In-Plane Shear Characterization of Sandwich
Laminates Using a Picture-Frame Test Configuration
F. C. STOLL and N. G. JOHNSTON
ABSTRACT
An investigation was conducted to characterize the accuracy of “picture-frame” test
configurations for measuring the in-plane shear strength of sandwich constructions as
may be limited by local instability of the face sheets including wrinkling and dimpling.
Two configurations were studied: a conventional configuration featuring corner pins
which extend between the two sides of the specimen and fixture, and a modified
configuration in which the corner pivot axes coincide with the corners of the square
gage area bounded by edge doublers. Shear strains were predicted using finite element
analysis and measured experimentally with strain gages and digital image correlation.
The conventional configuration produced shear strain at the specimen center
approximately 15% below the nominal value predicted for uniform shear stress, and the
strain was significantly non-uniform over the gage area. The modified configuration
provided a much more uniform shear strain distribution, at a strain level corresponding
closely to the nominal value. Ultimate strength measurements are provided for four
sandwich designs featuring both plain foam core and web reinforced core, tested in both
configurations. Practical details of the hardware implementation for the modified
configuration are presented.
INTRODUCTION
The cross section of a typical utility-scale wind turbine blade reveals an outer shell
which provides the external airfoil shape, and one or more internal “shear webs” which
are the primary load-carrying elements with respect to flap-wise shear loads induced by
the aerodynamic forces on the blade. Shear webs feature sandwich construction to
provide resistance to global (panel-level) buckling. The sandwich face sheets are fiberreinforced composite laminates, generally reinforced with double-bias [+-45] E-glass
fiber architecture in consideration of the in-plane shear load.
One potential sandwich failure initiation mode under in-plane shear is local buckling
of the face sheets. With quasi-homogeneous core products such as structural foams,
_____________
Frederick C. Stoll, Nathan G. Johnston, Milliken & Company, 920 Milliken Rd., M-169,
Spartanburg, SC 29303, U.S.A.
local buckling due to compression or shear takes the form of face sheet wrinkling, which
can be predicted with sufficient accuracy using an equation of the form
(1)
where
is the (compressive) face wrinkling stress, is the modulus of the face,
are the Young’s modulus and shear modulus of the core material, respectively,
and
and k is a theoretical or empirical coefficient [1]. However discretely reinforced core
products have entered the wind market (Figure 1) [2], introducing the possibility of face
sheet dimpling as the local buckling mode, for which there is no accurate general closedform analytical solution. An experimental method is needed to provide accurate values
for local in-plane shear strength as limited by face sheet dimpling.
The “picture-frame” test configuration (Figure 2) has been applied to investigate inplane shear response of composite laminates, sandwich constructions, dry fabrics, and
foams [3-8]. Although the configuration requires a relatively complex specimen and
fixture design, it provides the possibility of producing an approximately uniform shear
stress/strain field on large specimens with the application of a uniaxial load. While a
picture frame method was recently standardized for determining the elastic shear
properties of composite laminates [8], there is currently no international standard
method for sandwich constructions, although as of the time of writing, ASTM
Subcommittee D30.09 has a draft method under ballot. However a number of
uncertainties have previously been expressed about preferred details and accuracy of
picture frame testing [5] which, to the knowledge of the author, have not been fully
resolved in the literature.
This paper presents the results of an experimental and analytical investigation of
picture-frame test configurations applicable to sandwich laminates typical of shear webs
of wind turbine blades. The preferred attributes include the following:
1. Create a uniform shear stress/strain field of known value
2. Employ a large gage area to allow local buckling of the face sheets with
negligible influence from boundary effects
3. Minimize local stress concentrations at the edges or corners to avoid improper
failure modes.
