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Composite Structures 80 (2007) 105–116

Design, manufacture and analysis of a thermoplastic compositeframe structure for mass transit

Haibin Ning, Uday Vaidya *, Gregg M. Janowski, George Husman

Department of Materials Science and Engineering, The University of Alabama at Birmingham, 1530 3rd Avenue South, BEC 254,

Birmingham, AL 35294-4461, United States

Available online 5 June 2006

Abstract

Thin-walled composite sub-elements possess excellent properties, including high specific strength, lightweight, internal torque andmoment resistance which offer opportunities for applications in mass transit and ground transportation. In the present work, an opensection thin-walled thermoplastic composite frame segment (sub-element) of a mass transit bus was designed, analyzed and manufacturedto replace a conventional metal-based design. Three cross-section configurations, rectangular, V-shape and rounded C-shape, were con-sidered, and different lamina stacking sequences, (0/90)6, [±45/(0/90)2]s, and [±45/(0/90)]3 were compared. Carbon fiber/polyphenylenesulphide (carbon/PPS) was the material choice, and single diaphragm forming (SDF) process was adopted to manufacture the framesegment. In-plane compression testing was conducted on the manufactured carbon/PPS composite frame to validate the finite elementanalysis results. A successful design concept to manufacture strategy of the open thin-walled carbon/PPS thermoplastic composite framesegment was demonstrated.� 2006 Elsevier Ltd. All rights reserved.

Keywords: Thermoplastic composites; Carbon fiber; Polyphenylene sulphide (PPS); Mass transit; Thin-walled structure

1. Introduction

Polymer matrix composites (PMCs) are being usedextensively in aerospace and transportation due to theirsuperior specific modulus and strength. By introducingthem into mass transit applications, composites help toincrease fuel efficiency and decrease maintenance costsdue to their low weight. Various components have beendesigned and manufactured for ground transportationincluding structural roof panels in high-speed railway coa-ches, bus structures, front cabins of high-speed locomo-tives, and non-structural interior panels [1]. Compositeshave also been featured in the structural body shell of masstransit buses. For example, the CompoBusTM, introducedthrough the technology developed by Tillson Pearson,Inc. (TPI), and now produced by North American BusIndustries (NABI), is manufactured from two glass fiber

0263-8223/$ - see front matter � 2006 Elsevier Ltd. All rights reserved.

doi:10.1016/j.compstruct.2006.04.036

* Corresponding author. Tel.: +1 205 934 9199; fax: +1 205 934 8485.E-mail address: uvaidya@uab.edu (U. Vaidya).

reinforced vinyl-ester resin face sheets with balsa coremonocoque shells. This non-metallic bus is more than30% lighter than a typical conventional 9 m bus andrequires 60% less power to run [2].

Thermoplastic composites have been successfully intro-duced into a wide range of applications previously filledby thermoset composites. In general, thermoplastics havesuperior impact resistance, high toughness and ease ofshaping and recycling compared to thermosets. Theseproperty advantages are usually observed when thermo-plastics are utilized as composite matrices. The use of ther-moplastic composites had been limited historically due toimpregnation difficulties and high temperature processingrequirements. A variety of processes have been developedin recent years to improve the impregnation of reinforcingfibers with thermoplastic polymers. These methods include(but are not limited to) FulcrumTM thermoplastic compositetechnology [3], commingled thermoplastic fabric [4],powder/sheath-fiber bundles [5], film stacking [6], powderpre-impregnation [7], wet processing method [8], Direct

106 H. Ning et al. / Composite Structures 80 (2007) 105–116

ReInforcement Fabrication Technology (DRIFT) [9], andfilament winding [10]. Superior impact resistance and largevolume production potential make thermoplastic compos-ites particularly attractive as structural materials in masstransit, rail and ground vehicles, military and aircraft struc-tures. They have excellent potential to maintain integrityfollowing an impact event because they do not exhibit cat-astrophic failure.

In this paper, the design, manufacture and analysis of athin-walled frame segment of a carbon fiber reinforcedpolyphenylene sulphide (PPS) thermoplastic composite isdiscussed. The frame segment represents a sub-element ofa mass transit bus. Different geometries were consideredin the design of the frame and the various stackingsequences of carbon/PPS pre-preg (also referred to as pre-form) were compared. The frame segment was fabricatedusing a single diaphragm forming approach and subse-quently tested in compression as a partial design verifica-tion. The process-induced microstructure of the carbon/PPS composite is also analyzed.

2. Background and literature review

Thin-walled beam sub-elements with open and closedcross-sections are used extensively in the aerospace indus-try, both as direct load carrying members and as stiffenersin panel constructions [11]. In a conventional mass transitbus, steel frames are used for the load-bearing body struc-ture. However, with the increasing importance of improv-ing transportation efficiency, composites have beenintroduced in mass transportation systems to replacemetallic parts and structures.

