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Process simulation, design and manufacturing of a long fiber thermoplastic composite for mass transit application
Composites: Part A 39 (2008) 1512–1521
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Composites: Part A journal homepage: www.elsevier.com/locate/compositesa
Process simulation, design and manufacturing of a long fiber thermoplastic composite 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
Article history: Received 29 October 2007 Received in revised form 24 March 2008 Accepted 26 May 2008
Keywords: A. Long fiber thermoplastics E. Extrusion compression molding C. Process modeling B. Thermoplastic composites
a b s t r a c t 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 thermoplastic polymer matrix is reinforced with long glass or carbon fibers (3–25 mm) are processed by extrusioncompression 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 the flow of fiber filled viscous charge during the compression molding of the LFT composite. Studies on optimum charge size and placement in the tool, press force, temperature of mold, shrinkage and warpage were also conducted. The flow pattern of the molten charge in the mold and the resulting fiber orientation predicted by process simulation are verified experimentally. The studies have been applied for a mass transit/transportation component namely, a LFT battery box access door for form-fit-function to replace a heavy metal door. Weight reduction of 60% was achieved using 40% weight percent E-glass/ PP LFT over the metal design. Ó 2008 Elsevier Ltd. All rights reserved.
1. Introduction Long fiber thermoplastics (LFTs) are being used extensively in automotive and transportation industry due to their superior specific 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 contributing to environmental and economic benefits [2]. Global use of LFTs is expected to grow from around 40 million lbs in 2001 to 75 million lbs in 2007 [3]. In general, some of the advantages of using LFT over metals include high impact resistance, superior toughness, improved damping and corrosion resistance in conjunction with ease of shaping and recyclability [4,5]. The use of a thermoplastic matrix provides the molder the ability to modify and enhance the properties of the resin by blending additives, fillers and fire retardants depending on the nature of the application [6]. Various components have been designed and manufactured using LFTs for the transportation industry including, dashboard carriers, front ends, seat shells, battery trays, spare wheel dwells, etc. [2,7,8]. The typical applications of LFT components in an automobile are shown in Fig. 1 [8]. The mechanical properties of a part made of reinforced thermoplastics are defined by the matrix system, type of fibers, fiber con* Corresponding author. Tel.: +1 205 934 9199; fax: +1 205 934 8485. E-mail address: [email protected] (U.K. Vaidya). 1359-835X/$ - see front matter Ó 2008 Elsevier Ltd. All rights reserved. doi:10.1016/j.compositesa.2008.05.017
tent and orientation of the reinforcing fibers. The orientation and length of the fibers are influenced by the processing method and process parameters. LFTs possess starting fiber lengths of 3– 25 mm in contrast to short fiber thermoplastics (SFTs) compounds that possess 0.5 mm fiber length or less [9]. When processed optimally, LFTs possess a fiber length of 3–25 mm [2]. Hence the average fiber lengths of LFTs are an order of magnitude greater than the SFTs. The full strength of the reinforcements is utilized because the fiber length is above the critical fiber length for effective load transfer [10]. The stiffness of the laminate is directly proportional to the fiber concentration up to 40% by weight; and independent of fiber length above 0.5 mm [11]. Hence the use of long fibers has proven to increase the elastic modulus and the tensile strength of the material as close as to 90% of that obtained when using continuous fibers [12]. LFTs are manufactured by pulling continuous fiber tows through a thermoplastic polymer melt in a specialized processing die. Early manufacturing attempts mimicked wire-coating technology, crosshead extrusion or several pultrusion techniques that did not wet-out the individual fibers within the tow [13]. An alternate technique (Direct ReInforcement Fabrication Technology, DRIFT) [14], also referred to as hot-melt impregnation allows complete impregnation of continuous fibers with thermoplastics polymers at very high production rates, providing a high-quality, low cost thermoplastic composite. The hot-melt impregnation technology enables to produce products in various forms such as continuous rods, tapes, pultruded shapes, or pellets of any length for injection
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Fig. 1. Automobile components made of LFT [8].
and/or compression molding. This manufacturing process can be used to combine a wide variety of thermoplastic resins and reinforcing fibers. Fiber levels as high as 60% by weight are easily produced. The starting materials for LFTs are pellets of average length 3– 25 mm compared to the plate shaped semi-finished product of glass mat thermoplastics (GMT). LFT pellets are fed into the hopper of a plasticator (a single screw, low shear extruder) where they are metered down a barrel, heated above the melting point and extruded in low shear to form a molten charge. The molten charge is extruded to a predetermined size, and shape (usually cylindrical) that is transferred to the compression molding press for the forming operation. Thermal process parameters and the velocity gradient developed during the flow of the material influence the final mechanical property of the molded part. To optimize the processing of LFT, it is necessary to take into account a number of process variables and effects that are not typically encountered when processing unreinforced plastics. Molding pressures, fiber orientations, fiber distribution, and flow fronts of the molten charge within the tool are important parameters which determine the homogeneity of the molded product and therefore the success of the produced part. Recently developed software tools based on the finite element method (FEM) help in the design for manufacturing stage using these materials. By simulating the molding operation in a virtual environment, the effect of process variables on charge flow, mold filling, fiber orientation, shrinkage and warpage can be determined and observed before the tool is cut, ensuring a complete control of the process variables and being aware of its limitations. Several authors have implemented process models to verify injection molding of short fiber composites [15–21]. In the current work, a finite element (FE) simulation program for fiber filled polymer has been used to simulate the flow pattern, fiber orientation and process induced shrinkage/warpage of compression molded geometries [22]. It is during the filling stage that the flow induced fiber orientation develops, upon which the final mechanical and thermo-mechanical properties of the part are highly dependent. The material properties can be broadly classified into the type of analysis as shown in Fig. 2. The rheological behavior of the polymer and/or fiber filled resin under the molding conditions were used to calculate the flow front over the processing time. The flow front simulation predicts knit lines and entrapped air, the pressure and temperature distribution in the cavity and the clamping force. Because shrinkage and warpage have a decisive influence on the dimensional stability of the molded part, pressure volume temperature (PVT) characteristics exhibited by the material is also an input to the module. The mechanical analysis data specified (Young’s modulus, Poisson’s ratio, and aspect ratio) was used along with predicted fiber orientation distribution to calculate the final orthotropic material properties.
