Composites: Part A 39 (2008) 1512–1521

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Composites: Part A

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Process simulation, design and manufacturing of a long fiber thermoplasticcomposite for mass transit application

K. Balaji Thattaiparthasarathy, S. Pillay, H. Ning, U.K. Vaidya *

Department of Materials Science and Engineering, The University of Alabama at Birmingham (UAB), Birmingham, Alabama 35294, USA

a r t i c l e i n f o a b s t r a c t

Article history:Received 29 October 2007Received in revised form 24 March 2008Accepted 26 May 2008

Keywords:A. Long fiber thermoplasticsE. Extrusion compression moldingC. Process modelingB. Thermoplastic composites

1359-835X/$ - see front matter � 2008 Elsevier Ltd.doi:10.1016/j.compositesa.2008.05.017

* Corresponding author. Tel.: +1 205 934 9199; faxE-mail address: uvaidya@uab.edu (U.K. Vaidya).

Long fiber thermoplastics (LFTs) have witnessed rapid growth in thermoplastics matrix composites,mainly due to developments in the automotive and transportation sector. In LFTs, pelletized thermoplas-tic polymer matrix is reinforced with long glass or carbon fibers (3–25 mm) are processed by extrusion-compression molding. The current work focuses on the applied science and manufacturing of E-glass/polypropylene (E-glass/PP) LFT composite material. Process simulation was conducted to evaluate theflow of fiber filled viscous charge during the compression molding of the LFT composite. Studies on opti-mum charge size and placement in the tool, press force, temperature of mold, shrinkage and warpagewere also conducted. The flow pattern of the molten charge in the mold and the resulting fiber orienta-tion predicted by process simulation are verified experimentally. The studies have been applied for amass transit/transportation component namely, a LFT battery box access door for form-fit-function toreplace a heavy metal door. Weight reduction of 60% was achieved using 40% weight percent E-glass/PP LFT over the metal design.

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1. Introduction

Long fiber thermoplastics (LFTs) are being used extensively inautomotive and transportation industry due to their superior spe-cific strength and modulus resulting in substantial weight savings,combined with relative ease of fabrication and handling [1].Weight reduction in a vehicle increases overall fuel efficiency,thereby reducing the operating costs and significantly contributingto environmental and economic benefits [2]. Global use of LFTs isexpected to grow from around 40 million lbs in 2001 to 75 mil-lion lbs in 2007 [3]. In general, some of the advantages of usingLFT over metals include high impact resistance, superior tough-ness, improved damping and corrosion resistance in conjunctionwith ease of shaping and recyclability [4,5]. The use of a thermo-plastic matrix provides the molder the ability to modify and en-hance the properties of the resin by blending additives, fillersand fire retardants depending on the nature of the application[6]. Various components have been designed and manufacturedusing LFTs for the transportation industry including, dashboardcarriers, front ends, seat shells, battery trays, spare wheel dwells,etc. [2,7,8]. The typical applications of LFT components in an auto-mobile are shown in Fig. 1 [8].

The mechanical properties of a part made of reinforced thermo-plastics are defined by the matrix system, type of fibers, fiber con-

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tent and orientation of the reinforcing fibers. The orientation andlength of the fibers are influenced by the processing method andprocess parameters. LFTs possess starting fiber lengths of 3–25 mm in contrast to short fiber thermoplastics (SFTs) compoundsthat possess 0.5 mm fiber length or less [9]. When processed opti-mally, LFTs possess a fiber length of 3–25 mm [2]. Hence the aver-age fiber lengths of LFTs are an order of magnitude greater than theSFTs. The full strength of the reinforcements is utilized because thefiber length is above the critical fiber length for effective loadtransfer [10]. The stiffness of the laminate is directly proportionalto the fiber concentration up to 40% by weight; and independentof fiber length above 0.5 mm [11]. Hence the use of long fibershas proven to increase the elastic modulus and the tensile strengthof the material as close as to 90% of that obtained when using con-tinuous fibers [12].

