure and fracture of the specimens occurred at the ultimate load (Figure 7). The average and standard deviation of the ultimate load at failure was calculated for each group of specimens and is shown in terms of Ra for the 6.35 and 9.53 mm holes in Figure 9(a) and (b), respectively. Similar to the load at first fiber failure, Kt(app) was also calculated at ultimate failure and is shown in Table 5. There was minimal variation in Kt(app) for holes of similar size regardless of the hole quality. The average stress concentration factor at the ultimate load for all specimens was approximately 1.73, whereas the average Kt(app) at first fiber failure was 2.77 (Table 5).

DISCUSSION

The apparent stress concentration factors (Kt(app)) for the drilled and AWJ machined holes were estimated using results of tension tests. The superposition principle was then used to estimate the stress concentration factor associated with the hole quality (Kt(q))

Downloaded from http://jcm.sagepub.com at UNIV OF MARYLAND BALTIMORE CO on March 19, 2007 © 2003 SAGE Publications. All rights reserved. Not for commercial use or unauthorized distribution. 1454 D. AROLA AND M. L. MCCAIN according to Equation (6). Based on experimental results the mean Kt(app) at first fiber failure ranged from 2.50 to 3.40. Both the range in Kt(app) and its magnitude with respect to the stress concentration factor for open holes without edge effects suggests that hole quality influenced the tensile strength. It is important to emphasize that hole quality is comprised of both the surface texture and surface integrity, which describe the external and internal features of quality, respectively. The lowest stress concentration (2.50) was obtained from the AWJ machined holes with lowest Ra (AWJ B) as evident in Table 5. Holes introduced with the new diamond coated twist drill (Drill A) and tungsten carbide drill reamer (Drill C) exhibited the lowest Kt(app) of the drilled holes. However, the holes with lowest surface roughness (9.53 mm, Drill B) exhibited the highest Kt(app). Thus, there was no correlation between the surface texture and first fiber failure load of the Gr/Ep tensile specimens as evident from Figure 8. The apparent stress concentration factor for the drilled and AWJ machined specimens corresponding to the ultimate failure load ranged from 1.69 to 1.76. Uniformity in Kt(app) amongst the drilled and AWJ machined holes indicates that neither the surface texture or surface integrity influenced the ultimate strength of the Gr/Ep laminate tensile specimens (Figure 9). These results confirm reports of previous investigations indicating no correlation between the surface texture of open holes and ultimate strength of FRP laminates [17–19]. In addition to estimating Kt(q) from Kt(app) using superposition, the effective stress concentration factor (KKt) for the holes was estimated using the surface roughness parameters and effective profile valley radii according to Equation (2). A comparison of Kt(q) and KKt for specimens with 6.35 and 9.53 mm holes is shown in Figure 10. The average difference in KKt and mean value of Kt(q) for the 6.35 and 9.53 mm holes was 5.6 and 13.1%, respectively. The maximum difference in Kt(q) and KKt was 40.4% and occurred for the 9.53 mm holes introduced with the worn diamond coated drill (Drill B). Excluding the 9.53 mm holes introduced with the worn diamond coated tool (Drill B) the difference between Kt(q) and KKt for all specimens with drilled holes was less than 6%. The difference between KKt and Kt(q) for the AWJ machined holes was within 12%. Although KKt of the AWJ machined holes was always greater than Kt(q) it would not necessarily provide a

Figure 10. Comparison of the stress concentration estimated from the hole surface texture with experimental results.