In transportation applications, stringent design require-ments call for structural frame members to possess highstrength, stiffness, and high damage resistance in conjunc-tion with less weight. Thermoplastic polymers have a crit-ical role as the matrix in these composites. Amongvarious candidate thermoplastics, such as PPS, polypropyl-ene (PP), and polyamide (nylon), PPS has excellent tensilestrength and flexural modulus. Furthermore, PPS has goodsolvent and chemical resistance as a result of its semi-crys-talline structure [12,13] and a high temperature resistance(continuous service temperature approximately 220 �C)compared to other polymers. Good flame resistance alsomakes it a preferred material for structural applications.

Fiber reinforced PPS has been utilized in aerospace andother applications, such as high-performance pumps [14].Ghaseminejhad [15] compared the impact behavior anddamage tolerance of carbon fiber reinforced PPS to carbonfiber reinforced polyetheretherketone (carbon/PEEK). Theauthor reported that carbon/PEEK panels showed an abil-ity to confine damage zones, whereas carbon/PPS panelsexhibited a high resistance to perforation through themechanism of extensive delamination. Manufacturingparameters have also been examined. For example,spring-back of fiber reinforced PPS composite parts canbe minimized and the post forming strength can be maxi-

mized by reducing the preheating time and increasing theforming temperature [16].

PPS is semi-crystalline, and the crystallinity plays animportant role in the end-properties of PPS and PPS-basedcomposites. Freddy and Lee [17] found that higher anneal-ing temperatures and longer annealing times resulted in anincreased critical buckling load due to an increase in thedegree of crystallinity which resulted in an increased axialcompressive modulus and in-plane shear modulus. Boeyet al. [12] found that the values of the creep stress andstrain rate of fiber reinforced PPS composites decreasedexponentially with the percentage of crystallinity. Thecreep deformation for composite samples with 20% and40% fiber were relatively similar, despite the differences inthe amount of fiber reinforcement. In carbon/PPS compos-ites, Cole [18] found that aging PPS in air at temperaturesnear its melting point causes structural changes whichresult in a lower ultimate degree of crystallinity and a lowermelting point.

Composite thin-walled structures with open profileshave been used in aerospace structures [19]. They offeradvantages of higher internal torque and moment resis-tance at reduced weight [20]. Fiber reinforced polymercomposite materials offer additional advantages such ascorrosion resistance and improved fatigue life and hencecan provide further optimization of strength and weightrequirements for open-profile sections [20].

Rand [21] found that open thin-walled orthotropic com-posite beams are much more sensitive to in-plane warpingthan similar isotropic beams which substantially modifiesthe bending stiffness of the beams and creates an upperlimit for applied loads. Bauld [11] developed a theory toanalyze the buckling behavior of laminated curved beamswith open sections made from fiber reinforced laminateswith midplane symmetry by extending classical thin-walledtheory for isotropic materials. Bank [22] used a modifiedbeam theory to describe the combined bending and twist-ing of anisotropic composite material open-section beamssubjected to pure bending and transverse loading. He andcoworkers [23,24] found that there existed a good correla-tion between experimental, finite element analysis, and the-ory for the twisting and bending deflection of a compositebeam. Maddur [20] analyzed the dynamic response ofopen-profile sections made of laminated composite materi-als using modified first order shear deformation theorywithout violating the assumption of zero mid-plane shearstrain. Wang et al. [25] studied buckling and post-bucklingbehavior of thin-walled structures made of laminated com-posite materials and found good agreement betweennumerical and experimental results.

Sheet forming technology has been widely used for man-ufacturing thin-walled thermoplastic composite products.There are four types of sheet forming technology: matchdie forming [26], roll forming [27], stretch forming [28],and diaphragm forming [29]. One or more diaphragms,typically silicone rubber film, are required in the diaphragmforming method. During processing, pressure is applied to

Fig. 1. Concept of the bus body panel and the frame.

H. Ning et al. / Composite Structures 80 (2007) 105–116 107

the heated pre-preg such that it is consolidated and forcedto conform to a die or a tool. The advantage of this form-ing process is that the diaphragm stretches and keeps thepreform in tension during forming, and thereby preventscompressive instabilities such as wrinkling, splitting andlocal material thinning. There are several variations ofthe diaphragm forming process. Double diaphragm form-ing is widely used and is derived from both vacuum form-ing of thermoplastic sheet and superplastic forming ofmetals, such as aluminum [29]. It is a process where anunconsolidated pre-preg is placed between two dia-phragms. Vacuum is applied in the chamber formed by amold and the bottom diaphragm. The system is heated tothe processing temperature of the thermoplastic, and thepre-preg is consolidated into the mold under hydrostaticpressure [29–35]. Single diaphragm forming (SDF) wasdeveloped by Olsen [36] as another technique to manufac-ture thermoplastic products. The difference between thisforming process and double diaphragm forming is thatone of the diaphragms is replaced by a tool. Thermoplasticpre-pregs are placed between the tool and the single dia-phragm, which can be silicone rubber or Upilex, a polyi-mide film [31,34]. Vacuum is applied between thediaphragm and tool, and pressure is applied above the dia-phragm to consolidate the pre-pregs. SDF process wasadopted in this work.