LFTs are manufactured by pulling continuous fiber towsthrough a thermoplastic polymer melt in a specialized processingdie. Early manufacturing attempts mimicked wire-coating technol-ogy, crosshead extrusion or several pultrusion techniques that didnot wet-out the individual fibers within the tow [13]. An alternatetechnique (Direct ReInforcement Fabrication Technology, DRIFT)[14], also referred to as hot-melt impregnation allows completeimpregnation of continuous fibers with thermoplastics polymersat very high production rates, providing a high-quality, low costthermoplastic composite. The hot-melt impregnation technologyenables to produce products in various forms such as continuousrods, tapes, pultruded shapes, or pellets of any length for injection

Fig. 1. Automobile components made of LFT [8].

K. Balaji Thattaiparthasarathy et al. / Composites: Part A 39 (2008) 1512–1521 1513

and/or compression molding. This manufacturing process can beused to combine a wide variety of thermoplastic resins and rein-forcing fibers. Fiber levels as high as 60% by weight are easilyproduced.

The starting materials for LFTs are pellets of average length 3–25 mm compared to the plate shaped semi-finished product ofglass mat thermoplastics (GMT). LFT pellets are fed into the hopperof a plasticator (a single screw, low shear extruder) where they aremetered down a barrel, heated above the melting point and ex-truded in low shear to form a molten charge. The molten chargeis extruded to a predetermined size, and shape (usually cylindrical)that is transferred to the compression molding press for the form-ing operation. Thermal process parameters and the velocity gradi-ent developed during the flow of the material influence the finalmechanical property of the molded part. To optimize the process-ing of LFT, it is necessary to take into account a number of processvariables and effects that are not typically encountered when pro-cessing unreinforced plastics. Molding pressures, fiber orienta-tions, fiber distribution, and flow fronts of the molten chargewithin the tool are important parameters which determine thehomogeneity of the molded product and therefore the success ofthe produced part. Recently developed software tools based onthe finite element method (FEM) help in the design for manufac-turing stage using these materials. By simulating the molding oper-ation in a virtual environment, the effect of process variables oncharge flow, mold filling, fiber orientation, shrinkage and warpagecan be determined and observed before the tool is cut, ensuring acomplete control of the process variables and being aware of itslimitations.

Several authors have implemented process models to verifyinjection molding of short fiber composites [15–21]. In the currentwork, a finite element (FE) simulation program for fiber filled poly-mer has been used to simulate the flow pattern, fiber orientationand process induced shrinkage/warpage of compression moldedgeometries [22]. It is during the filling stage that the flow inducedfiber orientation develops, upon which the final mechanical andthermo-mechanical properties of the part are highly dependent.The material properties can be broadly classified into the type ofanalysis as shown in Fig. 2.

The rheological behavior of the polymer and/or fiber filled resinunder the molding conditions were used to calculate the flow frontover the processing time. The flow front simulation predicts knitlines and entrapped air, the pressure and temperature distributionin the cavity and the clamping force. Because shrinkage and warpagehave a decisive influence on the dimensional stability of the moldedpart, pressure volume temperature (PVT) characteristics exhibitedby the material is also an input to the module. The mechanical anal-ysis data specified (Young’s modulus, Poisson’s ratio, and aspect ra-tio) was used along with predicted fiber orientation distribution tocalculate the final orthotropic material properties.

The present work considers an LFT composite with plate likeand ribbed features. The LFT composite is investigated in termsof flow, fiber distribution, fiber orientation and design validationstudies. The applied science studies have been extended to areal-world application namely a battery box access door for a masstransit bus, manufactured from a ribbed LFT 40 wt.% E-glass/PP (E-glass/PP) material.

2. Process modeling of extrusion compression molding of LFTs

The process simulation used in the present work comprises fourmodules namely – flow, heat transfer, fiber orientation, and shrink-age/warpage. A solid model of the LFT composite was generatedusing Pro/Engineer Wildfire. The solid model was imported intothe process simulation software after generating a three-noded fi-nite element mesh in Hypermesh�. The shell element is considereda 2.5 dimension membrane element with the thickness specified.For the simulation of compression molding, charge placement isdefined by selecting an area on the finite element mesh, which cor-responds to an extruded charge placed on a mold maintained at alower temperature than the charge. The flow analysis in compres-sion molding is modeled as non-Newtonian, under non-isothermalthree-dimensional cavities using finite elements. This technique,commonly called as control volume approach (CVA), requires thatthe three-dimensional molding surface be divided into flat shellelements. The cells or control volumes are generated by connectingthe element centroid with element mid-sides. When applying themass balance to each cell, the resulting equations are identical tothose arising from a Galerkin method for finite elements.

