SHEAR DEFORMATION PROPERTIES OF GLASS-FABRIC SHEET MOLDING COMPOUND

Transcription

1 SHEAR DEFORMATION PROPERTIES OF GLASS-FABRIC SHEET MOLDING COMPOUND Caroline Dove Automotive Composites Consortium, United States Council for Automotive Research Abstract The Automotive Composites Consortium (ACC), a partnership of Chrysler Group LLC, Ford Motor Company, General Motors Company, and the U.S. Department of Energy, conducts pre-competitive research on structural and semi-structural polymer composites to advance high strength, lightweight solutions in automobile technology. An ACC focal project concerning the development of a structural composite underbody was established to provide methodologies and data for each ACC member company to implement lightweight, cost-effective structural composites in high-volume vehicles. This objective will be fulfilled through design, analysis, fabrication, and testing of a structural composite underbody. In the current phase of the project, a draping analysis of a full underbody made of woven glass-fabric sheet molding compound (SMC) was used to identify changes in local mechanical properties due to fabric shearing during compression molding. As a laboratory-scale effort, woven glass-fabric SMC was compression molded into 390 mm x 200 mm x 60 mm double dome shapes and 610 mm x 610 mm flat plaque configurations of three separate 4-ply layups: (0/90) s, (45/- 45) s, and (45/0) s. Double domes underwent static crush, impact, and mechanical testing; mechanical properties were further compared to corresponding flat plaque properties. All data was used to broaden the material property database and validate model predictions of strand orientations in a molded part. Background The ACC was formed in 1988 with a mission to provide pre-competitive, collaborative efforts towards researching structural and semi-structural polymer composites. An ACC focal project concerning the development of a structural composite underbody was established with the following objectives [1]: Develop methodologies and data for each ACC member company to implement lightweight, cost-effective structural composites in high-volume vehicles. Design and fabricate structural automotive components with reduced mass and cost, and with equivalent or superior performance to existing components. Develop new composite materials and processes for the manufacture of these high volume components. These objectives will be fulfilled through design, analysis, fabrication, and testing of a structural composite underbody with continuous oriented fibers, focusing on cycle time and structural joining of composite to steel. In Phase 1 of the project, a woven glassfabric SMC was established as the material of choice. In this phase of the project, a finite element draping analysis of a full underbody made of fabric SMC was designed to identify changes in local mechanical properties due to fabric shearing during compression molding [2]. There are several factors that must be accounted for in the Page 1

2 finite element model. However, not included are the affects of fiber bundling and spreading changes that occur during molding. As a laboratory-scale effort, a physical "double dome" shaped part was molded (Figure 1) to confirm the finite element model predictions of fabric shear angle and fiber tension in a molded part. This shape was chosen for analysis because of its ability to show different draping effects of representative curvatures associated with the full underbody. The double dome dimensions are approximately 390 mm long, 200 mm wide, and 60 mm tall. The top region is flat with side walls on two opposite sides and dome shapes on the ends. The side walls and dome ends taper into a lateral, flat flange section. The finite element model successfully predicted fabric shear deformations in a molded part with the exception of the flange area, which exhibited yarn sliding not accounted for in the model. While this modeling provided insight into the practicality of the chosen composite layup for the full underbody, more detail was needed on double dome properties to further establish how well the finite element model predicts molded part structural behavior. (a) (b) Figure 1: As-molded 4-ply (45/-45) s double dome top view (a) and bottom view (b). Once the molded part strand configurations validated the model, a second molding trial of double domes with three separate 4-ply woven fabric layups- (0/90) s, (45/-45) s, and (45/0) s - was completed. This paper discusses this molding and subsequent testing performed to evaluate the molded parts and update the model. Additionally, a comparison of tensile and compressive properties between the double domes and flat plaque configurations of the same layups was investigated. The results allow property extrapolations from flat plaques to the double domes, and aid in modeling the behavior of molded parts. Materials and Processing Experimental Material for double domes and plaques was a version of sheet molding compound. However, instead of chopped glass fiber reinforcement, a woven 1854 style glass-fabric sheet was fed between two layers of vinyl ester resin paste. Prior to molding, stacking sequences for both double domes and flat plaques were laid-up by hand. Individual plies were cut out of the fabric at 0, 90, 45, and -45 o angles. The dimensions of the double dome plies are shown in Figure 2; four plies were compiled to form each charge. For plaques, the individual plies as well as the charges for the flat plaques were 584 mm by 584 mm (24 in by 24 in). Double domes were compression molded in a hydraulic press for 180 s with an upper tool temperature of 140 C and lower tool temperature of Page 2