The present work considers an LFT composite with plate like and ribbed features. The LFT composite is investigated in terms of flow, fiber distribution, fiber orientation and design validation studies. The applied science studies have been extended to a real-world application namely a battery box access door for a mass transit bus, manufactured from a ribbed LFT 40 wt.% E-glass/PP (Eglass/PP) material. 2. Process modeling of extrusion compression molding of LFTs The process simulation used in the present work comprises four modules namely – flow, heat transfer, fiber orientation, and shrinkage/warpage. A solid model of the LFT composite was generated using Pro/Engineer Wildfire. The solid model was imported into the process simulation software after generating a three-noded finite element mesh in HypermeshÒ. The shell element is considered a 2.5 dimension membrane element with the thickness specified. For the simulation of compression molding, charge placement is defined by selecting an area on the finite element mesh, which corresponds to an extruded charge placed on a mold maintained at a lower temperature than the charge. The flow analysis in compression molding is modeled as non-Newtonian, under non-isothermal three-dimensional cavities using finite elements. This technique, commonly called as control volume approach (CVA), requires that the three-dimensional molding surface be divided into flat shell elements. The cells or control volumes are generated by connecting the element centroid with element mid-sides. When applying the mass balance to each cell, the resulting equations are identical to those arising from a Galerkin method for finite elements. The influence of the effect of temperature on the local viscosity of the material is captured by the Carreau–Williams Landel Ferry (WLF) model. This model is used to capture the temperature and deformation rate dependency of the viscosity [23] as given by:
g¼
P1 aT ð1 þ aT P2 jc_ jÞP3
ð1Þ
where
c_ corresponds to the shear rate, aT temperature shift coefficient, accounts for variation of viscosity 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 in a cold mold. The material that comes into contact with the mold walls is rapidly cooled; the local viscosity increases and the material in these regions will no longer flow. The filling stages of the compression molding process are temperature dependant, the
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Fig. 2. Material properties required for process modeling of LFTs.
calculation of the temperature distribution is an integral step of the overall simulation. The simplified form of the energy equations (2)–(4) used in the simulation of the heat transfer is as follows:
ume content, aspect ratio and a fiber interaction coefficient that depends on the number of fiber touches that occurred during the flow.
oT o2 T ¼k 2 ot oz oT oT Convection term: qcp vx þ vy ox oy ovx ovy Diffusion term: sxy syz oz oz
w ¼ f ð/; x; y; tÞ
Conduction term: qcp
ð2Þ ð3Þ ð4Þ
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 compression molding is essential to accurately predict the thermo-mechanical behavior and the final mechanical properties of the molded part. In general, the orientation of a particle, such as fiber is described by two angles namely; in plane orientation and out of plane orientation angle. These angles change in time as the melt flows thorough a die. In general the angular orientations of the fibers are represented only by the in plane orientation angle. Folger–Tucker model [24] is used to capture the flow induced fiber orientation. The model adopts a statistical approximation that is applied to the entire domain to predict the fiber orientation. The state of particle orientation at a point is described by an orientation distribution function, and is defined such that probability of a particle located at x, y at time t, being oriented between two angles is given by Eqs. (5) and (6). Assuming the fiber density is homogenous throughout, the continuity equation can be written as shown by Eqs. (7)–(9). The fiber distribution model accounts for the fiber vol-
Pð/1 < / < /2 Þ ¼
Z
ð5Þ /2
wð/; x; y; tÞ d/
ð6Þ
/1
ow o _ ¼ ðw/Þ ot o/ C 1 c_ ow ovx ovx 2 /_ ¼ cos / sin / sin / w o/ ox oy ov ov y y þ cos2 / þ sin / cos / ox oy ow o2 w ow ovx ovx 2 ¼ C 1 c_ 2 cos / sin / sin / ot ox oy o / o/ ov ov y y þ cos2 / þ sin / cos / ox oy o ovx ovx 2 cos / sin / sin / w o/ ox oy ovy ovy 2 þ cos / þ sin / cos / ox oy
ð7Þ
ð8Þ
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 interaction between the fibers, vx; vy; vz are the velocity component in x, y, and z, respectively. The anisotropic material properties resulting from the flow induced fiber orientation can be determined. Combinations of micro- and macromechanical theory are used to calculate the overall
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stiffness of the laminate. The Halpin–Tsai [25] micromechanical theory 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 configurations are shown in Fig. 3. Types 1–3 (Fig. 3a–c) have two small charges placed in different configuration, and these show the presence of knit lines when the two molten charges are compressed inside 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 premature freezing of the melt the approximate charge dimensions were deemed to be 650 mm in length and 170 mm in diameter and the force required to flow the molten charge inside the tool was predicted to be approximately 350 metric tons. The top tool temperate was maintained at 80 °C and the bottom tool at 90 °C. The flow front of the charge is seen to progress from the geometric center to the edges of the mold. The flow simulation shows that the four 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 simulation result. A short shot consist of placing a smaller volume dosing of the mold than required to complete fill the cavity. The part produced using a short shot provides information about the actual flow fronts developed inside the mold. Fig. 4 depicts the flow pattern comparison of the model which is 85% filled and a short shot of a charge respectively. In both the above cases the four corners did not fill and there is resemblance of the flow pattern predicted by the software. 3.2. Fiber orientation The fiber orientation in each element is represented by a plot of fiber distribution function and fiber angle. This plot represents the fiber orientation in an element with respect to the local XY plane of the element. The average fiber orientation distribution for five elements is illustrated in Fig. 5.
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The degree of orientation that occurred as the melt flows through the cavity is predicted by a fiber orientation scale. This scale is derived for five layers through half the thickness, from the top surface to the mid-plane. The 180° angle is divided into 25 sectors. For a randomly oriented layer there will be equal number of fibers in each sector or direction. On the other hand, for a preferential orientation, the fibers will tend to align in one direction and so most fibers will lie in just a few sectors. The value on the scale that represents no orientation is derived by dividing 1 by the 25 (total number of sectors), which yields the value 0.04. The more oriented the fibers become the less sectors, hence the fiber orientation scale value is greater than 0.04. Simulation results show areas where the top surface has random orientation compared to a preferential orientation at the center. X-ray radiographic studies were done to assess the fiber orientation of final molded part as shown in Fig. 6. The molded part was radiographed using a tungsten target Xray source at 40 kV. The part was placed between the X-ray source and an image intensifier connected to a charge coupled device (CCD) camera to obtain digital images. The images show a preferential fiber orientation in selected areas. Although the images show preferential orientation of fibers, it was difficult to determine the orientation of fibers through the thickness of the molded part. Hence an alternate method to determine the orientation through the thickness namely high resolution computerized tomography (micro-CT) was used [26]. Using X-radiation as a penetrating probe, the micro-CT affords detailed microstructural information from almost any material. To validate the fiber orientation predicted results, a representative sample from the molded part was analyzed for through the thickness by using a Scanco lCT40 Micro-CT apparatus. Cross-sectional images were obtained at various depths to capture the orientation effect through the thickness. The images obtained from Micro-CT were then analyzed using FiberScanÒ, advanced image processing software that determines the fiber distribution as a function of fiber orientation angle. Fig. 7 compares the fiber distribution plot obtained from simulation results with those generated from micro-CT images. The predicted results are in accordance with the micro-CT images showing the presence of a random
Fig. 3. Flow patterns for different charge placement and configurations: (a)–(c) shows flow patterns of two small charges placed in different orientations, and (d) shows the flow patterns of one long charge placed horizontally.