Fig. 2. (a) Rectangular profile; (b) 120� V-shape profile; (c) rounded C-shape

3. Design of thin-walled thermoplastic composite

The design of the thin-walled thermoplastic compositeframe segment is based on several factors that include:(a) the consideration of the cross-section profiles, (b) stack-ing sequence of the thermoplastic composite pre-preg and,(c) conformability to the mating/joining structural membersuch as side body panel segment [37]. The concept of theframe and side body panel is shown in Fig. 1. The bodypanel is a sandwich structure constructed with a thermo-plastic honeycomb core and face sheets made of glass fiberreinforced polypropylene on both exterior and interiorsurfaces. The work on the body panel, including the design,analysis and manufacture, is presented in detail in thereference [37]. The structural response of the frame seg-ment with different cross-section profiles, including rectan-gular, V-shape and rounded C-shape was compared usingfinite element analysis (FEA). These three cross-sectionprofiles are shown in Fig. 2. The final component configu-ration was determined from analysis on the three basicstacking sequences, namely, (0/90)6, [±45/(0/90)2]s, and[±45/(0/90)]3.

3.1. Linear static analysis

The geometries of the three profiles, rectangular, 120� V-shape and rounded C-shape, were created in Pro/Engineer2001 (Pro/E) as shown in Fig. 2 and then imported intoAltair� Hypermesh� 6.0 as IGS files for subsequent finiteelement analysis (FEA). SHELL 99 was used as the ele-ment type for the pre-processing in the Hypermesh 6.0.The mesh was imported as a CWD file to ANSYS 7.0 forFEA. The length, height and the width of the three profileswere kept identical for comparison purpose.

TowFlex� TFF-CPPS-103 12K carbon/2 · 2 Twill/PPSwas selected as the material for both modeling and eventualmanufacturing. It is a continuous fabric reinforced thermo-plastic fabric which has a woven 2 · 2 twill 12K carbonfiber tow impregnated with PPS powder. It has a resin con-tent of 43% by weight and 49% by volume [38].

The length, height and the width of the frame segmentwas 914 mm (3600), 63.5 mm (2.500), and 215.1 mm (8.4700),respectively. The element used in the analysis was a layeredstructural shell, SHELL99. This element has six degrees of

profile with the same length, a, the same width, b, and the same height, c.

Table 1Material properties of carbon/PPS

Elastic modulus Poisson’s ratio Shear modulus

EX (GPa) EY (GPa) EZ (GPa) mXY mXZ mYZ GXY (GPa) GXZ (GPa) GYZ (GPa)

64 64 7.2 0.3 0.25 0.25 5 3 3

Table 2ANSYS results for the three profiles with a 24 MPa applied pressure

Rectangular 120� V-shape RoundedC-shape

Stress applied (MPa) 24 24 24Force applied Per Node (N) 544 548 577r1 (MPa) 5.84 6.41 4.26r2 (MPa) 10.9 13.6 13.9r3 (MPa) 69.3 69.9 74.7s (MPa) 25.8 20.9 26.4Max. deflection (mm) 0.4 0.4 0.4Tsai–Wu ratio 0.62 0.49 0.49

108 H. Ning et al. / Composite Structures 80 (2007) 105–116

freedom at each node, namely, translation in the nodal x, y,and z directions, and rotations about the nodal x, y, and z-axes. Eight nodes, layer thicknesses, and orthotropic mate-rial properties define the element. Sixteen layers of wovencarbon/PPS were used in the construction of the laminatedframe segment for analysis. Orthotropic material proper-ties of the carbon/PPS laminate included elastic modulusin the x, y and z directions, Poisson’s ratio in the xy, yz,and xz directions, and shear modulus in the xy, yz, andxz directions, respectively. The nominal properties wereobtained from the manufacturer and are listed in Table 1.

The loading and boundary conditions were kept con-stant for all three profiles for FEA. The nodes at one endof the frame were fixed for rotation and translation in alldirections (x, y and z). Fig. 3 shows representative loadingand boundary conditions. A pressure of 24 MPa wasapplied at the other end. The 24 MPa pressure was basedon the weight of a representative bus, total number ofthe frames that would form the support structure, andthe cross-sectional area of each frame. The force was dis-tributed along the top nodes to simulate longitudinal com-pression as shown in Fig. 3. The end nodes where a verticalforce was applied were restricted from out-of-plane transla-tion to simulate a pinned boundary condition.

