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STATUS OF THE COMPOSITE UNDERBODY COMPONENT AND ASSEMBLY STRUCTURAL TEST-ANALYSIS CORRELATION
Hannes Fuchs and Eric Gillund
Multimatic Engineering
Abstract
Computer aided engineering-based design methodologies have been utilized throughout
the Automotive Composites Consortium Focal Project 4 to assess the vehicle level structural
stiffness and impact performance of the composite underbody design proposals, and to
estimate the potential mass reduction for several candidate material scenarios.
To increase confidence in the vehicle level CAE model predictions and to better understand
the effect of material and manufacturing variability, prototype molded underbody components
were fabricated, and subsequently built into underbody assemblies to assess their structural
performance. Non-destructive component and assembly tests were devised to assess the
general static and modal performance of the underbody component, and a quasi-static
destructive test of a built-up underbody assembly was developed to simulate the deformation
and loading observed in the worst case vehicle impact design load case.
The paper will discuss the preparation and fabrication of the built-up test assemblies, the
structural stiffness and modal performance testing of trimmed underbody molded components
and assemblies, and the destructive testing of assemblies. The predicted performance was
investigated for two composite thickness assumptions to account for the additional thickness
observed in the prototype components. Predictions were then compared to the measured test
results to understand the status of correlation between the response of idealized components
and the as-molded prototype test components.
A comparison of the non-destructive stiffness and modal test results to the predictions
indicated that the stiffness and modal response were reasonable. The destructive underbody
test was developed to better represent the physical composite and metallic components. The
destructive underbody test was limited by buckling of the longitudinal rail. The results correlated
well with predictions up until rail buckling occurred, after which significant local damage was
introduced into the underbody. The fiberglass SMC fabric composite material and the weld bond
joints were found to very robust in terms of their ability to tolerate large local deformations
without separation.
Additional correlation studies and modeling refinements were underway at the time of this
writing and these results will be reported at a future date.
Background
The Automotive Composites Consortium Focal Project 4 (ACC FP4) is a joint program
between GM, Ford, and Chrysler to develop structural automotive components from composite
materials. Part of this project is to develop a structural composite underbody capable of
carrying crash loads as has been previously reported in [1-4].
A compression molded multi-layered fiberglass fabric reinforced SMC composite [5-9] was
selected as the material and process system [10-11] to best meet the performance and cost
Page 2
objectives of the program. The process of weld bonding was selected as the means to join the
composite underbody to the steel body-in-white (BIW) structure [12-14].
A schematic of the fabric SMC composite underbody structure design and the integration
into the donor BIW via weld bonding is illustrated in Figure 1. The underbody was designed as
single piece molded component to replace 14 major stamped components and partially replace
4 minor stamped components. The underbody assembly was designed to be weld bonded to
the vehicle BIW to compatible with conventional body assembly.
(a) Donor vehicle
BIW and composite
underbody
(b) Weld bonded
joints
Fiberglass
fabric SMC
composite
(a) Donor vehicle
BIW and composite
underbody
(b) Weld bonded
joints
Fiberglass
fabric SMC
composite
Figure 1: Composite underbody design overview
CAE-based design methodologies were utilized throughout the program to assess the
structural performance of the underbody design concept at the vehicle level and to determine
the potential mass reduction. A net mass saving of ~11.3kg (26%) was predicted for the
composite underbody design based on the following structural load cases [1]:
EuroNCAP/IIHS 40 mph frontal offset deformable barrier (ODB)
NCAP 35 mph full frontal impact
FMVSS 214 33.5 mph side impact
FMVSS 301 50 mph rear offset impact
BIW static torsion and bending stiffness
BIW modal response
Several physical tests, including material coupon and sub-element tests, were conducted
throughout the program to determine material properties, validate test-analysis correlation, and
to develop modeling strategies [5-9, 11-13]. In order to evaluate the underbody materials and
processing, and to evaluate the performance for the actual underbody components, a prototype
component mold tool was build and prototype components were fabricated by the ACC team [1,
10].