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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.
orientation on the face (skin) of the molded part and a preferential orientation through the thickness. This is termed as the skin–core effect in the molding of thermoplastic parts. During the solidification of the compression molded part, the residual stresses continue to build. The molded part experiences varying temperatures and stages of solidification. The flow simulation, fiber orientation calculation and the influence of material properties are then used as input to evaluate the thermo-mechanical response (shrinkage and warpage) of the molded part at the end of the compression molding process. The effect of temperature and pressure on thermal expansion is obtained from the pressure volume temperature (PVT) data for matrix material. The effect of fibers on thermal expansion coefficient is obtained by the combination of the PVT data of the matrix and the micromechanical Halpin–Tsai model developed for unidirectional orientation. Fig. 8
shows a representative deformation after the part has been demolded 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 manufactured 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 protect and house the several batteries needed for the regular operation of the electrical systems of the bus. It is currently comprised of an all steel sheet metal fascia which is bent to shape and then welded to a tubular steel frame which provides additional stiffness to the part. The metallic battery door is approximately 1 0.6 0.003 m and currently weights about 12 kg.
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Fig. 6. Radiograph image of molded part.
A LFT battery door was redesigned using the current metallic battery box door as the baseline. For the LFT battery door, the goal was to replace the steel frame by integrating the features offered by the frame into external face and therefore decreasing weight and assembly time and cost. In the integrated LFT battery door design the stiffness offered by the steel frame is achieved by the use of ribs integrated with the face (Fig. 10). The ribs provide extra dimensional stability to the part by controlling out of plane displacements caused by shrinkage and warpage of the parts as they are withdrawn from the tool and cooled to ambient temperature. The material selected for the LFT battery door had to meet various criteria including, but not limited to: (a) equivalent stiffness to that of 3 mm thick steel face sheet; (b) possess low weight and cost; (c) resist humidity and salt rich environments including battery acids; (d) possess dimensional stability, (e) ease of processing,
Fig. 8. Deformation predicted by software after the part cools.
and (f) paintable surface. Based on these requirements, long glass fiber reinforced polypropylene, E-glass/PP CelstranÒ PP-GF40-03 (40% fiber weight fraction, 25 mm long) produced by Ticona Inc. was selected. The mechanical properties 40% fiber weight fraction of 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 conducted using ANSYSÒ. The boundary conditions for the model were
Fig. 7. Fiber orientation for a representative sample obtained from Micro CT: (a) top surface showing a random orientation, (b) mid-plane through half the thickness showing a preferential orientation, (c) fiber distribution graph obtained from modeling showing random orientation on top surface, with a preferential orientation through the thickness, and (d) representative fiber frequency plot obtained from Micro CT images using FiberScanÒ.
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K. Balaji Thattaiparthasarathy et al. / Composites: Part A 39 (2008) 1512–1521 Table 1 Mechanical properties of E-Glass/PP CelstranÒ PP-GF40-03 (40% fiber weight fraction) [28] Property
Value
Units
Tensile modulus (1 mm/min) Tensile stress at break (5 mm/min) Tensile strain at break (5 mm/min) Flexural modulus (23 °C) Flexural strength (23 °C) Charpy notched impact strength (23 °C) Density
7900 100 2 8000 175 20 1210
MPa MPa % MPa MPa kJ/m2 kg/m3
5. Processing and component verification 5.1. Process verification
Fig. 9. Detail of Battery Box Door on the 60 BRT Bus Model.
selected based upon the assembly of the part to the exterior shell of a bus. Static FEA simulations were conducted for three different thicknesses (3, 4, and 5 mm) of the door shell using the properties of the LFT (PP-GF40-03) material. Fig. 11 shows a typical von Mises stress plot and the maximum deflection of the door for a shell thickness of 5 mm. The von Mises stress plot illustrates the stress concentration in the region of loading and the stress profile in the region of the stiffening ribs. The mid-span deflection for the 3 mm shell thickness LFT battery box door was excessive, and hence failed based on failure criteria (FC = 2.4, where a value of FC exceeding 1.0 denotes failure). A similar scenario was observed with the 4 mm shell thickness case (FC = 1.7). With the 5 mm thick version, the safety factor is close to the strength of the steel counterpart, and the deflection (19.2 mm) is not significant. A summary of the results from the FEA and the solid model is provided in Table 2. The weight between the steel frame and the LFT battery door design for 5 mm thickness shell is compared in Table 3. The percentage weight savings on the final LFT molded design was calculated to be approximately 60% compared to the steel frame design.
An oil heated/cooled two side matched steel tool was selected as the prototype/production tool. The tool has two main parts, a top tool which is a solid steel block that can be heated to the required molding temperature and has the required machining and surface finish to provide a class-A finish to the fascia (exterior) of the produced battery door and a bottom tool, which includes all the machined cavities to generate the ribbed structure on the back of the door and includes the detail to accommodate the lock housing and door handles. Fig. 12 shows the top and bottom tool mounted on the press with an extruded charge placed on the bottom tool. The tool was placed in a 400 metric ton press. A comparison of the process variables predicted by process modeling and actual values are tabulated in Table 4. 5.2. Design verification by mechanical testing The molded LFT battery door was mechanically tested to verify the stiffness predicted by the FEA model. The displacement predicted by the FEA model of the battery door was verified experimentally with the same set of boundary conditions. A test frame was fabricated, and a hydraulic jack was used to apply a load of 5500 N. The boundary conditions applied during testing were designed to replicate those seen by the door once mounted on the bus and during service. The load was applied from the exterior surface at the geometric center and the corresponding displacement on the transverse direction was measured using an electronic dial gauge that was placed in the geometric center on the interior. Fig. 13 shows the FEA model with the applied boundary conditions
Fig. 10. (a) Single component design of the battery box access door with rib stiffened structure and (b) view of the door from the cosmetic side.