3 145 C. Flat plaques were compression molded in a hydraulic press for 180 s with an upper tool temperature of 149 C and lower tool temperature of 154 C. (a) (b) Figure 2: 0/90 ply dimensions (a) and 45/-45 ply dimensions (b) for double domes. Testing, Characterization Methods, and Standards After molding, the double domes were divided into groups for three separate types of testing: static crush testing, impact testing, and mechanical testing. Table I displays the test matrix. Table I. Test matrix for crush, impact, and mechanical testing of double domes. Layup Total # of Parts # Molds for Crush Testing # Molds for Impact Testing # Molds for Mechanical Testing Top Tensile (1 ea.) Mechanical Testing Breakdown (# specimens from each mold) Top Compression (1 ea.) Side Compression (4 ea.) Flange Compression (2 ea.) (0/90) s (45/-45) s (45/0) s Crush Testing Flash was removed from the edges of the molded double domes using a diamond blade band saw. Samples consisted of five double domes for each layup, totaling 15 specimens. Each double dome was placed between two parallel plates in a MTS Sintech 30/G and crushed to a crosshead displacement of 50 mm. Load vs. displacement curves were recorded in MTS Testworks Software for later analysis. Impact Testing A 50.8 mm diameter steel ball was dropped from between 152 to 914 mm heights (approximately 0.82 J, 1.73 m/s to 4.94 J, 4.24 m/s) on the dome ends of (45/0) s double domes. Additionally, a (45/-45) s double dome was impacted from 914 mm. Some domes were inadvertently hit twice with the ball as it dropped, and are noted in the discussion as double-strike. Impacted double domes underwent dye penetrant testing and vibrothermography to aide in determining a threshold for damage onset as well as level of damage. A MagnaFlux Dye Penetrant Kit with ultra-violet lamp was used for Page 3

4 both NDE techniques. Testing was completed on the top (impact side) surface and bottom (underside) surface of all samples. A FLIR SC-3000 QWIP Cooled Detector and FLIR Researcher v. 2.8 were used for infrared detection and analysis, respectively. Branson Sonic Excitation settings included a 15 khz sonic source, 2 s pulse length, 1100 W of power applied, 360 W of power absorbed, 20 lb air horn pressure, and a wet tissue coupler. To see exterior surface damage, one sonic pulse was needed, while interior surface damage required three sonic pulses to generate enough heat to see results. Impacted double domes were subsequently crush tested under the same conditions mentioned above to assess any changes in peak load and load vs. displacement characteristics after impact. Mechanical Testing For mechanical testing, 127 x 19 mm rectangular compression and 173 x 25 mm (16 mm gage width) dog-bone tensile specimens were cut from the double domes. These sizes were not standard due to size limitations of the double domes, but all other specifications were followed for testing [3,4]. Per the test matrix, each double dome allowed two compression specimens from the flange, four compression specimens from the side walls, and one compression specimen from the top flat surface. In lieu of a top compression specimen, some double domes had a tensile bar cut from the top flat surface (Refer to Table I for number of specimens). Specimens were also cut from flat plaques to compare with the mechanical properties of the double dome specimens. Each plaque sample set consisted of six specimens. Plaque samples were cut in 0, 90, 45, and -45 o directions, while tensile samples were cut from double domes in a 0/90 o orientation. All tension and compression specimens were tested using a MTS Sintech 30/G and data recorded with MTS Testworks Software. For tension testing, a MTS 50.8 mm longitudinal extensometer was used for determining strain and modulus values; for compression testing, an 8 mm MTS compression extensometer was used. Fiber content and density [5] were also determined for all samples. Square specimens cut to approximately 25 mm 2 were weighed, placed in a furnace set between o C for two hours to burn off resin, and weighed again. Burn-offs were digested using dilute sulfuric acid to remove calcium carbonate filler. Digested specimens were placed in an oven overnight at 100 o C to dry and weighed a final time. Fiber weight fractions were calculated. To determine density, the Archimedes' method was used, comparing specimen weights in air and in water. Results and Discussion Peak loads (and coefficients of variance) obtained from crush testing for the (0/90) s, (45/-45) s, and (45/0) s layups were 8.1 (7.2%), 12.2 (5.9%), and 10.9 kn (4.4%), respectively. Figure 3 shows representative load vs. displacement curves obtained for each layup from the crush testing. As the double domes were crushed, the flanges exhibited increasing curvature and cracking. Additionally, when the applied load was released from the double domes, they reverted to their original dome-shape, although damage was visible. The measured load vs. displacement curves were used to compare to structural finite element analysis simulations discussed in [6]. Page 4