Page 3
A test plan was developed to conduct non-destructive component and assembly tests, and
also destructive tests of the assembly to simulate the impact performance of the underbody in a
vehicle environment.
The build of the test components and assemblies, the subsequent testing, and the
comparison to analysis predictions conducted to date are discussed in the following. As some of
the test and analysis result interpretations were being completed at the time of this writing, these
are presented as “work-in-progress”.
Test Assembly Component Build
As-molded underbody components were supplied by the ACC to the Multimatic Technical
Centre for component and assembly preparation and build, subsequent testing, and test-
analysis correlation. The underbody components were trimmed according to the 3-D CAD
design using a 5-axis laser to demonstrate the feasibility of this potential production trimming
methodology. The as-molded components, trimmed component design, laser trim fixturing, and
trimming detail are illustrated in Figure 2.
As will be described in a subsequent section, trimmed components were tested using non-
destructive tests only, and trimmed components were built into test assemblies for subsequent
assembly level non-destructive and destructive tests.
X~1647mm
Y~1470mm
Z~445mm
Component fixture / laser trimming Trimmed components
As-molded
components
Laser trimming detail
X~1647mm
Y~1470mm
Z~445mm
X~1647mm
Y~1470mm
Z~445mm
Component fixture / laser trimming Trimmed components
As-molded
components
As-molded
components
Laser trimming detail
Figure 2: Component trimming
During the molding of the prototype underbody components, it was found that the
components were thicker than the design intent [10]. This was verified during the trimming and
preparation of the seven underbody components. The mass summary in Figure 3 indicates that
the mass of the trimmed components varied between 28.2kg and 30.0kg, with a coefficient of
variation of (COV) of 2.3%. The average mass of all prototype underbodies was found to be
5.8kg, or nearly 20%, heavier than the design intent mass of 23.6kg.
Page 4
Trimmed Underbody Mass
30.029.5
30.0
28.6
29.5
28.2
29.5
20.0
22.0
24.0
26.0
28.0
30.0
32.0
121510-5 121510-3 121510-6 121410-1 121510-4 120910-4 121510-8
Underbody ID
Ma
ss
[k
g]
Average
29.3kg
Design
23.6kg
Difference 5.8kg (19.6%)
COV = 2.3%
Trimmed Underbody Mass
30.029.5
30.0
28.6
29.5
28.2
29.5
20.0
22.0
24.0
26.0
28.0
30.0
32.0
121510-5 121510-3 121510-6 121410-1 121510-4 120910-4 121510-8
Underbody ID
Ma
ss
[k
g]
Average
29.3kg
Design
23.6kg
Difference 5.8kg (19.6%)
Trimmed Underbody Mass
30.029.5
30.0
28.6
29.5
28.2
29.5
20.0
22.0
24.0
26.0
28.0
30.0
32.0
121510-5 121510-3 121510-6 121410-1 121510-4 120910-4 121510-8
Underbody ID
Ma
ss
[k
g]
Average
29.3kg
Design
23.6kg
Difference 5.8kg (19.6%)
Trimmed Underbody Mass
30.029.5
30.0
28.6
29.5
28.2
29.5
20.0
22.0
24.0
26.0
28.0
30.0
32.0
121510-5 121510-3 121510-6 121410-1 121510-4 120910-4 121510-8
Underbody ID
Ma
ss
[k
g]
Average
29.3kg
Design
23.6kg
Difference 5.8kg (19.6%)
COV = 2.3%
Figure 3 : Design vs. as-molded underbody component mass
The thickness of all seven underbody components was measured at each trimmed hole
location to gain insight into the thickness distribution relative to the design thickness. The image
on the left in Figure 4 shows the design thickness based on the 0.45mm ply thickness. The
image on the right in Figure 4 indicates the average deviation of all seven components relative
to the design thickness, showing that the average floor thickness was greater than the design
thickness nearly everywhere. The maximum average thickness deviation was found to be
~2.7mm thicker (38%) at the thickest area at the rear of the tunnel (locations T08-T11). The
maximum percentage increase relative to the design intent was ~47% (+2.4mm at locations T02
& T03) at the front of the tunnel.