Downloaded from http://jcm.sagepub.com at UNIV OF MARYLAND BALTIMORE CO on March 19, 2007 © 2003 SAGE Publications. All rights reserved. Not for commercial use or unauthorized distribution. Surface Texture and the Stress Concentration Factor for FRP Components 1455 reliable design allowance. Variations in first fiber failure load (Figure 8) identified using AE and the range in profile valley radii (Table 4) limits the significance in trend between Kt(q) and KKt evident in Figure 10. Previous studies on the influence of surface texture and strength of FRPs in flexure [33] have indicated that KKt could be used to account for the influence of machined edge effects on the mechanical behavior of FRPs. However, a comparison of the stress concentration factors estimated from experimental results with KKt indicates that it cannot fully describe the effects of hole quality on the open hole tensile strength of FRP components. Through a comparison between Kt(q) and KKt of the drilled and AWJ machined specimens it appears that the mechanical behavior of the Gr/Ep specimens was dependent on the method of machining. Drilling and AWJ machining of FRPs invoke different mechanisms of material removal [25,33,41] and undoubtedly resulted in a different surface integrity. An evaluation of the machined surface was conducted using a SEM to further understand the mechanical behavior and limited correlation between surface texture and open hole tensile strength. Micrographs from the surface of representative holes within the Gr/Ep laminate are shown in Figure 11. Figure 11(a) shows pitting of a ply with À45 fiber orientation and matrix smearing along the surface of 0, 45 and 90 plies from the surface of a 6.35 mm hole introduced using Drill A. Subsurface delamination and pitting was evident in all plies oriented À45 with respect to the cutting direction and resulted from fiber fracture beneath the tool nose. Edge trimming and drilling of FRPs materials with single point cutting tools often initiates damage in plies with À45 fiber orientation (Figure 3(c)). The comparatively low interfacial shear strength of FRPs results in delamination along the fiber–matrix interface and bending induced fiber fracture beneath the machined surface [6,11,25,41,43,44]. A micrograph from the surface of a 9.53 mm hole obtained with Drill B is shown in Figure 11(b). The epoxy matrix has been smeared over the entire drilled surface (all plies) and obscured details of the machined surface. Thus, the low roughness estimated from surface profiles (Table 3) was attributed to the smear layer. Features from a drilled hole introduced with Drill C are shown in Figure 11(c) and are similar to those resulting from the use of Drill A. Subsurface delamination is evident within the –45 ply similar to that highlighted in Figure 11(a). All holes prepared using Drills A and C exhibited subsurface damage in plies with À45 fiber orientation. In comparison, holes resulting from drilling the AWJ machined holes (Figure 9(d)) exhibited considerably different microscopic characteristics. Matrix smearing was not evident on any plies and the machined surface showed no distinct signs of microscopic damage. At high magnification the AWJ machined holes did not exhibit any disruption to the constituents. In addition, there was no apparent change to the fiber–matrix interface as shown in Figure 11(e) for a 90 ply. Despite the larger surface roughness the AWJ machined holes exhibited a higher surface integrity than the drilled holes and resulted in lower Kt(q) than expected from the surface roughness and KKt. The difference in surface integrity of the drilled and AWJ machined holes may have obscured the correlation between surface texture, KKt and first fiber failure load. It appears that the comparison of hole quality and tensile strength was complicated by differences in material removal in drilling and AWJ machining, as well as the discrepancies in surface integrity. It is important to note that KKt and Kt(q) estimated for the 6.35 and 9.53 mm holes introduced with the worn drill (Drill B) were not consistent. Differences in the response of specimens with 6.35 and 9.53 mm holes may have resulted from differences in wear introduced to the drills prior to the study; the wear induced to each drill was not quantified. Surface profiles from specimens obtained with Drill B were nearly perfect; no prominent

Figure 11. Micrographs from the surface of drilled and AWJ machined fastener holes in the Gr/Ep laminate. Arrows indicate the cutting direction. Note the subsurface delamination that is evident in (a) and (c) resulting from the use of the twist drills: (a) pitting on the surface of a À45 ply (Drill A, ¼ 6.35 mm) 270X; (b) matrix smearing covers machined surface (Drill B, ¼ 9.53 mm) 1000X; (c) subsurface delamination within a –45 ply (Drill C, ¼ 6.35 mm) 500X; (d) macroscopic view of AWJ machined surface (AWJ A, ¼ 6.35 mm) 160X; (e) high integrity of a 90 ply (AWJ A, ¼ 6.35 mm) 2000X.