5 Load vs. Displacement for Double Dome Crush Tests Load (kn) Displacement (mm) 0/90 45/-45 45/0 Figure 3. Representative load vs. displacement curves from crush testing for (0/90) s, (45/-45) s, and (45/0) s. Dye penetrant (DP) testing and vibrothermography (VT) indicated damage from all drop heights. DP testing showed damage on top and bottom surfaces from 457, 610, a- nd 914 mm drop heights. For 152 and 305 drop heights, no top surface damage was visible, however bottom surface damage was detected. The same behavior was also observed for a 152 mm drop height double-strike. Vibrothermography required multiple sonic pulses to identify bottom surface damage compared to one pulse for the top surface. Figure 4 shows dye penetrant and vibrothermographic images of a (45/0) s layup double dome impacted from 610 mm. The scale in the figure identifies the amount of heat produced in a specific region by the vibrations induced into the crack, thus making the crack visible. While VT successfully shows the presence of cracking, a quantifiable value defining the degree of damage is unknown. (a) (b) (c) 23.3 C C Figure 4. Images of dye penetrant inspected damage sites for the top surface (a) and bottom surface (b) of a (45/0) s layup dome impacted from 610 mm and corresponding vibrothermographic image (c). Crush testing after impact showed no depleting affect on the peak load compared to unimpacted double domes. Load vs. displacement curves also exhibited similar shapes to those seen with unimpacted double domes. Additionally, while impacted double domes also reverted to their original shape, more damage was visible which lessened the pop-back effect. Figure 5 shows a comparison of (45/0) s unimpacted and impacted double domes after crush testing. Page 5

6 Figure 5. Top views of unimpacted (left) and impacted (right) (45/0) s layup double domes after crush testing. More pop-back behavior and less damage were seen for the unimpacted double dome. Double dome fiber content on the top section from where the tensile bars were taken from was approximately 70 wt%, while flat plaques contained approximately wt%. Fiber content variability was determined across the double dome, and showed a decrease towards the outer regions of the flanges attributed to extra squeeze out from the shear edge mold. Additionally, top regions of the double domes exhibited less yarn sliding than expected, while the flange regions showed more yarn sliding than expected. Double dome densities averaged 1.94 g/cm 3 versus 1.78 g/cm 3 for plaques. Higher compaction of the double domes during molding and the lack of hydrostolic pressure from shear edges account for increased fiber content and densities as compared to flat plaques. Double dome tensile and compressive properties were higher than plaque properties. This is attributed to the higher fiber content present in the double domes as well as a 35% decrease in specimen thickness from plaques to double domes. Table II displays the results of the mechanical testing. Plaque tensile and compressive properties are comparable and exhibit reasonable standard deviations. However, it is difficult to draw conclusions based on the double dome data. There were only two tensile specimens within each sample, leading to large coefficients of variance. Additionally, there were no tensile bars tested from the side walls or flange region because of geometry restrictions, and therefore no comparison can be made between tensile and compressive properties in those regions. However, it is clear that higher tensile strengths are seen in the top sections of the double domes than compressive strengths. This is due to the large amount of fibers in this region. During compression testing, these fibers rub against each other with little resin interaction, causing a negative effect on compressive strength. It may also be noted that certain samples contain large standard deviations which is a result of fewer specimens. Page 6