Several reasons for increased underbody mass and thickness are being investigated by the
ACC team, however it seems plausible that the higher than expected areal weight of the
compounded fabric SMC material contributed to the increased thickness due to the increased
resin content.
Test Assembly Build
A test assembly was designed to simulate the worst case ODB design load case. The
objective was to create a simplified test to induce deformation and loadings similar to those
observed in the full vehicle impact simulation. To achieve a reasonable test configuration, the
test was designed around carry-over components from a donor vehicle BIW. The final test
assembly design is shown in Figure 5 along with the original donor vehicle components.
The test assembly was built-up from modified production BIW sub-components, a trimmed
underbody molding, and custom components including all of the weld bond doublers. The
molded composite underbody components were obtained from the ACC, the BIW components
were provided by GM, and the remaining components were fabricated. All of the components
were sequentially positioned in a custom designed assembly fixture according to a build
process, and then joined via adhesive bonding spot- and MIG welding. An automotive grade
crash-toughened epoxy adhesive was used to bond all components, with 0.75mm diameter
glass beads used to maintain the minimum bond gap. The adhesive was cured according to the
manufacturer’s specifications using a large walk-in oven. The various BIW components and the
assembly fixture used to build the assembly are depicted in Figure 6. The multi-step process
that was used to build each test assembly is illustrated in Figure 7.
Page 5
Figure 4: Design vs. measured average thickness deviation of as-molded components (average of 7 components)
2125mm
1650mm
600mm
Original donor
vehicle components
Test assembly
design
Cross-member
attachments
2125mm
1650mm
600mm
2125mm
1650mm
600mm
Original donor
vehicle components
Test assembly
design
Cross-member
attachments
Figure 5: Test assembly
Page 6
Figure 6: Test assembly build components
Figure 7: Test assembly build procedure
Test and Analysis
A total of six assemblies were built out of the seven composite underbody components that
were trimmed. The test matrix for all component and assembly tests is shown in Table 1. Note
that assembly 121510-6 was not tested to allow it to be used for display, and component
121510-8 was non-destructively tested to provide an additional test data point.
Page 7
Table 1 : Test matrix
Stiffness Modal Stiffness Modal Destructive
120910-4 Assembly X X X X #1
121410-1 Assembly X X - - #3
121510-3 Assembly X X - - #4
121510-4 Assembly X X X X #5
121510-5 Assembly X X - - #2
121510-6 Assembly X X - - -
121510-8 Component X X - - -
AssemblyComponentUnderbody
Component /
Assembly
A finite element analysis (FEA) model was developed for the design intent layup. The layup
model included all design intent ply overlaps specified for the layup of each component which
resulted in the definition 315 individual laminates. A version of the FE model was created to
conduct linear elastic stiffness analyses and to evaluate modal performance of the underbody
using MSC NASTRAN [15]. Another version of the FE model was created to analyze the
nonlinear response in the destructive ODB test using LS-DYNA [16]. Additional thickness and
layup variations were also investigated. Adhesive bonding was modeled in each case using
solid elements, while the sheet metal and composite components were modeled using shell
elements. The composite and adhesive material properties and modeling strategies were based
on properties used in previous studies [11, 12].
Thickness Bounding Cases for Analysis
Ideally, the measured component thickness and actual layup definitions would we used
predict the response of the underbody. However, as seen in Figure 4, the thickness deviation of
the underbody varied throughout the component. An additional complication is that the laminate
thickness is associated with the fiber volume of the laminate, which affects the local laminate
stiffness and strength. Thus, material properties may even vary in the same laminate if the
thickness varies.
As a first step in accounting for the observed differences between the as-molded underbody
thickness and the design intent thickness, two bounding cases were considered to compare
with the test results. The bounding cases are illustrated in Figure 8, where Trial 57 represents
the design intent thickness and layup, and Trial 58 represents an assumed 33% average
increase in ply thickness from 0.45mm per ply to 0.60mm per ply. Note that the ply overlap
regions can be identified by the vertical and horizontal lines in Figure 8. Due to the prototype
nature of the test components, it is recognized that the actual as-molded thickness, layup,
overlap locations, and effective stiffness and strength properties may vary from those assumed
in Figure 8, and may influence the component stiffness, modal, and strength performance.