The influence of the effect of temperature on the local viscosityof the material is captured by the Carreau–Williams Landel Ferry(WLF) model. This model is used to capture the temperature anddeformation rate dependency of the viscosity [23] as given by:

g ¼ P1aT

ð1þ aT P2j _cjÞP3ð1Þ

where

_c corresponds to the shear rate,aT temperature shift coefficient, accounts for variation of viscos-ity at various temperature,P1 is the zero shear viscosity,P2 is a time constant,P3 is the exponent index.

Compression molding of LFT involves placing a heated charge ina cold mold. The material that comes into contact with the moldwalls is rapidly cooled; the local viscosity increases and the mate-rial in these regions will no longer flow. The filling stages of thecompression molding process are temperature dependant, the

Fig. 2. Material properties required for process modeling of LFTs.

1514 K. Balaji Thattaiparthasarathy et al. / Composites: Part A 39 (2008) 1512–1521

calculation of the temperature distribution is an integral step of theoverall simulation. The simplified form of the energy equations(2)–(4) used in the simulation of the heat transfer is as follows:

Conduction term: qcpoTot¼ k

o2Toz2 ð2Þ

Convection term: �qcp vxoToxþ vy

oToy

� �ð3Þ

Diffusion term: �sxyovx

oz� syz

ovy

ozð4Þ

where

q is the density;cp is the specific heat capacity;k is the thermal conduction coefficient;sxz; syz, shear stress is xz and yz plane, respectively;vx; vy; vz are the velocity component in x, y, and z, respectively.

The simulation of fiber orientation during the compressionmolding is essential to accurately predict the thermo-mechanicalbehavior and the final mechanical properties of the molded part.In general, the orientation of a particle, such as fiber is describedby two angles namely; in plane orientation and out of plane orien-tation angle. These angles change in time as the melt flows thor-ough a die. In general the angular orientations of the fibers arerepresented only by the in plane orientation angle. Folger–Tuckermodel [24] is used to capture the flow induced fiber orientation.The model adopts a statistical approximation that is applied tothe entire domain to predict the fiber orientation. The state of par-ticle orientation at a point is described by an orientation distribu-tion function, and is defined such that probability of a particlelocated at x, y at time t, being oriented between two angles is givenby Eqs. (5) and (6). Assuming the fiber density is homogenousthroughout, the continuity equation can be written as shown byEqs. (7)–(9). The fiber distribution model accounts for the fiber vol-

ume content, aspect ratio and a fiber interaction coefficient thatdepends on the number of fiber touches that occurred during theflow.

w ¼ f ð/; x; y; tÞ ð5Þ

Pð/1 < / < /2Þ ¼Z /2

/1

wð/; x; y; tÞd/ ð6Þ

owot¼ � o

o/ðw _/Þ ð7Þ

_/ ¼ �C1 _cw

owo/� cos / sin /

ovx

ox� sin2 /

ovx

oy

þ cos2 /ovy

oxþ sin / cos /

ovy

oyð8Þ

owot¼ �C1 _c

o2w

o2/� ow

o/cos / sin /

ovx

ox� sin2 /

ovx

oy

�

þ cos2 /ovy

oxþ sin / cos /

ovy

oy

�

� wo

o/cos / sin /

ovx

ox� sin2 /

ovx

oy

�

þ cos2 /ovy

oxþ sin / cos /

ovy

oy

�ð9Þ

where

w is the orientation distribution function,/1; /2 are the orientation angles,_c is the magnitude of the strain rate tensor,C1 is the phenomological coefficient which models the interac-tion between the fibers,vx; vy; vz are the velocity component in x, y, and z, respectively.

The anisotropic material properties resulting from the flowinduced fiber orientation can be determined. Combinations of mi-cro- and macromechanical theory are used to calculate the overall

K. Balaji Thattaiparthasarathy et al. / Composites: Part A 39 (2008) 1512–1521 1515

stiffness of the laminate. The Halpin–Tsai [25] micromechanicaltheory is used to calculate the anisotropic material property.