Table 2 lists the results from ANSYS analysis for thethree profiles. The results indicate that the third principalstress, r3, which is along the thickness direction, had thehighest value, 69.3 MPa, 69.9 MPa, and 74.7 MPa, for allthree profiles (the rectangular, 120� V-shape, and roundedC-shape profiles), respectively. The analysis also showed

Fig. 3. Loading and boundary conditions of the rectangular framestructure.

that the rounded C-shape and the 120 �C V-shape profilehas the lowest maximum Tsai–Wu ratio, which is directlyrelated to the strength of the structure. The Tsai–Wu ratiocriterion states that failure occurs when

A11r21 þ 2A12r1r2 þ A22r

22 þ A66s12 þ B1r1 þ B2r2 P 1 ð1Þ

Based on the FEA results, it was determined that therounded C-shape frame structure would be the best candi-date for the bus frame structure because it has the lowestTsai–Wu failure ratio, 0.49. In addition, the carbon/PPSframe segment had the required conformability to thegeometry of the thermoplastic glass/polypropylene body

Fig. 4. (a) 1905 mm frame with loading and boundary conditions; (b)transverse load to initiate pure buckling; (c) transverse load to initiatetorsion buckling.

Table 3Stress and deflection produced at the critical buckling load

Max. principal stress1 (r1) (MPa)

Max. principal stress2 (r2) (MPa)

Max. principal stress3 (r3) (MPa)

Max. In-plane shearstress (s) (MPa)

Max. Tsai–Wuratio

Max. out-of-planedeflection (mm)

(0/90)6 61.9 58.9 65.0 23.4 0.64 5.4[±45/(0/90)2]s 58.7 57.3 57.2 33.1 0.53 5.4[±45/(0/90)]3 59.1 56.6 69.4 22.8 0.70 5.9

0

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25000

30000

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Lo

ad (

N)

(0/90)6[45/(0/90)2]s[45/(0/90)]3

Fig. 5. Load vs. displacement for the pure buckling mode.

H. Ning et al. / Composite Structures 80 (2007) 105–116 109

panel segment (reported in [37]) to which it mates. Further-more, the gradual contours of the rounded C-shape profilemake it amenable to the SDF processing.

3.2. Non-linear analysis: Buckling

Following the initial study from the linear analysis,work was extended to a geometric non-linear analysis ofonly the rounded C-profile. Geometric non-linear analysiswas carried out on the rounded C-shape profile of1905 mm (7500) length representing the entire length of thesingle frame structure as shown in Fig. 1. A non-linear lay-ered structural shell, SHELL 91, was used for FEA. Theelement has six degrees of freedom at each node, namely,translation in the nodal x, y, and z directions and rotationsabout the nodal x, y, and z-axes. Eight nodes, fiber direc-tion angles, average or corner layer thicknesses, and ortho-tropic material properties define the element. The sameorthotropic material properties that were used in the linearanalysis (Table 1) were utilized for the non-linear analysis.

Pure buckling analysis was carried out to evaluate theperformance of the three stacking sequences of pre-pregsin this paper. The three configurations have the sameboundary and loading conditions as shown in Fig. 4(a).All nodes at one end are restrained for rotation and trans-lation in all directions. A vertical force of 15 kN based onthe bus weight and number of the structural frames, wasdistributed evenly along the top nodes to simulate in-planeloading. The nodes were restricted from out-of-plane trans-lation to simulate a pinned boundary condition at the endwhere the load was applied. A load of magnitude equal to5% of the vertical load, i.e. 750 N, was applied horizontallyin the transverse direction to initiate buckling in the framemember as shown in Fig. 4(b).

Six layers of the woven fiber (0�/90� balanced weave or±45� balanced weave) were considered for all of the lami-nated frame configurations. Each woven fabric layer wasmodeled as two individual layers. A total of 12 plies wereused for the three difference stacking sequences, namely(0/90)6, [±45/(0/90)2]s, and [±45/(0/90)]3. The principalstress directions correspond to the local shell element coor-dinates (x, y, and z respectively). Table 3 shows the resultsof the non-linear analysis for the three stacking sequences.It indicates that the (0/90)6 and [±45/(0/90)2]s configura-tions had similar maximum out-of-plane displacement,which was approximately 5.4 mm. The [±45/(0/90)2]s hadthe lowest Tsai–Wu ratio, 0.53.

The critical load for the three stacking sequences atwhich the deformation reaches non-linearity is shown inFig. 5. This figure also shows that the (0/90)6 configurationpossesses higher rigidity among the three due to its higherload-to-displacement ratio, 50,144 N/mm, while the [±45/(0/90)]3 configuration attains failure at the largestdisplacement.

The three laminate configurations were also analyzed fortorsion-buckling mode. The boundary conditions wereidentical to the ones in the pure buckling mode inFig. 4(a). A load for initiating torsion buckling was appliedat the center of the frame with loads acting in oppositedirections as shown in Fig. 4(c).