7 Table II: Comparison of average tensile and compressive properties of double domes and flat plaques (standard deviations in parentheses). Layup, Sample Type, & Test Direction Thickness (mm) Fiber Content (wt%) Young's Modulus (GPa) Tensile Properties Ultimate Tensile Strength (MPa) Strain at Break (%) Compressive Properties Compressive Modulus (GPa) Max Compressive Stress (MPa) (0/90) s Double Dome Top Side Flange (0/90) s Plaque (45/-45) s Double Dome Top Side Flange (45/-45)s Plaque (5.4) 21.2 (1.5) 23.0 (3.6) 11.8 (0.6) 13.4 (1.0) 32.4 (1.5) 13.2 (1.0) 12.3 (1.4) 19.1 (0.7) 19.6 (2.6) 398 (20) 277 (23) 283 (43) 93 (5) 133 (7) 197 (2) 115 (18) 113 (30) 254 (8) 241 (46) 1.38 (0.19) 1.79 (0.51) 1.60 (0.23) 2.02 (0.34) 2.05 (0.13) 1.34 (0.02) 1.87 (0.30) 2.18 (0.39) 2.06 (0.21) 1.80 (0.21) TBD 26.5 (9.2) 18.1 (7.0) 23.9 (4.4) 26.1 (5.4) 14.4 (5.4) 17.6 (3.9) 14.7 (2.4) 14.5 (1.7) 9.1 (1.3) 14.0 (4.5) 26.9 (2.4) 22.8 (3.0) 333 (1) 137 (31) 99 (31) 254 (17) 284 (26) 123 (7) 129 (8) 94 (8) 96 (1) 116 (5) 101 (11) 294 (8) 243 (13) (45/0) s Double Dome Top Side Flange (45/0) s Plaque (1.1) 17.8 (1.8) 17.8 (2.3) 19.8 (1.0) 12.7 (1.1) 338 (2) 163 (17) 205 (23) 305 (23) 111 (6) 1.29 (0.09) 1.30 (0.14) 1.69 (0.36) 2.01 (0.26) 1.34 (0.14) TBD 20.1 (4.5) 19.9 (3.1) 20.6 (3.0) 15.5 (0.3) 23.7 (2.9) 17.1 (2.3) 217 (11) 99 (12) 104 (5) 179 (26) 197 (4) 232 (14) 165 (14) Summary and Next Steps This project successfully provided insight on the structural behavior and properties of a molded part in comparison to flat plaques. This data will allow for better finite element model predictions moving forward. Crush testing supplied the range of peak loads for the three different double dome layups, as well as an idea of the spring-back behavior of the composite. Additionally, crush testing of impacted double domes comparatively showed more damage and less spring-back than unimpacted double domes. Dye penetrant testing and vibrothermography illustrated impact heights at which damage is Page 7

8 detectable. DP testing was able to detect damage on bottom surfaces only at lower impact heights. Also, a quantifiable value determining the degree of damage will need to be identified for vibrothermography to better evaluate the technique's results. As previously mentioned, flat plaque tensile and compressive properties are comparable with reasonable standard deviations, but it is difficult to draw conclusions from the double dome mechanical properties due to the large standard deviations and small tensile sample sizes. In the future, a larger sample size is needed for the top sections of double domes to get a more accurate idea of their tensile properties. Additionally, an alternative geometry for tensile measurements will be established so that tensile bars can be taken from side walls and flange sections to record the differences throughout the molded part and compare to compression data. Finally, a better molding method should be determined to produce parts with more evenly spread fibers throughout. Acknowledgments The author thanks the ACC Underbody team for their significant contributions, especially Libby Berger (General Motors Co.) for her assistance and consultation in evaluating test results, Dan Simon (General Motors Co.) for impact testing and NDE analysis, Chuck Knakal (USCAR ACC) for impact testing, and Dan Houston (Ford Motor Co.) for sharing his technical insight in evaluating test results. Additional thanks to Jeff Dahl (Ford Motor Co.) and Ray Silva (Ford Motor Co.) for assistance with double dome molding, Century Tool & Gage for molding of flat plaques, and Continental Structural Plastics for compounding the fabric SMC. Funding for this project was provided by the Automotive Composites Consortium and U.S. Department of Energy Cooperative Agreement No. DE-FC05-02OR Page 8

9 References [1] Berger, L., "Automotive Composites Consortium Structural Composite Underbody," Society of Plastics Engineers, Automotive Composites Conference and Exhibition, Troy. MI, Sept , [2] Sherwood, J., Jauffrès, D., Fetfatsidis, K., Winchester, D., Chen, J., "Mesoscopic Finite Element Simulation of the Compression Forming of Sheet Molding Compound Woven-Fabric Composites," Society of Plastics Engineers Automotive Composites Conference and Exhibition, Troy, MI, Sept , [3] ASTM Standard D3039, 2000 (2006), "Tensile Properties of Polymer Matrix Composite Materials," ASTM International, West Conshohocken, PA, [4] ASTM Standard D3410, 2003 (2008), "Compressive Properties of Polymer Matrix Composite Materials with Unsupported Gage Section by Shear Loading," ASTM International, West Conshohocken, PA, [5] ASTM Standard D792, 2000 (2008), "Density and Specific Gravity (Relative Density) of Plastics by Displacement," ASTM International, West Conshohocken, PA, [6] Fuchs, H., Deslauriers, P. and Conrod, B, "ACC Composite Underbody Structural Test-Analysis Correlation Studies," Society of Plastics Engineers, Automotive Composites Conference and Exhibition, Troy. MI, Sept , Page 9