3. Process simulation results and validation

3.1. Flow fronts of molten charge

The flow patterns for four different charge locations and config-urations are shown in Fig. 3. Types 1–3 (Fig. 3a–c) have two smallcharges placed in different configuration, and these show the pres-ence of knit lines when the two molten charges are compressed in-side the mold. Hence, a charge parallel to the longer edge (Fig. 3d)was adopted to mold the LFT composite part.

For the mold to fill completely without any voids or prematurefreezing of the melt the approximate charge dimensions weredeemed to be 650 mm in length and 170 mm in diameter andthe force required to flow the molten charge inside the tool waspredicted to be approximately 350 metric tons. The top tool tem-perate was maintained at 80 �C and the bottom tool at 90 �C. Theflow front of the charge is seen to progress from the geometric cen-ter to the edges of the mold. The flow simulation shows that thefour corners of the mold fill at the very end of the molding process.A short shot of the LFT charge was used to verify the flow simula-tion result. A short shot consist of placing a smaller volume dosingof the mold than required to complete fill the cavity. The part pro-duced using a short shot provides information about the actualflow fronts developed inside the mold. Fig. 4 depicts the flow pat-tern comparison of the model which is 85% filled and a short shotof a charge respectively. In both the above cases the four cornersdid not fill and there is resemblance of the flow pattern predictedby the software.

3.2. Fiber orientation

The fiber orientation in each element is represented by a plot offiber distribution function and fiber angle. This plot represents thefiber orientation in an element with respect to the local XY plane ofthe element. The average fiber orientation distribution for five ele-ments is illustrated in Fig. 5.

Fig. 3. Flow patterns for different charge placement and configurations: (a)–(c) shows floflow patterns of one long charge placed horizontally.

The degree of orientation that occurred as the melt flowsthrough the cavity is predicted by a fiber orientation scale. Thisscale is derived for five layers through half the thickness, fromthe top surface to the mid-plane. The 180� angle is divided into25 sectors. For a randomly oriented layer there will be equal num-ber of fibers in each sector or direction. On the other hand, for apreferential orientation, the fibers will tend to align in one direc-tion and so most fibers will lie in just a few sectors. The value onthe scale that represents no orientation is derived by dividing 1by the 25 (total number of sectors), which yields the value 0.04.The more oriented the fibers become the less sectors, hence the fi-ber orientation scale value is greater than 0.04. Simulation resultsshow areas where the top surface has random orientation com-pared to a preferential orientation at the center.

X-ray radiographic studies were done to assess the fiber orien-tation of final molded part as shown in Fig. 6.

The molded part was radiographed using a tungsten target X-ray source at 40 kV. The part was placed between the X-ray sourceand an image intensifier connected to a charge coupled device(CCD) camera to obtain digital images. The images show a prefer-ential fiber orientation in selected areas. Although the imagesshow preferential orientation of fibers, it was difficult to determinethe orientation of fibers through the thickness of the molded part.Hence an alternate method to determine the orientation throughthe thickness namely high resolution computerized tomography(micro-CT) was used [26]. Using X-radiation as a penetratingprobe, the micro-CT affords detailed microstructural informationfrom almost any material.

To validate the fiber orientation predicted results, a representa-tive sample from the molded part was analyzed for through thethickness by using a Scanco lCT40 Micro-CT apparatus. Cross-sec-tional images were obtained at various depths to capture the orien-tation effect through the thickness. The images obtained fromMicro-CT were then analyzed using FiberScan�, advanced imageprocessing software that determines the fiber distribution as afunction of fiber orientation angle. Fig. 7 compares the fiber distri-bution plot obtained from simulation results with those generatedfrom micro-CT images. The predicted results are in accordancewith the micro-CT images showing the presence of a random

w patterns of two small charges placed in different orientations, and (d) shows the

Fig. 4. Flow front comparison of molten charge under compression molding: (a) short shot of charge compressed partially and (b) predicted flow front of charge.

Fig. 5. Fiber orientation distribution plots for adjacent elements.

1516 K. Balaji Thattaiparthasarathy et al. / Composites: Part A 39 (2008) 1512–1521

orientation on the face (skin) of the molded part and a preferentialorientation through the thickness. This is termed as the skin–coreeffect in the molding of thermoplastic parts.