Page 8
Trial 57 – design thickness Trial 58 – increased ply thickness
Max=5.4mm
Min=1.8mm
Max=7.2mm
Min=2.4mm
Trial 57 – design thickness Trial 58 – increased ply thickness
Max=5.4mm
Min=1.8mm
Max=7.2mm
Min=2.4mm
Figure 8: Analysis model thickness assumptions for bounding cases
Component and Assembly Stiffness Performance
Test Setup
Both molded components and built-up test assemblies were quasi-statically loaded in
bending and torsion to measure the static stiffness values. Displacements were measured at
various locations using string pots for the purpose of calculating the required stiffness values.
The physical test setup for the components and assemblies is shown in Figure 9. The
associated FE model setup is shown in Figure 10.
The physical test setup consisted of a stiff beam at the forward and aft ends of the
underbody. The forward beam was mounted to a pivot in the center of the beam, while the aft
beam was fixed. The test components were mounted upside-down in the fixture to facilitate
loading, with the component ends attached to the stiff beams via spherical joints. For the
component tests, loading plates were bolted to the composite component to distribute the
loading. For the assembly tests, load was introduced at attachment points to the front of the
longitudinal rails and sides of the rear rails.
In the case of torsion, a load was applied to the outboard end of the forward beam to induce
a rotation at the front of the floor. For the case of bending, the forward pivot was locked to inhibit
rotation, and a vertical load was applied to the loading bar at the center of the floor to induce
global bending. In each case, the average static torsion and bending stiffness was calculated
from the input force and the resulting rotation or bending displacement, respectively. Each
component and assembly was tested three times to obtain an average value for the torsion and
bending stiffness.
Page 9
Figure 9: Component and assembly stiffness test setup
(a) Component test setup (b) Assembly test setup(a) Component test setup (b) Assembly test setup
Figure 10: Component and assembly stiffness FEA model setup
Results
The static component mass and stiffness results are summarized in Table 2. The mass and
stiffness results are averaged for all tests, and the results were seen to be consistent between
tests with a COV of less than 10% for all data, and a maximum individual test deviation of 16%
relative to the average. The components that were subsequently tested in assembly are
indicated in the table.
Page 10
Table 2 : Component mass, bending, and torsion stiffness results
120910-4* 121410-1 121510-3 121510-4* 121510-5 121510-6 121510-8Test
AverageCOV
Mass [kg] 28.2 28.6 29.5 29.5 30.0 30.0 29.5 29.3 2.3%
Bending [N/mm] 2341 2016 2279 2005 2510 2001 2002 2165 9.7%
Torsion [N-m/deg] 66.2 55.1 64.4 57.2 62.6 63.0 63.1 61.6 6.5%
*Also tested in assembly
Component Test
The average test results from Table 2 are compared to the bounding case FEA predictions
in Table 3. Is can be seen that the average test results fall in between the predicted bounding
case results indicating that the bounding results are reasonable given the variation in
component thickness.
Table 3: Comparison between average component stiffness test results and bounding case predictions
FEA Tr57%Diff vs.
Test Avg.FEA Tr58
%Diff vs.
Test Avg.
Mass [kg] 29.3 23.6 -20% 30.6 4%
Bending [N/mm] 2165 1904 -12% 2658 23%
Torsion [N-m/deg] 61.6 49.8 -19% 70.1 14%
Increased ThicknessDesign ThicknessTest
Average
The static assembly stiffness test results are summarized in Table 4. The stiffness results
were seen to be consistent with an individual test difference of less than 6% relative to the
average for both tests. Predicted stiffness results for the assembly were not available at the
time of this writing.