Table 4 shows the results of the torsion buckling analy-sis. The load vs. deflection of the three laminate configura-tions under torsion-buckling analysis is shown in Fig. 6.The results indicate that the (0/90)6 and [±45/(0/90)2]sconfigurations have similar Tsai–Wu ratio, 0.55 and 0.53respectively. The (0/90)6 configuration has the lowestmaximum out-of-plane deflection, 5.0 mm.

3.3. Non-linear analysis: Compression

Non-linear analysis was also carried out on a 914 mm(3600) frame segment with six consolidated layers of car-bon/PPS pre-preg. The shorter length (914 mm) as opposedto 1905 mm used in the buckling analysis was to inducecompression failure in the frame segment. The purpose ofthis analysis was to evaluate the performance of the frameduring in-plane compression. A load of 30 kN, determinedafter trials of analysis, was applied on one end of the frame

Table 4Stress and deflection produced at the critical buckling load in torsion

Max. principal stress1 (r1) (MPa)

Max. principal stress2 (r2) (MPa)

Max. principal stress3 (r3) (MPa)

Max. in-plane shearstress (s) (MPa)

Max. Tsai–Wu ratio Max. out-of-planedeflection (mm)

(0/90)6 49.9 48.7 73.4 21.2 0.55 5.0[±45/(0/90)2]s 38.9 52.2 61.6 33.7 0.53 5.7[±45/(0/90)]3 51.0 51.8 72.4 23.7 0.72 6.3

0

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20000

0.00 0.20 0.40 0.60 0.80 1.00Displacement (mm)

Lo

ad (

N)

(0/90)6[45/(0/90)2]s[45/(0/90)]3

Fig. 6. Load vs. displacement for the torsion buckling mode.

Fig. 7. Out-of-plane (buckling) displacements developed at failure load.

Fig. 8. Out-of-plane stress contour developed at failure load.

Fig. 9. Aluminum tool attached with three heating elements for manu-facturing the frame structure.

110 H. Ning et al. / Composite Structures 80 (2007) 105–116

towards the other end. The boundary conditions were iden-tical to the ones used in pure buckling and torsion bucklinganalysis. From the FEA result, failure occurred in theframe when the load reached to 25 kN. At the failure,out-of-plane displacements were formed. The largest out-of-plane displacement appeared to be 5.2 mm and occurredat the geometrical center as shown in Fig. 7 and the maxi-mum out-of-plane principal stress was found to be98.2 MPa at the failure as shown in Fig. 8.

4. Manufacturing

The rounded C-shape profile with six layers of 0�/90�carbon/PPS plain woven architectures was chosen for man-

ufacturing based on the FEA results. An aluminum moldwas designed consistent with the geometry of the roundedC-shape frame segment as shown in Fig. 9. Due to dimen-sion limitations of the molding press, the length of the toolwas determined to be 914 mm (3600) instead of the fulllength of one single frame, 1905 mm (7500). The aluminummold (equipped with cartridge heaters) was placed in acompression molding press of 400 metric ton capacity.The same press was also for manufacturing the body panelsegment reported in our earlier work [37]. Thermal insula-tion is critical for maintaining the pre-pregs at 320 �C

Fig. 10. SDF used for manufacturing the frame structure.

Fig. 11. A representative carbon/PPS frame segment manufactured by theSDF process.

H. Ning et al. / Composite Structures 80 (2007) 105–116 111

(608 �F), the processing temperature of PPS. Ceramicblocks were placed underneath the mold for insulating pur-poses. The SDF process was used in manufacturing thethin-walled carbon/PPS frame segment. A schematic ofthe SDF processing setup is shown in Fig. 10. Carbon/PPS pre-pregs were placed on the aluminum tool, equippedwith heaters. A pressure of 344 kPa and a vacuum of85 kPa were used to process the frame. The part was main-tained at a temperature of 320 �C for 50 min to obtaingood consolidation and then cooled to room temperature.A representative carbon/PPS frame segment manufacturedby the SDF process is shown in Fig. 11.

5. Mechanical testing and design verification

Samples were cut from a typical carbon/PPS frame seg-ment for flexural testing and low velocity impact (LVI)testing.