During the solidification of the compression molded part, theresidual stresses continue to build. The molded part experiencesvarying temperatures and stages of solidification. The flow simula-tion, fiber orientation calculation and the influence of materialproperties are then used as input to evaluate the thermo-mechan-ical response (shrinkage and warpage) of the molded part at theend of the compression molding process. The effect of temperatureand pressure on thermal expansion is obtained from the pressurevolume temperature (PVT) data for matrix material. The effect offibers on thermal expansion coefficient is obtained by the combi-nation of the PVT data of the matrix and the micromechanical Hal-pin–Tsai model developed for unidirectional orientation. Fig. 8

shows a representative deformation after the part has been demol-ded and cooled to ambient temperature.

4. Design, and analysis of the LFT battery box door

As explained earlier, a mass transit part was designed and man-ufactured using 40 wt.% LFT E-glass/PP. A battery box access door(referred to as battery door) is an external part of the 20 m(60 ft) articulated mass transit bus (Fig. 9) which functions to pro-tect and house the several batteries needed for the regular opera-tion of the electrical systems of the bus. It is currently comprised ofan all steel sheet metal fascia which is bent to shape and thenwelded to a tubular steel frame which provides additional stiffnessto the part. The metallic battery door is approximately1 � 0.6 � 0.003 m and currently weights about 12 kg.

Fig. 6. Radiograph image of molded part.

Fig. 8. Deformation predicted by software after the part cools.

K. Balaji Thattaiparthasarathy et al. / Composites: Part A 39 (2008) 1512–1521 1517

A LFT battery door was redesigned using the current metallicbattery box door as the baseline. For the LFT battery door, the goalwas to replace the steel frame by integrating the features offeredby the frame into external face and therefore decreasing weightand assembly time and cost. In the integrated LFT battery door de-sign the stiffness offered by the steel frame is achieved by the useof ribs integrated with the face (Fig. 10). The ribs provide extradimensional stability to the part by controlling out of plane dis-placements caused by shrinkage and warpage of the parts as theyare withdrawn from the tool and cooled to ambient temperature.

The material selected for the LFT battery door had to meet var-ious criteria including, but not limited to: (a) equivalent stiffness tothat of 3 mm thick steel face sheet; (b) possess low weight andcost; (c) resist humidity and salt rich environments including bat-tery acids; (d) possess dimensional stability, (e) ease of processing,

Fig. 7. Fiber orientation for a representative sample obtained from Micro CT: (a) top surfaa preferential orientation, (c) fiber distribution graph obtained from modeling showinthickness, and (d) representative fiber frequency plot obtained from Micro CT images u

and (f) paintable surface. Based on these requirements, long glassfiber reinforced polypropylene, E-glass/PP Celstran� PP-GF40-03(40% fiber weight fraction, 25 mm long) produced by Ticona Inc.was selected. The mechanical properties 40% fiber weight fractionof E-glass/PP LFT (25.4 mm long) are listed and classified in Table 1.

4.1. Finite element analysis of the battery door

Finite element analysis (FEA) of the LFT battery door was con-ducted using ANSYS�. The boundary conditions for the model were

ce showing a random orientation, (b) mid-plane through half the thickness showingg random orientation on top surface, with a preferential orientation through thesing FiberScan�.

Fig. 9. Detail of Battery Box Door on the 60 BRT Bus Model.

Table 1Mechanical properties of E-Glass/PP Celstran� PP-GF40-03 (40% fiber weight fraction)[28]

Property Value Units

Tensile modulus (1 mm/min) 7900 MPaTensile stress at break (5 mm/min) 100 MPaTensile strain at break (5 mm/min) 2 %Flexural modulus (23 �C) 8000 MPaFlexural strength (23 �C) 175 MPaCharpy notched impact strength (23 �C) 20 kJ/m2

Density 1210 kg/m3

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selected based upon the assembly of the part to the exterior shellof a bus. Static FEA simulations were conducted for three differentthicknesses (3, 4, and 5 mm) of the door shell using the propertiesof the LFT (PP-GF40-03) material. Fig. 11 shows a typical von Misesstress plot and the maximum deflection of the door for a shellthickness of 5 mm.