Table 4: Assembly bending and torsion stiffness results
120910-4 121510-4Test
AverageCOV
Bending [N/mm] 975 1029 1002 3.8%
Torsion [N-m/deg] 564 629 597 7.7%
Assembly Test
Component and Assembly Modal Performance
Test Setup
The modal performance was evaluated for the seven molded components and the two built-
up test assemblies using “free-free” boundary conditions. The basic test setup for the
components and assemblies is shown in Figure 11. In the test, the component or assembly was
suspended from a stiff frame using bungee cords. The component or assembly was then
randomly excited in the normal “Z-direction” using a shaker and triaxial accelerometers were
used to map out the mode shapes and the associated frequencies. The component bending
Page 11
and torsion FEA model used to predict the component stiffness response were the same as the
models shown in Figure 10 but with free-free boundary conditions applied (i.e. without any
fixturing). The measured mode shapes were visualized and compared to the predicted mode
shapes for comparison.
Figure 11: Assembly and component modal test setup
Results
The predicted bounding results are provided for reference as the test-analysis correlation
was in process at the time of this writing.
The primary bending and torsion mode results are summarize in Table 5 for the molded
components. The modal results were generally consistent with the exception of the torsion
response of sample 121510-3, which exhibited a higher frequency value (12.4 Hz) than all other
samples. The mode shape for this sample appeared to be different than the other samples,
possibly explaining the difference. The components that were subsequently tested in assembly
are indicated in the table.
Table 5: Component bending and torsion modal results
Component Test
120910-4* 121410-1 121510-3 121510-4* 121510-5 121510-6 121510-8Test
AverageCOV
Mass [kg] 28.2 28.6 29.5 29.5 30.0 30.0 29.5 29.3 2.3%
Bending [Hz] 23.5 24.2 24.8 24.0 23.5 24.8 23.7 24.1 2.4%
Torsion [Hz] 8.8 9.5 12.4 7.8 9.6 8.2 7.7 9.2 17.8%
*Also tested in assembly
The average test results from Table 5 are compared to the bounding case FEA predictions
Page 12
in Table 6. The associated first bending and torsion mode shapes for the “design thickness”
bounding case (Trial 57) are shown in Figure 12, where the shapes in the figure represent the
deformed shapes at the maximum positive and negative amplitudes. The predicted mode
shapes were observed to be very similar to the measured mode shapes (not shown). The
average test results for bending are closer to the “increased thickness” Trial 58 results, while the
torsion results lie closer to the “design thickness” Trial 57 results.
Table 6: Comparison between average component modal test results and bounding case predictions
FEA Tr57%Diff vs.
Test Avg.FEA Tr58
%Diff vs.
Test Avg.
Mass [kg] 29.3 23.6 -20% 30.6 4%
Bending [Hz] 24.1 19.2 -20% 21.8 -9%
Torsion [Hz] 9.2 9.1 -1% 10.1 10%
Design Thickness Increased ThicknessTest
Average
The modal component FEA analyses were found to be sensitive to thickness distribution,
which not only would affect the local stiffness, but also the local mass distribution. Given the
differences between the actual components and the bounding assumptions, the mode shapes
and frequency results for the 1st bending and torsion frequencies were found to be reasonable,
with the predicted frequencies within 10% of the measured values for the case of increased
thickness (Trial 58).
Figure 12: Predicted component 1st torsion and bending mode shapes (Trial 57)
The modal test results for two assemblies are summarized in Table 7. The test results were
Page 13
very consistent with the magnitude of the assembly bending frequency results being similar to
the component test results, and the magnitude of the torsion frequencies nearly twice that of the
component tests.
Table 7: Assembly bending, and torsion modal results
120910-4 121510-4Test
AverageCOV
Bending [Hz] 23.6 23.4 23.5 0.4%
Torsion [Hz] 16.3 16.9 16.6 2.7%
Assembly Test
The average test results from Table 7 are compared to the bounding case FEA predictions
in Table 8. The associated predicted first bending and torsion modes for the “design thickness”
bounding case (Trial 57) are shown in Figure 13. The predicted mode shapes were observed to
be very similar to the measured mode shapes (not shown). The average bending test results
are higher than the predicted results, while the average torsion test results are lower than the
predicted results.