5.1. Flexural testing

Flexural testing was conducted according to ASTM D790M. Five carbon/PPS samples with dimensions of50 · 25 · 2 mm3 were prepared from the flat portion ofthe frame as shown in Fig. 11. The testing was carriedout on a screw operated universal testing machine with across-head motion rate of 0.85 mm/min. The results ofthe flexural testing are shown in Table 5. The average flex-ural modulus and ultimate tensile strength (UTS) of theSDF processed carbon/PPS composite was found to be28.50 GPa and 322.63 MPa, respectively. There was no vis-ible fiber breakage on the tensile face while apparent fiberfracture is observed around the loading line on the com-

pression face. A typical load vs. displacement plot for flex-ural test of a carbon/PPS specimen is shown in Fig. 12(a).The failure mode was wrinkling on the compression faceand debonding between the plies, as shown in theFig. 12(c) and (d), respectively. Multiple peaks in theload–displacement curve of Fig. 12(a) indicate debonding

Table 5Summary of flexural test on carbon/PPS specimens

Specimen Modulus (GPa) UTS (MPa) Statistical data Modulus (GPa) UTS (MPa)

1 22.54 255.94 Mean 28.50 322.632 26.85 295.30 Standard error 1.86 22.303 33.60 374.55 Standard deviation 4.15 49.864 30.63 367.86 Minimum 22.54 255.945 28.88 319.49 Maximum 33.60 374.55

0

5

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25

0 2 4 6 8 10

En

erg

y (j

ou

le)

0

1

2

3

4

5

Lo

ad (

KN

)EnergyLoad

Time (msec)(b)(a)

Fig. 13. (a) Impacted carbon/PPS specimen with the punctuation; (b) energy vs. time and load vs. time plot of LVI test.

Fig. 12. (a) Load vs. Displacement plot for the flexural test; (b) flexural specimen with the dimension of 25 · 50 mm2; (c) the wrinkling failure on theloading position; (d) side view of the specimen showing the debonding.

112 H. Ning et al. / Composite Structures 80 (2007) 105–116

between the plies. The first peak is due to the wrinkling onthe compression face, and the other peaks are attributed todebonding of the plies.

5.2. Low velocity impact (LVI) testing

LVI samples with dimension 100 · 100 mm2 were cutfrom the flat portion of the carbon/PPS frame segment(Fig. 11). The testing was conducted using a Dynatup8250 impact-testing machine. A hemispherical shaped headtup of diameter 19.5 mm was used. A representative plot ofenergy vs. time and load vs. time is shown in Fig. 13(a).The failure mode was mainly tensile fracture, as shown inFig. 13(a).

Thermogravimetric analysis (TGA) (TGA 2950, Du PontInstruments) and differential scanning calorimetry (DSC)

(DSC Q100, TA Instruments) were conducted on samplesfrom the frame segment to obtain the degree of crystallinity(DOC) and degradation temperature of PPS. Figs. 14 and 15show the TGA and DSC results of carbon/PPS produced bythe SDF process. The heat of fusion of PPS (43% by weightin the carbon/PPS) was found to be 52.6 J/g. The heat offusion of 100% crystalline PPS was estimated to be 80 J/gaccording to the work by Lovinger and coworkers [39].Therefore, it was estimated that the DOC of the PPS matrixin the carbon/PPS frame is 66% based on

DOC ¼ DH=DH f ð2Þ

where DH is the change of heat of fusion from temperatureT1 to temperature T2 and DHf: heat of fusion for a 100%crystalline PPS. The TGA graph shows that the degrada-tion temperature is 420 �C at room temperature.

Fig. 14. TGA showing the degrading temperature, 420 �C.

Fig. 15. DSC of carbon/PPS showing that the change of heat of fusion of PPS in the carbon/PPS is 52.6 J/g.

H. Ning et al. / Composite Structures 80 (2007) 105–116 113

Microstructural analysis was conducted on the carbon/PPS specimens. The samples were cut from the flat andthe curved portion of the frame segment and mountedin epoxy resin. After polishing, the carbon/PPS specimens

were examined under microscope. The representativemicrographs in Fig. 16 show that there is excellentimpregnation between the carbon fibers and the PPSmatrix.

Fig. 16. Microstructure of carbon/PPS frame: (a) flat portion; (b) curvedportion; and (c) higher magnification of flat portion.

114 H. Ning et al. / Composite Structures 80 (2007) 105–116

5.3. Component verification

The carbon/PPS frame segment produced by the SDFapproach was subjected to a compression test. The purpose

of the experiment was to validate the design and FEA pre-sented earlier. The frame segment was potted in aluminafilled epoxy at both ends in order to provide two parallelloading surfaces and alignment required for loading undercompression. Care was taken to prevent buckling failureand to promote failure primarily in compression. A dialgage was attached at the geometrical center transverse tothe frame segment to record transverse deflection, if any.The carbon/PPS frame segment was placed in the testingframe supported on rubber plates on both sides to preventits movement. Fig. 17 shows the carbon/PPS component‘before’ and ‘after’ failure.

6. Analysis and discussion: Component verification

Barreling was initially noted along the length of the car-bon/PPS frame segment at the early stages of loading in thecompression test. As the load increased, small ripples (localout-of-plane deformation) were observed at the geometriccenter of the frame as was predicted in FEA (Fig. 7). Thelocal crippling mode (ripple formation) was dominant indi-cating the frame segment underwent in-plane compression.