The von Mises stress plot illustrates the stress concentration inthe region of loading and the stress profile in the region of the stiff-ening ribs. The mid-span deflection for the 3 mm shell thicknessLFT battery box door was excessive, and hence failed based on fail-ure criteria (FC = 2.4, where a value of FC exceeding 1.0 denotesfailure). A similar scenario was observed with the 4 mm shellthickness case (FC = 1.7). With the 5 mm thick version, the safetyfactor is close to the strength of the steel counterpart, and thedeflection (19.2 mm) is not significant. A summary of the resultsfrom the FEA and the solid model is provided in Table 2. The weightbetween the steel frame and the LFT battery door design for 5 mmthickness shell is compared in Table 3. The percentage weight sav-ings on the final LFT molded design was calculated to be approxi-mately 60% compared to the steel frame design.

Fig. 10. (a) Single component design of the battery box access door with r

5. Processing and component verification

5.1. Process verification

An oil heated/cooled two side matched steel tool was selectedas the prototype/production tool. The tool has two main parts, atop tool which is a solid steel block that can be heated to the re-quired molding temperature and has the required machining andsurface finish to provide a class-A finish to the fascia (exterior) ofthe produced battery door and a bottom tool, which includes allthe machined cavities to generate the ribbed structure on the backof the door and includes the detail to accommodate the lock hous-ing and door handles. Fig. 12 shows the top and bottom toolmounted on the press with an extruded charge placed on the bot-tom tool. The tool was placed in a 400 metric ton press. A compar-ison of the process variables predicted by process modeling andactual values are tabulated in Table 4.

5.2. Design verification by mechanical testing

The molded LFT battery door was mechanically tested to verifythe stiffness predicted by the FEA model. The displacement pre-dicted by the FEA model of the battery door was verified experi-mentally with the same set of boundary conditions. A test framewas fabricated, and a hydraulic jack was used to apply a load of5500 N. The boundary conditions applied during testing were de-signed to replicate those seen by the door once mounted on thebus and during service. The load was applied from the exterior sur-face at the geometric center and the corresponding displacementon the transverse direction was measured using an electronic dialgauge that was placed in the geometric center on the interior.Fig. 13 shows the FEA model with the applied boundary conditions

ib stiffened structure and (b) view of the door from the cosmetic side.

Fig. 11. (a) Shows the Von Mises stresses (MPa) and (b) shows the displacement (mm) plots of 5 mm thick Battery Box Access Door subjected to 2223 N in the center.

Table 2FEA for a load of 2223 N (500 lbf) for various shell thickness

Shell thickness(mm)

Displacement(mm)

Von Misesstress (MPa)

Maximum stresscriterion (MPa)

Calculatedmass (kg)

3 44.70 224 2.40 2.104 28.19 156 1.70 2.805 19.20 99 1.00 3.50

Table 3Comparison of weights between steel frame design and LFT design (5 mm thick shell)

Physical property Steel door LFT door

Face sheet Tubular frame 5 mm shell design

Volume (m3) 9.40E�04 5.80E�04 3.14E�03Density (kg/m3) 7.86E+03 7.86E+03 1.21E+03Mass (kg) 7.39 4.57 3.51Total mass (kg) 11.96 3.51

Table 4Processing parameters for extrusion compression molding of E-Glass/PP

Process parameter Predicted value Actual value Units

Mold closing velocity 25 25 mm/sMold temperature (top tool) 80 80 �CMold temperature (bottom tool) 90 130 �CMaximum press force 3500 3900 kNCharge length 650 650 mmCharge diameter 170 170 mmCooling time 120 180 S

K. Balaji Thattaiparthasarathy et al. / Composites: Part A 39 (2008) 1512–1521 1519

and the experimental set-up used to compare the predicted dis-placement. Load vs. displacement data obtained from both theFEA model and experimentally are compared in Fig. 14.

The stiffness of the door as predicted by the FEA is consistentwith the experimental stiffness until a deflection of 15 mm (78%of failure deflection). At this point the stiffness response transitionsfrom a linear to a nonlinear state. The onset of nonlinearity can beattributed to the local plasticity effects and/or damage initiation.

Fig. 12. Top and Bottom tool clamped on the press with an extruded charge.