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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 4 Processing parameters for extrusion compression molding of E-Glass/PP
Table 2 FEA for a load of 2223 N (500 lbf) for various shell thickness Shell thickness (mm)
Displacement (mm)
Von Mises stress (MPa)
Maximum stress criterion (MPa)
Calculated mass (kg)
3 4 5
44.70 28.19 19.20
224 156 99
2.40 1.70 1.00
2.10 2.80 3.50
Process parameter
Predicted value
Actual value
Units
Mold closing velocity Mold temperature (top tool) Mold temperature (bottom tool) Maximum press force Charge length Charge diameter Cooling time
25 80 90 3500 650 170 120
25 80 130 3900 650 170 180
mm/s °C °C kN mm mm S
Table 3 Comparison 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) Density (kg/m3) Mass (kg) Total mass (kg)
9.40E04 7.86E+03 7.39 11.96
5.80E04 7.86E+03 4.57
3.14E03 1.21E+03 3.51 3.51
and the experimental set-up used to compare the predicted displacement. Load vs. displacement data obtained from both the FEA model and experimentally are compared in Fig. 14. The stiffness of the door as predicted by the FEA is consistent with the experimental stiffness until a deflection of 15 mm (78% of failure deflection). At this point the stiffness response transitions from a linear to a nonlinear state. The onset of nonlinearity can be attributed 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 loading in the mid-span and a stress concentration in the vertical rib in the surrounding area. The mechanical tests conducted on the battery door show similar failure like the FEA occurring on the vertical rib (Fig. 15). FEA predicted stress levels in excess of 155 MPa in the region which corresponds to the flexural strength reported for the E-glass/PP LFT material. 5.3. Fiber weight fractions and fiber length verification Fiber distribution and orientation in substructures, such as ribs or bosses, change and there by the expected stiffness of products cannot be obtained. These substructures also lead to fiber matrix separation in SMC molded parts [27]. Fiber–matrix separation leads to ribs with poor fiber content and resin-rich edges in large parts, weakening the structural integrity of the product. To validate the fiber distribution (fiber weight fractions) in the final molded part, representative samples from various sections of the door (Ribs and Skin) were sectioned and subjected to burn-off to separate the fiber from the matrix. The results of the burn-off study are tabulated in Table 5. The results show that there is no significant fiber matrix separation and the overall fiber weight fraction remained to be constant. Unlike SMC the LFT seem to flow through ribs of narrow width without any fiber matrix separation, there by the retaining the structural integrity. Fiber lengths were determined by image analysis and optical microscopy on fiber samples removed from the molded part after high temperature ashing. The fiber lengths of 600 individual fibers were measured from the molded part to determine the fiber length distribution. The fibers were dispersed in an aluminum pan and a stereoscope was used to capture several images. Post-processing of the dispersed fibers was done using the software Image Pro-Plus. Fig. 16 shows the fibers separated from the matrix and dispersed for length analysis. The fiber length distribution (Fig. 17) shows that the majority of the fibers are greater in length than the critical fiber length necessary for effective load transfer. The average fiber length for 600 fibers was 9.54 mm and approximately 80% of the fibers were greater than the critical fiber length of 3 mm.
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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.
Table 5 Results of fiber weight fraction at various regions of molded battery door
Fig. 14. Load vs. displacement obtained from FEA and experiment.
6. Summary and conclusions A LFT bus battery door was designed and fabricated using the extrusion–compression molding process. The LFT design incorporated ribs to the shell to add stiffness and rigidity in a single
Sample ID
Length (mm)
Width (mm)
Height (mm)
Fiber weight fraction
Rib – 1 Rib – 2 Rib – 3 Rib – 4 Rib – 5 Skin center – 1 Skin center – 2 Skin corner – 1 Skin corner – 2
25.75 26.25 26.30 29.25 27.50 22.50 22.50 22.50 22.50
4.50 4.50 4.50 9.00 9.00 22.50 22.50 22.50 22.50
14.50 14.50 14.50 19.88 19.88 5.00 5.00 5.00 5.00
40.15 39.80 41.04 38.28 38.42 39.25 39.93 40.28 39.23
component as opposed to the steel frame which was welded to the face sheet. The LFT design offered an approximate weight savings of 60%. The extrusion–compression molding process was simulated using the Cadpress-Thermoplastics. The flow fronts of the molten charge inside the tool and the fiber orientation in the final molded part were verified experimentally. Quantitative verification was performed to compare the stiffness predicted by the FEA model and the actual molded part. Fiber distribution was uniform throughout the molded area and there was no significant fiber matrix separation in narrow ribs. Fiber length analysis show minimum fiber degradation favoring effective load transfer.
Fig. 15. (a) Failure initiation on the ribs during mechanical testing and (b) FEA showing stress concentration (von Mises stress) on the ribs at the same location.
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Fig. 16. Fibers separated from resin and dispersed for fiber length analysis.
Fig. 17. Fiber length distribution plot.
Acknowledgements The authors gratefully acknowledge the support provided by the Federal Transit Administration (FTA), Department of Transportation, Project No. AL-26-7002 and Program Manager Terrell Williams of FTA. Technical help received from Juan Serrano and George Husman is also gratefully acknowledged. References [1] Häuptli A, Winski J. Direct processing of long fibre reinforced thermoplastics: selecting a feeding system. Plast Additives Compound 2003;5(5):36–9. [2] Steffens M, Himmel N, Maier M. Design and analysis of discontinuous long fiber reinforced thermoplastic structures for car seat applications. In: Proceedings of the International Conference on Computer Methods in Composite Materials, CADCOMP; 1998. p. 35–44.
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[3] Schut JH. Long-fiber thermoplastics extend their reach. Plast Technol 2003;49(4):56–61. [4] Bartus SD, Vaidya UK. Performance of long fiber reinforced thermoplastics subjected to transverse intermediate velocity blunt object impact. Compos Struct 2005;67(3):263–77. [5] Vaidya UK, Sriram R. Vibration damping of glass/polypropylene thermoplastic composites. In: International SAMPE Symposium and Exhibition (Proceedings), SAMPE 2004 Conference Proceedings, vol. 49; 2004. p. 757–62. [6] Schut J. Why long-glass molders are compounding in-line. Plast Technol 2002;48(4):52–9. [7] Bartus SD, Vaidya UK, Ulven CA. Design and development of a long fiber thermoplastic bus seat. J Thermoplast Compos Mater 2006;19(2):131–54. [8] Krause W, Henning F, Troster S, Geiger O, Eyerer P. LFT-D – A process technology for large scale production of fiber reinforced thermoplastic components. J Thermoplast Compos Mater 2003;16(4):289–302. [9] Fu S-Y, Lauke B. Effects of fiber length and fiber orientation distributions on the tensile strength of short-fiber-reinforced polymers. Compos Sci Technol 1996;56(10):1179–90. [10] Chawla KK. Interfaces. New York, NY: Springer-Verlag, New York, Inc.; 1998. p. 126. [11] Thomason JL, Vlug MA. Influence of fibre length and concentration on the properties of glass fibre-reinforced polypropylene: 1. Tensile and flexural modulus. Composites