The failure occurred at a load of 34 kN. This experimen-tal failure load was 36% higher than the predicted failureload, which was 25 kN from FEA. Fig. 18 compares thecompression load vs. deflection of the carbon/PPS framefor the experiment and FEA. The figure shows that theexperimental deflection is larger than the FEA predictionat the same load. The difference was attributed to theboundary conditions achieved in experimental compressiontesting and the difference in thickness.

6.1. Weight savings

The weight of a 700 mm (2700) frame segment from thePro/E solid model was estimated to be 1 kg. The 700 mm(2700) segment length was chosen because it matches (andmates) to the manufactured glass/polypropylene (glass/PP) body panel segment in our previous work [37]. Theweight of the as-manufactured frame segment manufac-tured was measured to be 0.9 kg. An equivalent steel frame(as it exists on present buses) providing the same functionsand load-bearing is approximately 10 kg [40]. Hence, thecarbon/PPS, thermoplastic composite solution providesweight-saving on the order of 90%. In addition, the designcan be made flexible by adding or decreasing layers of thepre-pregs for different vehicle sizes and needs.

7. Conclusions

In this work, an open thin-walled carbon fiber rein-forced polyphenylene sulfide composite frame structurefor a mass transit bus was designed, manufactured andanalyzed successfully. Finite element analysis (FEA) ofthe frame configurations was carried out with the aid ofPro/Engineer 2001, HypermeshTM, and ANSYS�. A framewith an open cross-section featuring a rounded C-shape

Fig. 17. (a) Frame attached with dial gage prior to testing; (b) frame showing buckling failure after testing.

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10000

12000

14000

16000

18000

0 1 2 3 4

Load

(N)

Analysis

Experiment

Deflection (mm)

Fig. 18. Load vs. deflection plot of carbon/PPS under compression(experiment vs. analysis).

H. Ning et al. / Composite Structures 80 (2007) 105–116 115

profile with the stacking sequence of (0/90)6 was adoptedfor manufacturing, after different modes such as pure buck-ling mode and torsion buckling mode were analyzed. Theframe segment was successfully manufactured using poly-phenylene sulfide powder pre-impregnated carbon fabricand the single diaphragm forming process. The microstruc-tural investigation indicated that adequate consolidationresulted from the low-cost single diaphragm process. Thecompression testing of the frame segment validated thedesign and analysis. Flexure and low velocity impact testson coupons of single diaphragm formed carbon polyphen-ylene sulfide composites indicated primarily tensile domi-nated fracture behavior. A high degree of crystallinitywas noted in these specimens.

Acknowledgements

The authors gratefully acknowledge the financial sup-port from the Federal Transit Administration, Departmentof Transportation project # FTA-AL-26-7001 and techni-cal assistance from the National Composite Center(NCC), Kettering Ohio. The authors express their thanksto Chad Ulven, Selvum Pillay, Juan C. Serrano, AndrewGrabany and Patrick Moriarty for their assistance.

References

[1] Potluri P, Kusak E, Reddy TY. Novel stitch-bonded sandwichcomposite structures. Compos Struct 2003;59:251–9.

[2] Vaidya UK, Samalot F, Pillay S, Janowski GM, Husman G, GleichK. Design and manufacture of woven reinforced glass/polypropylenecomposites for mass transit floor structure. J Compos Mater2004;38(21):1949–72.

[3] D’Hooghe EL, Edwards C. Thermoplastic composite technology:tougher than you think. Adv Mater 2000;12(23):1865–8.

[4] Van West BP, Pipes RB, Advani SG. The consolidation of commin-gled thermoplastic fabrics. Polym Compos 1991;12:417–27.

[5] Ye L, Friedrich K. Processing of thermoplastic composites frompowder/sheath-fibre bundles. J Mater Process Technol 1995;48:317–24.

[6] Gamstedt EK, Berglund LA, Peijs T. Fatigue mechanisms inunidirectional glass-fibre-reinforced polypropylene. Compos SciTechnol 1999;59:759–68.

[7] Price RV. Production of impregnated rovings. US Patent # 03742106,1973.

[8] Goguelin M. Reinforced thermoplastics sheet and its manufacturingprocess. US Patent # 05008306, 1991.

[9] Hartness T, Husman G, Koenig J, Dyksterhouse J. The character-ization of low cost fiber reinforced thermoplastic composites pro-duced by the DRIFT process. Composites: Part A 2001;32(8):1155–60.

116 H. Ning et al. / Composite Structures 80 (2007) 105–116

[10] Henninger F, Hoffman J, Friedrich K. Thermoplastic filamentwinding with online-impregnation. Part B. Experimental study ofprocessing parameters. Composites: Part A 2002;33:1677–88.

[11] Bauld N, Tzeng L. A Vlasov theory for fiber-reinforced beamswith thin-walled open cross sections. Int J Solids Struct 1984;20:277–97.