Figure 13: Predicted assembly 1st torsion and bending mode shapes (Trial 63)
Table 8: Comparison between average assembly modal test results and bounding case predictions
FEA Tr63%Diff vs.
Test Avg.FEA Tr64
%Diff vs.
Test Avg.
Bending 23.5 20.0 -15% 20.7 -12%
Torsion 16.6 19.3 16% 20.5 23%
Test
Average
Design Thickness Increased Thickness
Page 14
ODB Destructive Testing
Test Setup
A test setup was developed to subject the underbody to a simulated worst-case loading in
the vehicle environment. The setup was designed to apply deformations similar to those that
are predicted to occur during a full vehicle ODB impact loading.
As can be seen in Figure 14, the underbody was clamped at the rear side edges and the
rear of the floor while loads were applied simultaneously via hydraulic cylinders to the driver side
front rail and to the front of the transmission housing. The rear side attachments were made by
welding sleeves into the rocker sections, and then bolting through these into a 25mm thick steel
plate that was attached to bed plate. The rear of the floor was bolted to stiff box beam which
was attached to the bed plate.
The loading was applied quasi-statically using load control at a ratio of 2.375:1 and at a
specified initial loading angle, with the displacement of the loaded end of the driver side rail
constrained by a pivot link pinned on both ends. Additional details of the setup can be seen in
the photos shown in Figure 15. The loads were applied at specified angles through spherical
joints. The transmission housing was only attached to a bushed cross-member that was bolted
to the composite floor (see Figure 5). Therefore, the motion of the transmission was controlled
by the initial transmission load application direction, the rotation defined by the cross-member
bushings, and the eventual contact between the transmission housing and the tunnel at higher
loads. Loads were measured at the load application points with load cells and displacements
were taken from the stroke of the hydraulic actuators. Displacements and surface strains were
measured at specific locations in some tests to monitor the response during loading.
Results
A summary of the measured rail and transmission force vs. displacement responses
obtained from all five ODB tests are shown in Figure 16. The limiting factor is noted for each
test, either premature rail buckling or failure of the cast aluminum cross-member.
For test #1 (120910-4), the maximum input load was limited by the local buckling of the
loaded rail at the front of the floor. Additional analytical test development and correlation was
conducted after test #1 to improve subsequent tests and to maximize the load input capacity
into the underbody. During this development, it was found that the analysis model did not
adequately represent the connection and contact between the longitudinal rails and the internal
foam core. As a result, the model predicted a higher rail buckling force than was observed in the
test. To help compensate for the lower rail buckling load, the load application angles were
adjusted to minimize the lateral component of the rail load input.
Based on test #1, the revised test setup was used for tests #2 (121510-5) and #3 (121410-
1) with the result that the maximum input load was limited by the failure of the cast aluminum
transmission cross-member. For tests #4 (121510-3) and #5 (121510-4), which have been re-
plotted in Figure 17 for clarity, the cast aluminum transmission cross-member was reinforced
with a 25mm thick steel plate (shown in Figure 15), which enabled a higher underbody load
input due to suppression of the cross-member failure mode. Both tests #4 and #5 were very
similar in terms of overall response and maximum load. A small difference can be seen in the
slope of the load-displacement response in Figure 17, where it appears that test #5 is stiffer
above ~18kN. This difference has been attributed to the motion of the transmission housing and
the timing of the contact with the tunnel.
Page 15
Figure 14: ODB destructive test setup
Figure 15 : Rail and transmission loading details
As noted previously, another key factor already identified as having an influence on the
response of the underbody was the observed differences between the design and the
measured as-molded thickness. Areas with variations in thickness will affect the local
component stiffness and strength, which may further alter the load path and failure modes.
Analytical investigations have shown that the failure locations and failure modes can change as
a result of variations in the local part thickness, stiffness, and strength. Studies were ongoing at
the time of this writing.