Part A: Appl Sci Manufact 1996;27(6):477–84. [12] Thomason JL, Vlug MA, Schipper G, Krikor HGLT. Influence of fibre length and concentration on the properties of glass fibre-reinforced polypropylene: Part 3. Strength and strain at failure. Composites Part A: Appl Sci Manufact 1996;27(11):1075–84. [13] Bockstedt RJ, Skarlupka RJ. Overview of long fiber reinforced thermoplastics. In: Technical Papers, Regional Technical Conference – Society of Plastics Engineers, Design for the 21st Century; 1996. p. 69–75. [14] Hartness T, Husman G, Koenig J, Dyksterhouse J. The characterization of low cost fiber reinforced thermoplastic composites produced by the DRIFT [trademark] process. Composites Part A: Appl Sci Manufact 2001;32(8):1155–60. [15] Kang S, Hieber CA, Wang KK. Optimum design of process conditions to minimize residual stresses in injection-molded parts. J Thermal Stresses 1998;21(2):141–56. [16] Wang VW, Hieber CA, Wang KK. Dynamic simulation and graphics for the injection molding of three-dimensional thin parts. J Polym Eng 1986;7(1):21–45. [17] Foss PH. Coupling of flow simulation and structural analysis for glass-filled thermoplastics. Polym Compos 2004;25(4):343–54. [18] Foss PH, Chiang HH, Inzinna LP, Tucker III CL, Heitzmann KF. Experimental verification of C-mold fiber orientation and modulus predictions. In: Annual Technical Conference – ANTEC, Conference Proceedings, vol. 1, 1995. p. 674–8. [19] Foss PH, Harris JP, O’Gara JF, Inzinna LP, Liang EW, Dunbar CM, et al. Prediction of fiber orientation and mechanical properties using C-MOLD and ABAQUS, In: Annual Technical Conference – ANTEC, Conference Proceedings, vol. 1, 1996. p. 501–5. [20] Shay Jr RM, Foss PH, Mentzer CC. Comparison of C-MOLD predictions and experimental shrinkages: fiber-filled materials. In: Annual Technical Conference – ANTEC, Conference Proceedings, vol. 1, 1996. p. 496–500. [21] VerWeyst BE, Tucker III CL, Foss PH. Performing an accurate fiber orientation analysis. In: Annual Technical Conference – ANTEC, Conference Proceedings, vol. 1, 1996. p. 726–30. [22] N.N., EXPRESS 6.1, User and Theroy Manual, Aachen, Germany, 2000. [23] Heber M. Modell zur rheologischen auslegung faserverstarkter thermoplastischer prebauteile. Aachen: RWTH; 1995. [24] Folgar F, Tucker III CL. Orientation behavior of fibers in concentrated suspensions. J Reinforced Plast Compos 1984;3(2):98–119. [25] Halpin JC, Kardos JL. Halpin–Tsai equations: a review. Polym Eng Sci 1976;16(5):344–52. [26] Shen H, Nutt S, Hull D. Direct observation and measurement of fiber architecture in short fiber–polymer composite foam through micro-CT imaging. Compos Sci Technol 2004;64(13–14):2113–20. [27] Christensen SK, Hutchinson B, Sun EM, Osswald TA, Davis BA. Fiber–matrix separation in ribbed SMC and BMC parts. In: Annual Technical Conference – ANTEC, Conference Proceedings, vol. 1, 1997. p. 782–7. [28] [cited 2007 August 28]; Available from: http://tools.ticona.com/tools/mcbasei/ product-tools.php?sPolymer=&sProduct=.
Contents lists available at ScienceDirect
Composites: Part A journal homepage: www.elsevier.com/locate/compositesa
Process simulation, design and manufacturing of a long fiber thermoplastic composite 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
Article history: Received 29 October 2007 Received in revised form 24 March 2008 Accepted 26 May 2008
Keywords: A. Long fiber thermoplastics E. Extrusion compression molding C. Process modeling B. Thermoplastic composites
a b s t r a c t 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 thermoplastic polymer matrix is reinforced with long glass or carbon fibers (3–25 mm) are processed by extrusioncompression 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 the flow of fiber filled viscous charge during the compression molding of the LFT composite. Studies on optimum charge size and placement in the tool, press force, temperature of mold, shrinkage and warpage were also conducted. The flow pattern of the molten charge in the mold and the resulting fiber orientation predicted by process simulation are verified experimentally. The studies have been applied for a mass transit/transportation component namely, a LFT battery box access door for form-fit-function to replace a heavy metal door. Weight reduction of 60% was achieved using 40% weight percent E-glass/ PP LFT over the metal design. Ó 2008 Elsevier Ltd. All rights reserved.
1. Introduction Long fiber thermoplastics (LFTs) are being used extensively in automotive and transportation industry due to their superior specific 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 contributing to environmental and economic benefits [2]. Global use of LFTs is expected to grow from around 40 million lbs in 2001 to 75 million lbs in 2007 [3]. In general, some of the advantages of using LFT over metals include high impact resistance, superior toughness, improved damping and corrosion resistance in conjunction with ease of shaping and recyclability [4,5]. The use of a thermoplastic matrix provides the molder the ability to modify and enhance the properties of the resin by blending additives, fillers and fire retardants depending on the nature of the application [6]. Various components have been designed and manufactured using LFTs for the transportation industry including, dashboard carriers, front ends, seat shells, battery trays, spare wheel dwells, etc. [2,7,8]. The typical applications of LFT components in an automobile are shown in Fig. 1 [8]. The mechanical properties of a part made of reinforced thermoplastics are defined by the matrix system, type of fibers, fiber con* Corresponding author. Tel.: +1 205 934 9199; fax: +1 205 934 8485. E-mail address: [email protected] (U.K. Vaidya). 1359-835X/$ - see front matter Ó 2008 Elsevier Ltd. All rights reserved. doi:10.1016/j.compositesa.2008.05.017
tent and orientation of the reinforcing fibers. The orientation and length of the fibers are influenced by the processing method and process parameters. LFTs possess starting fiber lengths of 3– 25 mm in contrast to short fiber thermoplastics (SFTs) compounds that possess 0.5 mm fiber length or less [9]. When processed optimally, LFTs possess a fiber length of 3–25 mm [2]. Hence the average fiber lengths of LFTs are an order of magnitude greater than the SFTs. The full strength of the reinforcements is utilized because the fiber length is above the critical fiber length for effective load transfer [10]. The stiffness of the laminate is directly proportional to the fiber concentration up to 40% by weight; and independent of fiber length above 0.5 mm [11]. Hence the use of long fibers has proven to increase the elastic modulus and the tensile strength of the material as close as to 90% of that obtained when using continuous fibers [12]. LFTs are manufactured by pulling continuous fiber tows through a thermoplastic polymer melt in a specialized processing die. Early manufacturing attempts mimicked wire-coating technology, crosshead extrusion or several pultrusion techniques that did not wet-out the individual fibers within the tow [13]. An alternate technique (Direct ReInforcement Fabrication Technology, DRIFT) [14], also referred to as hot-melt impregnation allows complete impregnation of continuous fibers with thermoplastics polymers at very high production rates, providing a high-quality, low cost thermoplastic composite. The hot-melt impregnation technology enables to produce products in various forms such as continuous rods, tapes, pultruded shapes, or pellets of any length for injection
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Fig. 1. Automobile components made of LFT [8].