The FEA results show a maximum deflection at the point of loadingin the mid-span and a stress concentration in the vertical rib in thesurrounding area. The mechanical tests conducted on the batterydoor show similar failure like the FEA occurring on the verticalrib (Fig. 15). FEA predicted stress levels in excess of 155 MPa inthe region which corresponds to the flexural strength reportedfor the E-glass/PP LFT material.

5.3. Fiber weight fractions and fiber length verification

Fiber distribution and orientation in substructures, such as ribsor bosses, change and there by the expected stiffness of productscannot be obtained. These substructures also lead to fiber matrixseparation in SMC molded parts [27]. Fiber–matrix separationleads to ribs with poor fiber content and resin-rich edges in largeparts, weakening the structural integrity of the product. To validatethe fiber distribution (fiber weight fractions) in the final moldedpart, representative samples from various sections of the door(Ribs and Skin) were sectioned and subjected to burn-off to sepa-rate the fiber from the matrix. The results of the burn-off studyare tabulated in Table 5. The results show that there is no signifi-cant fiber matrix separation and the overall fiber weight fractionremained to be constant. Unlike SMC the LFT seem to flow throughribs of narrow width without any fiber matrix separation, there bythe retaining the structural integrity.

Fiber lengths were determined by image analysis and opticalmicroscopy on fiber samples removed from the molded part afterhigh temperature ashing. The fiber lengths of 600 individual fiberswere measured from the molded part to determine the fiber lengthdistribution. The fibers were dispersed in an aluminum pan and astereoscope was used to capture several images. Post-processingof the dispersed fibers was done using the software Image Pro-Plus.Fig. 16 shows the fibers separated from the matrix and dispersedfor length analysis. The fiber length distribution (Fig. 17) showsthat the majority of the fibers are greater in length than the criticalfiber length necessary for effective load transfer. The average fiberlength for 600 fibers was 9.54 mm and approximately 80% of thefibers were greater than the critical fiber length of 3 mm.

Fig. 13. (a) FEA model with applied boundary conditions and (b) experimental set-up to measure load vs. displacement on the molded LFT battery door.

Fig. 14. Load vs. displacement obtained from FEA and experiment.

Table 5Results of fiber weight fraction at various regions of molded battery door

Sample ID Length(mm)

Width(mm)

Height(mm)

Fiber weightfraction

Rib – 1 25.75 4.50 14.50 40.15Rib – 2 26.25 4.50 14.50 39.80Rib – 3 26.30 4.50 14.50 41.04Rib – 4 29.25 9.00 19.88 38.28Rib – 5 27.50 9.00 19.88 38.42Skin center – 1 22.50 22.50 5.00 39.25Skin center – 2 22.50 22.50 5.00 39.93Skin corner – 1 22.50 22.50 5.00 40.28Skin corner – 2 22.50 22.50 5.00 39.23

1520 K. Balaji Thattaiparthasarathy et al. / Composites: Part A 39 (2008) 1512–1521

6. Summary and conclusions

A LFT bus battery door was designed and fabricated using theextrusion–compression molding process. The LFT design incorpo-rated ribs to the shell to add stiffness and rigidity in a single

Fig. 15. (a) Failure initiation on the ribs during mechanical testing and (b) FEA sho

component as opposed to the steel frame which was welded tothe face sheet. The LFT design offered an approximate weightsavings of 60%. The extrusion–compression molding process wassimulated using the Cadpress-Thermoplastics. The flow fronts ofthe molten charge inside the tool and the fiber orientation inthe final molded part were verified experimentally. Quantitativeverification was performed to compare the stiffness predictedby the FEA model and the actual molded part. Fiber distributionwas uniform throughout the molded area and there was no sig-nificant fiber matrix separation in narrow ribs. Fiber length anal-ysis show minimum fiber degradation favoring effective loadtransfer.

wing stress concentration (von Mises stress) on the ribs at the same location.

Fig. 16. Fibers separated from resin and dispersed for fiber length analysis.

Fig. 17. Fiber length distribution plot.

K. Balaji Thattaiparthasarathy et al. / Composites: Part A 39 (2008) 1512–1521 1521

Acknowledgements

The authors gratefully acknowledge the support provided bythe Federal Transit Administration (FTA), Department of Transpor-tation, Project No. AL-26-7002 and Program Manager Terrell Wil-liams of FTA. Technical help received from Juan Serrano andGeorge Husman is also gratefully acknowledged.

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