valley radii could be detected and the average roughness (Ra) was 0.6 mm. Consequently, KKt for holes obtained with Drill B was 1.0 and suggested that there was no drilling related damage. However, the tensile strength of specimens with 9.53 mm holes introduced using Drill B resulted in Kt(q) of 1.40 (Table 5). Subsurface damage resulting from drilling with the worn tool must have been present beneath the redistributed epoxy matrix. An evaluation was conducted using the SEM to identify drilling damage that reduced the tensile strength of Drill B specimens with 9.53 mm holes. A representative specimen (9.53 mm, Drill B) not subjected to tensile loading was sectioned and examined and micrographs taken from the surface are shown in Figure 12. Discrete pockets without smeared matrix were found occasionally in regions of the hole with À45 fiber orientation (Figure 12(a)) and delamination pits were also identified at the intersection between the drilled hole and section

Figure 12. Micrographs from the surface of a drilled hole obtained with Drill B (9.53 mm): (a) discrete pockets in the smeared matrix and evidence of subsurface delamination resulting from drilling (x900); (b) a single subsurface pit within a ply with À45 fiber orientation that was exposed after sectioning (x1000). surface (e.g. Figure 12(b)). These features were not identified from the axial or circumferential surface profiles due to the smeared matrix (Figure 11(b)). Therefore, the difference in KKt and Kt(q) for 9.53 mm holes introduced with the worn drill was attributed to the limitations in contact profilometry and suggests that surface texture is not a reliable measure of drilled hole quality, especially for holes produced with worn cutting tools. Consequently, KKt may not provide an accurate measure of hole quality and corresponding stress concentration, especially for holes introduced with worn cutting tools.

CONCLUSIONS

The effects of drilled hole quality on the tensile strength of Fiber Reinforced Plastics (FRPs) was examined. Holes with nominal diameter of 6.35 and 9.53 mm were introduced in a graphite/epoxy (Gr/Ep) laminate using traditional cutting tools (twist drill and drill-reamer) and an Abrasive Waterjet (AWJ). The surface texture of

Downloaded from http://jcm.sagepub.com at UNIV OF MARYLAND BALTIMORE CO on March 19, 2007 © 2003 SAGE Publications. All rights reserved. Not for commercial use or unauthorized distribution. 1458 D. AROLA AND M. L. MCCAIN the holes was characterized using contact profilometry and the effective stress concentration factor (KKt) was estimated using the Arola–Ramulu model. Finally, the apparent stress concentration factor (KtðappÞ) was determined from experimental results of the tensile tests and was used to estimate the stress concentration posed by the hole quality (Kt(q)). 1. The apparent stress concentration factor for the drilled holes at first fiber failure ranged from 2.50 to 3.40. Using experimental results and the principle of superposition the stress concentration from the drilled hole damage (Kt(q)) was estimated and ranged from 1.0 to 1.40. There was no correlation between the hole surface texture and the open hole strength at first fiber failure. 2. The apparent stress concentration factor for the drilled holes at ultimate failure ranged from 1.69 to 1.76. There was no correlation between the hole quality (surface texture and surface integrity) and ultimate tensile strength of the Gr/Ep laminate. 3. When used to predict first fiber failure of FRPs with open holes, contact profilometry and standard surface roughness parameters may prove unreliable, especially in the evaluation of surfaces produced using worn cutting tools. Holes introduced using a worn diamond coated twist drill exhibited matrix smearing. Further evaluation using electron microscopy showed that smeared epoxy matrix obscured the surface topography and prevented detection of damage using surface profilometry.

ACKNOWLEDGMENTS

The authors gratefully acknowledge Paul Wienhold of the Applied Physics Laboratory of Johns Hopkins University for donation of materials and access to facilities. The authors are also thankful to the OMAX Corp. for partial support of the investigation through an equipment grant.

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