and/or compression molding. This manufacturing process can be used to combine a wide variety of thermoplastic resins and reinforcing fibers. Fiber levels as high as 60% by weight are easily produced. The starting materials for LFTs are pellets of average length 3– 25 mm compared to the plate shaped semi-finished product of glass mat thermoplastics (GMT). LFT pellets are fed into the hopper of a plasticator (a single screw, low shear extruder) where they are metered down a barrel, heated above the melting point and extruded in low shear to form a molten charge. The molten charge is extruded to a predetermined size, and shape (usually cylindrical) that is transferred to the compression molding press for the forming operation. Thermal process parameters and the velocity gradient developed during the flow of the material influence the final mechanical property of the molded part. To optimize the processing of LFT, it is necessary to take into account a number of process variables and effects that are not typically encountered when processing unreinforced plastics. Molding pressures, fiber orientations, fiber distribution, and flow fronts of the molten charge within the tool are important parameters which determine the homogeneity of the molded product and therefore the success of the produced part. Recently developed software tools based on the finite element method (FEM) help in the design for manufacturing stage using these materials. By simulating the molding operation in a virtual environment, the effect of process variables on charge flow, mold filling, fiber orientation, shrinkage and warpage can be determined and observed before the tool is cut, ensuring a complete control of the process variables and being aware of its limitations. Several authors have implemented process models to verify injection molding of short fiber composites [15–21]. In the current work, a finite element (FE) simulation program for fiber filled polymer has been used to simulate the flow pattern, fiber orientation and process induced shrinkage/warpage of compression molded geometries [22]. It is during the filling stage that the flow induced fiber orientation develops, upon which the final mechanical and thermo-mechanical properties of the part are highly dependent. The material properties can be broadly classified into the type of analysis as shown in Fig. 2. The rheological behavior of the polymer and/or fiber filled resin under the molding conditions were used to calculate the flow front over the processing time. The flow front simulation predicts knit lines and entrapped air, the pressure and temperature distribution in the cavity and the clamping force. Because shrinkage and warpage have a decisive influence on the dimensional stability of the molded part, pressure volume temperature (PVT) characteristics exhibited by the material is also an input to the module. The mechanical analysis data specified (Young’s modulus, Poisson’s ratio, and aspect ratio) was used along with predicted fiber orientation distribution to calculate the final orthotropic material properties.
The present work considers an LFT composite with plate like and ribbed features. The LFT composite is investigated in terms of flow, fiber distribution, fiber orientation and design validation studies. The applied science studies have been extended to a real-world application namely a battery box access door for a mass transit bus, manufactured from a ribbed LFT 40 wt.% E-glass/PP (Eglass/PP) material. 2. Process modeling of extrusion compression molding of LFTs The process simulation used in the present work comprises four modules namely – flow, heat transfer, fiber orientation, and shrinkage/warpage. A solid model of the LFT composite was generated using Pro/Engineer Wildfire. The solid model was imported into the process simulation software after generating a three-noded finite element mesh in HypermeshÒ. The shell element is considered a 2.5 dimension membrane element with the thickness specified. For the simulation of compression molding, charge placement is defined by selecting an area on the finite element mesh, which corresponds to an extruded charge placed on a mold maintained at a lower temperature than the charge. The flow analysis in compression molding is modeled as non-Newtonian, under non-isothermal three-dimensional cavities using finite elements. This technique, commonly called as control volume approach (CVA), requires that the three-dimensional molding surface be divided into flat shell elements. The cells or control volumes are generated by connecting the element centroid with element mid-sides. When applying the mass balance to each cell, the resulting equations are identical to those arising from a Galerkin method for finite elements. The influence of the effect of temperature on the local viscosity of the material is captured by the Carreau–Williams Landel Ferry (WLF) model. This model is used to capture the temperature and deformation rate dependency of the viscosity [23] as given by:
g¼
P1 aT ð1 þ aT P2 jc_ jÞP3
ð1Þ
where
c_ corresponds to the shear rate, aT temperature shift coefficient, accounts for variation of viscosity 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 in a cold mold. The material that comes into contact with the mold walls is rapidly cooled; the local viscosity increases and the material in these regions will no longer flow. The filling stages of the compression molding process are temperature dependant, the
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Fig. 2. Material properties required for process modeling of LFTs.
calculation of the temperature distribution is an integral step of the overall simulation. The simplified form of the energy equations (2)–(4) used in the simulation of the heat transfer is as follows:
ume content, aspect ratio and a fiber interaction coefficient that depends on the number of fiber touches that occurred during the flow.
oT o2 T ¼k 2 ot oz oT oT Convection term: qcp vx þ vy ox oy ovx ovy Diffusion term: sxy syz oz oz
w ¼ f ð/; x; y; tÞ
Conduction term: qcp
ð2Þ ð3Þ ð4Þ
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 compression molding is essential to accurately predict the thermo-mechanical behavior and the final mechanical properties of the molded part. In general, the orientation of a particle, such as fiber is described by two angles namely; in plane orientation and out of plane orientation angle. These angles change in time as the melt flows thorough a die. In general the angular orientations of the fibers are represented only by the in plane orientation angle. Folger–Tucker model [24] is used to capture the flow induced fiber orientation. The model adopts a statistical approximation that is applied to the entire domain to predict the fiber orientation. The state of particle orientation at a point is described by an orientation distribution function, and is defined such that probability of a particle located at x, y at time t, being oriented between two angles is given by Eqs. (5) and (6). Assuming the fiber density is homogenous throughout, the continuity equation can be written as shown by Eqs. (7)–(9). The fiber distribution model accounts for the fiber vol-
Pð/1 < / < /2 Þ ¼
Z
ð5Þ /2
wð/; x; y; tÞ d/
ð6Þ
/1
ow o _ ¼ ðw/Þ ot o/ C 1 c_ ow ovx ovx 2 /_ ¼ cos / sin / sin / w o/ ox oy ov ov y y þ cos2 / þ sin / cos / ox oy ow o2 w ow ovx ovx 2 ¼ C 1 c_ 2 cos / sin / sin / ot ox oy o / o/ ov ov y y þ cos2 / þ sin / cos / ox oy o ovx ovx 2 cos / sin / sin / w o/ ox oy ovy ovy 2 þ cos / þ sin / cos / ox oy
ð7Þ
ð8Þ
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 interaction between the fibers, vx; vy; vz are the velocity component in x, y, and z, respectively. The anisotropic material properties resulting from the flow induced fiber orientation can be determined. Combinations of micro- and macromechanical theory are used to calculate the overall
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stiffness of the laminate. The Halpin–Tsai [25] micromechanical theory 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 configurations are shown in Fig. 3. Types 1–3 (Fig. 3a–c) have two small charges placed in different configuration, and these show the presence of knit lines when the two molten charges are compressed inside 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 premature freezing of the melt the approximate charge dimensions were deemed to be 650 mm in length and 170 mm in diameter and the force required to flow the molten charge inside the tool was predicted to be approximately 350 metric tons. The top tool temperate was maintained at 80 °C and the bottom tool at 90 °C. The flow front of the charge is seen to progress from the geometric center to the edges of the mold. The flow simulation shows that the four 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 simulation result. A short shot consist of placing a smaller volume dosing of the mold than required to complete fill the cavity. The part produced using a short shot provides information about the actual flow fronts developed inside the mold. Fig. 4 depicts the flow pattern comparison of the model which is 85% filled and a short shot of a charge respectively. In both the above cases the four corners did not fill and there is resemblance of the flow pattern predicted by the software. 3.2. Fiber orientation The fiber orientation in each element is represented by a plot of fiber distribution function and fiber angle. This plot represents the fiber orientation in an element with respect to the local XY plane of the element. The average fiber orientation distribution for five elements is illustrated in Fig. 5.