[12] Boey F, Lee TH, Khor KA. Polymer crystallinity and its effect on thenon-linear bending creep rate for a polyphenylene sulphide thermo-plastic composite. Polym Test 1995;14:425–38.

[13] Fernandez I, Blas F, Frovel M. Autoclave forming of thermoplasticcomposite parts. J Mater Process Technol 2003;143–144:266–9.

[14] Using polyphenylene sulphide in high-performance pumps. WorldPumps 2002, 2002. p. 27–31.

[15] Ghaseminejhad MN, Parvizi-Majidi A. Impact behaviour anddamage tolerance of woven carbon fibre-reinforced thermoplasticcomposites. Construct Build Mater 1990;4:194–207.

[16] Boey F, Lua AC, Yue CY. Hot forming of a polyphenylene sulphidecomposite. J Mater Process Technol 1993;38:327–35.

[17] Boey F, Lee TH. Effect of matrix crystallinity on the buckling failureof a PPS thermoplastic composite. Polymer Testing 1994;13:47–53.

[18] Cole KC, Noel D, Hechler JJ. Crystallinity in PPS – carboncomposites: a study using diffuse reflection FT-IR spectroscopy anddifferential scanning calorimetry. J Appl Polym Sci 1990;39:1887–902.

[19] Jung SN, Lee JY. Closed-form analysis of thin-walled composite I-beams considering non-classical effects. Compos Struct 2003;60:9–17.

[20] Maddur SS, Chaturvedi SK. Laminated composite open profilesections: first order shear deformation theory. Compos Struct1999;45:105–14.

[21] Rand O. Nonlinear in-plane warping deformation in elasticallycoupled open thin-walled beams. Comput Struct 2001;79:281–91.

[22] Bank LC. Modifications to beam theory for bending and twisting ofopen-section composite beams. Compos Struct 1990;15:93–114.

[23] Bank LC, Cofie E. A modified beam theory for bending and twistingof open-section composite beams – Numerical verification. ComposStruct 1992;21:29–39.

[24] Smith SJ, Bank LC. Modifications to beam theory for bending andtwisting of open-section composite beams – Experimental verification.Compos Struct 1992;22:169–77.

[25] Wang C, Pian T. Hybrid semiloop element for buckling of thin-walledstructures. Comput Struct 1988;30:811–6.

[26] Krebs J, Friedrich K, Bhattacharyya D. A direct comparison ofmatched-die versus diaphragm forming. Composites: Part A1998;29A:183–8.

[27] Henninger F, Friedrich K. Production of textile reinforced thermo-plastics profiles by roll forming. Composites: Part A 2004;35:573–83.

[28] Lim TC, Ramakrishna S, Shang HM. Axisymmetric sheet forming ofknitted fabric composite by combined stretch forming and deepdrawing. Composites: Part B 1999;30:495–502.

[29] O’Bradaigh CM, Mallon PJ. Effect of forming temperature on theproperties of polymeric diaphragm formed thermoplastic composites.Compos Sci Technol 1989;35:235–55.

[30] Krebs J, Bhattacharyya D, Friedrich K. Production and evaluation ofsecondary composite aircraft components – a comprehensive casestudy. Composites: Part A 1997;28A:481–9.

[31] Bersee HEN, Beukers A. Diaphragm forming of continuous fibrereinforced thermoplastics: influence of temperature, pressure andforming velocity on the forming of Upilex-R diaphragm. Composites:Part A 2002;33:949–58.

[32] Monaghan MR, Mallon PJ. High temperature mechanical charac-terization of polyimide diaphragm forming films. Composites: Part A1998;29:265–72.

[33] Channer KJ, Cosgriff W, Smith GF, Okoli OI. Development of thedouble RIFT diaphragm forming process. J Reinforced PlasticsCompos 2002;21:1629–35.

[34] Pantelakis SG, Baxevani EA. Optimization of the diaphragm formingprocess with regard to product quality and cost. Composites: Part A2002;33:459–70.

[35] Walczyk DF, Hosford JF, Papazian JM. Using reconfigurable toolingand surface heating for incremental forming of composite aircraftparts. J Manufact Sci Eng 2003;125:333–43.

[36] Olsen SH. Single diaphragm forming of drapeable thermoplasticimpregnated composite materials. US Patent # 05037599, 1991.

[37] Haibin Ning, Gregg M. Janowski, Uday Vaidya, George Husman.Thermoplastic sandwich structure design and manufacturing for thebody panel of mass transit vehicle. Compos Struct, in press,doi:10.1016/j.compstruct.2006.04.090.

[38] www.hexcelcomposites.com.[39] Lovinger AJ, Davis DD, Padden Jr FJ. Kinetic analysis of the

crystallization of poly(p-phenylene sulphide). Polymer 1985;26(11):1595–604.

[40] Husman G, Vaidya UK, Janowski GM, Serrano JC. FTA reportFTA-AL-26-7001-2004.1, July 2004.