Page 16
Figure 18 shows several video frames that depict the failure sequence in test #4. Figure (a)
illustrates that there is no damage during the initial elastic loading of the underbody. Figure (b)
shows the appearance of local composite damage at several locations as a result of the
initiation of rail buckling at the left side of the dash panel. Figure (c) shows significant upward
rotation of the rail leads to significant local composite damage near the rail attachment and also
in the tunnel area. Figure (d) indicates that significant cracking around the tunnel occurs near
the end of the test when the rail rotation is at a maximum. However, as can be seen in the post
test images shown in Figure 19, net section failure was not observed in the composite despite
the significant cracking near the rail attachment and around the tunnel. Additionally, it was
observed that the weld bond doublers performed well in terms of maintaining joint integrity
between the steel and composites components.
Figure 20 compares several predicted results to the measured rail load vs. displacement
response for tests #4 and #5. The predicted results represent three different material
assumptions as indicated in the figure. Trial 1378 and 1386 are identical other than that the
composite material stiffness and strength values are reduced for Trial 1386 based on coupon
tests conducted from in-situ tension and compression samples that were taken from test #1.
Trial 1385 is based on increased material thickness values measured from test #5 and also
uses material stiffness and strength properties that take into account the local material
thickness from samples taken from test #1. All predicted results represent the measured
loading curves well up until rail buckling occurs, with the predicted responses starting to deviate
from one another above approximately 45kN based on the various material assumptions.
Figure 16: Load-displacement response for all destructive tests
Page 17
Figure 17: Load-displacement response for destructive tests #4 and #5
Figure 18: Destructive test #4 (121510-3)
Page 18
Significant local
damage due to
rail buckling
Local damage
around tunnel
Significant local
damage due to
rail buckling
Local damage
around tunnel
Figure 19: Observed post-test damage for destructive test #4 (121510-3)
Figure 20: Comparison between measured and predicted force-displacement response
Summary
A composite underbody was designed to perform within a donor vehicle BIW environment
based on a CAE approach. Prototype composite underbody components were molded and test
components were fabricated and subjected to both non-destructive component and assembly
tests, and destructive assembly tests. The destructive test was devised to subject the
underbody test assembly to the worst case design load case. A test assembly was designed
Page 19
and fabricated from composite underbody components and modified production BIW
components. It was confirmed that the thickness and mass of the underbody components
provided by the ACC was greater than the assumed design values.
FEA predictions were compared to test results to understand the status of correlation to the
as-molded components. Bounding studies indicated that the measured non-destructive
component and assembly stiffness and modal response were reasonable. The destructive
underbody test was developed to better represent the physical components and as-molded
material thickness values. The predicted force vs. deflection response of the destructive
underbody test was found to give a good representation of the tests up until rail buckling
occurred.
Overall, it was observed that the load carrying capacity of the as-molded composite
underbodies exceeded the load carrying capability of the adjacent metallic components.
Further, it was found that the fiberglass SMC fabric composite was very robust and was able to
tolerate large local deformations by cracking but without separating. Lastly, it was observed that
the weld bond doublers performed well in that the integrity of the composite-steel joints was
maintained.
All underbody testing was completed with correlation efforts ongoing at the time of this
writing. Therefore, final correlation results will be reported at a later date.
Acknowledgments
The authors would like to thank the ACC Underbody team for their valuable inputs, and in
particular Libby Berger (General Motors), Bhavesh Shah for facilitating the BIW components
(General Motors), Chuck Knakal (USCAR), and Dan Houston (Ford Motor Company).
This material is based upon work supported by the Department of Energy National Energy
Technology Laboratory under Award Numbers DE-FC26-02OR22910 and DE-EE0003583. This
report was prepared as an account of work sponsored by an agency of the United States
Government. Neither the United States Government nor any agency thereof, nor any of their
employees, makes any warranty, express or implied, or assumes any legal liability or
responsibility for the accuracy, completeness, or usefulness of any information, apparatus,
product, or process disclosed, or represents that its use would not infringe privately owned
rights. Reference herein to any specific commercial product, process, or service by trade name,
trademark, manufacturer, or otherwise does not necessarily constitute or imply its endorsement,
recommendation, or favoring by the United States Government or any agency thereof. The
views and opinions of authors expressed herein do not necessarily state or reflect those of the
United States Government or any agency thereof.
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