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The degree of orientation that occurred as the melt flows through the cavity is predicted by a fiber orientation scale. This scale is derived for five layers through half the thickness, from the top surface to the mid-plane. The 180° angle is divided into 25 sectors. For a randomly oriented layer there will be equal number of fibers in each sector or direction. On the other hand, for a preferential orientation, the fibers will tend to align in one direction and so most fibers will lie in just a few sectors. The value on the scale that represents no orientation is derived by dividing 1 by the 25 (total number of sectors), which yields the value 0.04. The more oriented the fibers become the less sectors, hence the fiber orientation scale value is greater than 0.04. Simulation results show areas where the top surface has random orientation compared to a preferential orientation at the center. X-ray radiographic studies were done to assess the fiber orientation of final molded part as shown in Fig. 6. The molded part was radiographed using a tungsten target Xray source at 40 kV. The part was placed between the X-ray source and an image intensifier connected to a charge coupled device (CCD) camera to obtain digital images. The images show a preferential fiber orientation in selected areas. Although the images show preferential orientation of fibers, it was difficult to determine the orientation of fibers through the thickness of the molded part. Hence an alternate method to determine the orientation through the thickness namely high resolution computerized tomography (micro-CT) was used [26]. Using X-radiation as a penetrating probe, the micro-CT affords detailed microstructural information from almost any material. To validate the fiber orientation predicted results, a representative sample from the molded part was analyzed for through the thickness by using a Scanco lCT40 Micro-CT apparatus. Cross-sectional images were obtained at various depths to capture the orientation effect through the thickness. The images obtained from Micro-CT were then analyzed using FiberScanÒ, advanced image processing software that determines the fiber distribution as a function of fiber orientation angle. Fig. 7 compares the fiber distribution plot obtained from simulation results with those generated from micro-CT images. The predicted results are in accordance with the micro-CT images showing the presence of a random
Fig. 3. Flow patterns for different charge placement and configurations: (a)–(c) shows flow patterns of two small charges placed in different orientations, and (d) shows the flow patterns of one long charge placed horizontally.
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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.
orientation on the face (skin) of the molded part and a preferential orientation through the thickness. This is termed as the skin–core effect in the molding of thermoplastic parts. During the solidification of the compression molded part, the residual stresses continue to build. The molded part experiences varying temperatures and stages of solidification. The flow simulation, fiber orientation calculation and the influence of material properties are then used as input to evaluate the thermo-mechanical response (shrinkage and warpage) of the molded part at the end of the compression molding process. The effect of temperature and pressure on thermal expansion is obtained from the pressure volume temperature (PVT) data for matrix material. The effect of fibers on thermal expansion coefficient is obtained by the combination of the PVT data of the matrix and the micromechanical Halpin–Tsai model developed for unidirectional orientation. Fig. 8
shows a representative deformation after the part has been demolded 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 manufactured 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 protect and house the several batteries needed for the regular operation of the electrical systems of the bus. It is currently comprised of an all steel sheet metal fascia which is bent to shape and then welded to a tubular steel frame which provides additional stiffness to the part. The metallic battery door is approximately 1 0.6 0.003 m and currently weights about 12 kg.
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Fig. 6. Radiograph image of molded part.
A LFT battery door was redesigned using the current metallic battery box door as the baseline. For the LFT battery door, the goal was to replace the steel frame by integrating the features offered by the frame into external face and therefore decreasing weight and assembly time and cost. In the integrated LFT battery door design the stiffness offered by the steel frame is achieved by the use of ribs integrated with the face (Fig. 10). The ribs provide extra dimensional stability to the part by controlling out of plane displacements caused by shrinkage and warpage of the parts as they are withdrawn from the tool and cooled to ambient temperature. The material selected for the LFT battery door had to meet various criteria including, but not limited to: (a) equivalent stiffness to that of 3 mm thick steel face sheet; (b) possess low weight and cost; (c) resist humidity and salt rich environments including battery acids; (d) possess dimensional stability, (e) ease of processing,
Fig. 8. Deformation predicted by software after the part cools.
and (f) paintable surface. Based on these requirements, long glass fiber reinforced polypropylene, E-glass/PP CelstranÒ PP-GF40-03 (40% fiber weight fraction, 25 mm long) produced by Ticona Inc. was selected. The mechanical properties 40% fiber weight fraction of 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 conducted using ANSYSÒ. The boundary conditions for the model were
Fig. 7. Fiber orientation for a representative sample obtained from Micro CT: (a) top surface showing a random orientation, (b) mid-plane through half the thickness showing a preferential orientation, (c) fiber distribution graph obtained from modeling showing random orientation on top surface, with a preferential orientation through the thickness, and (d) representative fiber frequency plot obtained from Micro CT images using FiberScanÒ.
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K. Balaji Thattaiparthasarathy et al. / Composites: Part A 39 (2008) 1512–1521 Table 1 Mechanical properties of E-Glass/PP CelstranÒ PP-GF40-03 (40% fiber weight fraction) [28] Property
Value
Units
Tensile modulus (1 mm/min) Tensile stress at break (5 mm/min) Tensile strain at break (5 mm/min) Flexural modulus (23 °C) Flexural strength (23 °C) Charpy notched impact strength (23 °C) Density
7900 100 2 8000 175 20 1210
MPa MPa % MPa MPa kJ/m2 kg/m3
5. Processing and component verification 5.1. Process verification
Fig. 9. Detail of Battery Box Door on the 60 BRT Bus Model.
selected based upon the assembly of the part to the exterior shell of a bus. Static FEA simulations were conducted for three different thicknesses (3, 4, and 5 mm) of the door shell using the properties of the LFT (PP-GF40-03) material. Fig. 11 shows a typical von Mises stress plot and the maximum deflection of the door for a shell thickness of 5 mm. The von Mises stress plot illustrates the stress concentration in the region of loading and the stress profile in the region of the stiffening ribs. The mid-span deflection for the 3 mm shell thickness LFT battery box door was excessive, and hence failed based on failure criteria (FC = 2.4, where a value of FC exceeding 1.0 denotes failure). A similar scenario was observed with the 4 mm shell thickness case (FC = 1.7). With the 5 mm thick version, the safety factor is close to the strength of the steel counterpart, and the deflection (19.2 mm) is not significant. A summary of the results from the FEA and the solid model is provided in Table 2. The weight between the steel frame and the LFT battery door design for 5 mm thickness shell is compared in Table 3. The percentage weight savings on the final LFT molded design was calculated to be approximately 60% compared to the steel frame design.
An oil heated/cooled two side matched steel tool was selected as the prototype/production tool. The tool has two main parts, a top tool which is a solid steel block that can be heated to the required molding temperature and has the required machining and surface finish to provide a class-A finish to the fascia (exterior) of the produced battery door and a bottom tool, which includes all the machined cavities to generate the ribbed structure on the back of the door and includes the detail to accommodate the lock housing and door handles. Fig. 12 shows the top and bottom tool mounted on the press with an extruded charge placed on the bottom tool. The tool was placed in a 400 metric ton press. A comparison of the process variables predicted by process modeling and actual values are tabulated in Table 4. 5.2. Design verification by mechanical testing The molded LFT battery door was mechanically tested to verify the stiffness predicted by the FEA model. The displacement predicted by the FEA model of the battery door was verified experimentally with the same set of boundary conditions. A test frame was fabricated, and a hydraulic jack was used to apply a load of 5500 N. The boundary conditions applied during testing were designed to replicate those seen by the door once mounted on the bus and during service. The load was applied from the exterior surface at the geometric center and the corresponding displacement on the transverse direction was measured using an electronic dial gauge that was placed in the geometric center on the interior. Fig. 13 shows the FEA model with the applied boundary conditions
Fig. 10. (a) Single component design of the battery box access door with rib stiffened structure and (b) view of the door from the cosmetic side.
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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 4 Processing parameters for extrusion compression molding of E-Glass/PP
Table 2 FEA for a load of 2223 N (500 lbf) for various shell thickness Shell thickness (mm)
Displacement (mm)
Von Mises stress (MPa)
Maximum stress criterion (MPa)
Calculated mass (kg)
3 4 5
44.70 28.19 19.20
224 156 99
2.40 1.70 1.00
2.10 2.80 3.50
Process parameter
Predicted value
Actual value
Units
Mold closing velocity Mold temperature (top tool) Mold temperature (bottom tool) Maximum press force Charge length Charge diameter Cooling time
25 80 90 3500 650 170 120
25 80 130 3900 650 170 180
mm/s °C °C kN mm mm S
Table 3 Comparison 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) Density (kg/m3) Mass (kg) Total mass (kg)
9.40E04 7.86E+03 7.39 11.96
5.80E04 7.86E+03 4.57
3.14E03 1.21E+03 3.51 3.51
and the experimental set-up used to compare the predicted displacement. Load vs. displacement data obtained from both the FEA model and experimentally are compared in Fig. 14. The stiffness of the door as predicted by the FEA is consistent with the experimental stiffness until a deflection of 15 mm (78% of failure deflection). At this point the stiffness response transitions from a linear to a nonlinear state. The onset of nonlinearity can be attributed 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 loading in the mid-span and a stress concentration in the vertical rib in the surrounding area. The mechanical tests conducted on the battery door show similar failure like the FEA occurring on the vertical rib (Fig. 15). FEA predicted stress levels in excess of 155 MPa in the region which corresponds to the flexural strength reported for the E-glass/PP LFT material. 5.3. Fiber weight fractions and fiber length verification Fiber distribution and orientation in substructures, such as ribs or bosses, change and there by the expected stiffness of products cannot be obtained. These substructures also lead to fiber matrix separation in SMC molded parts [27]. Fiber–matrix separation leads to ribs with poor fiber content and resin-rich edges in large parts, weakening the structural integrity of the product. To validate the fiber distribution (fiber weight fractions) in the final molded part, representative samples from various sections of the door (Ribs and Skin) were sectioned and subjected to burn-off to separate the fiber from the matrix. The results of the burn-off study are tabulated in Table 5. The results show that there is no significant fiber matrix separation and the overall fiber weight fraction remained to be constant. Unlike SMC the LFT seem to flow through ribs of narrow width without any fiber matrix separation, there by the retaining the structural integrity. Fiber lengths were determined by image analysis and optical microscopy on fiber samples removed from the molded part after high temperature ashing. The fiber lengths of 600 individual fibers were measured from the molded part to determine the fiber length distribution. The fibers were dispersed in an aluminum pan and a stereoscope was used to capture several images. Post-processing of the dispersed fibers was done using the software Image Pro-Plus. Fig. 16 shows the fibers separated from the matrix and dispersed for length analysis. The fiber length distribution (Fig. 17) shows that the majority of the fibers are greater in length than the critical fiber length necessary for effective load transfer. The average fiber length for 600 fibers was 9.54 mm and approximately 80% of the fibers were greater than the critical fiber length of 3 mm.
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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.
Table 5 Results of fiber weight fraction at various regions of molded battery door
Fig. 14. Load vs. displacement obtained from FEA and experiment.
6. Summary and conclusions A LFT bus battery door was designed and fabricated using the extrusion–compression molding process. The LFT design incorporated ribs to the shell to add stiffness and rigidity in a single
Sample ID
Length (mm)
Width (mm)
Height (mm)
Fiber weight fraction
Rib – 1 Rib – 2 Rib – 3 Rib – 4 Rib – 5 Skin center – 1 Skin center – 2 Skin corner – 1 Skin corner – 2
25.75 26.25 26.30 29.25 27.50 22.50 22.50 22.50 22.50
4.50 4.50 4.50 9.00 9.00 22.50 22.50 22.50 22.50
14.50 14.50 14.50 19.88 19.88 5.00 5.00 5.00 5.00
40.15 39.80 41.04 38.28 38.42 39.25 39.93 40.28 39.23
component as opposed to the steel frame which was welded to the face sheet. The LFT design offered an approximate weight savings of 60%. The extrusion–compression molding process was simulated using the Cadpress-Thermoplastics. The flow fronts of the molten charge inside the tool and the fiber orientation in the final molded part were verified experimentally. Quantitative verification was performed to compare the stiffness predicted by the FEA model and the actual molded part. Fiber distribution was uniform throughout the molded area and there was no significant fiber matrix separation in narrow ribs. Fiber length analysis show minimum fiber degradation favoring effective load transfer.
Fig. 15. (a) Failure initiation on the ribs during mechanical testing and (b) FEA showing stress concentration (von Mises stress) on the ribs at the same location.
K. Balaji Thattaiparthasarathy et al. / Composites: Part A 39 (2008) 1512–1521
Fig. 16. Fibers separated from resin and dispersed for fiber length analysis.
Fig. 17. Fiber length distribution plot.
Acknowledgements The authors gratefully acknowledge the support provided by the Federal Transit Administration (FTA), Department of Transportation, Project No. AL-26-7002 and Program Manager Terrell Williams of FTA. Technical help received from Juan Serrano and George Husman is also gratefully acknowledged. References [1] Häuptli A, Winski J. Direct processing of long fibre reinforced thermoplastics: selecting a feeding system. Plast Additives Compound 2003;5(5):36–9. [2] Steffens M, Himmel N, Maier M. Design and analysis of discontinuous long fiber reinforced thermoplastic structures for car seat applications. In: Proceedings of the International Conference on Computer Methods in Composite Materials, CADCOMP; 1998. p. 35–44.
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