A convention picture frame configuration applied to large (~230mm gage
dimension) plate and sandwich constructions is shown in Figure 2. A square frame on
y
Figure 1. Web-reinforced sandwich core used in
wind turbine blades [2].
x
Figure 2. Tension-loaded “picture frame”
configuration for in-plane shear testing.
each side of the test specimen is comprised of four rigid bars which feature a series of
holes for mounting the specimen with through-bolts. Doublers (tabs) are often bonded
to the specimen edges to reduce bearing stresses. At the four corners, large-diameter
pins extend between the two sides of the specimen, defining the corner pivot axes,
serving to align the two sides of the fixture, and providing load introduction points at
two corners. A notch must be cut at each corner of the specimen to clear the corner pins,
preventing the use of the kinematically ideal gage area which would consist of a square
with corners coinciding with the corner pivot axes. This results in uncertainty about the
best placement and extent of the doublers, and the optimal shape of the corner notches.
Findings on preferred configuration details for uniform shear stress/strain
distribution have been reported in the literature [5]. Among the key findings: (1) Not
surprisingly, for a rectangular gage area bounded by doublers, the four corner pivot axes
should coincide with the corners of the gage area; and (2) The doubler should be very
stiff (in-plane) compared to the test laminate. Item (1) requires a departure from the
conventional configuration, and serves as the basis for a second configuration
considered here given the goal of uniform shear stress in the gage area.
The following section contains detailed descriptions of the baseline (conventional)
test configuration, and a modified configuration featuring the preferred corner pin
location. An analytical and experimental program to study the shear strain field is
documented, employing finite element analysis, and strain measurements using both
strain gages and digital image correlation. Strength measurements for four different
sandwich laminates are compared for the two configurations. Details of a test fixture
developed in the modified configuration are provided.
PICTURE FRAME TEST CONFIGURATIONS
Baseline (conventional) and modified test configurations are described here. Both
configurations used a common specimen planform shown in Figure 3(a). The overall
specimen dimension is 305mm (12”) square, with 38mm (1.5”) wide bonded edge
doublers, defining an interior gage area 229mm (9”) square. A 38mm (1.5”) square
notch was cut at each corner.
The baseline fixture featured frame members 25.4mm (1”) thick with planform
dimensions shown in Figure 3(b). Each bar was machined down to 12.7mm (0.5”) thick
where two bars overlap at the corners. Each bar featured 17 holes, 6.4mm (0.25”)
diameter, in two rows, though only 15 were used due to interference at the ends. The
corner pin diameter was 25.4mm (1”). All frame members and pins were steel. The
baseline configuration was used with bonded doublers 6mm (0.25”) thick. Doubler
materials used in the investigation included both FRP (fiber-reinforced polymer)
composite plates with the same reinforcement as the face sheets, and steel.
In the modified configuration (Figure 3(c)), the same general fixture dimensions
were used except the corner pivot holes were moved to achieve 228.6mm (9”). Pin
spacing. Accommodations were required to compensate for the absence of the four
corner through-pins. At the load points, a stiff assemblage was outfitted with 25.4mm
diameter studs which extend into the corners of the fixture bars in a clearance fit (Figure
3(d)). At each passive corner, one bar featured a 25.4mm diameter stud, mounted by
press fit, which extended into the neighboring bar with a clearance fit. A set of bonded
a) Planform of specimen configuration for
method study.
b) Planform dimensions of the baseline
fixture frame.
228.6mm
(9.00”)
c) Modified fixture frame.
d) Load fixture for modified configuration.
Figure 3. Details of specimen and test fixtures.
steel doublers was used, each 12.7mm (0.5”) thick featuring blind threaded holes. Other
fixture details are described in a later section.
ANALYTICAL AND EXPERIMENTAL STUDY
An analytical and experimental study of the two test configurations was performed
consisting of the following elements:
• Uniaxial testing of the face laminate material to obtain in-plane stiffness
parameters.
• Finite element analyses (FEA) of the shear strain distributions for the two
configurations.
• Experimental shear strain measurements, including strain gages at the specimen
center, and digital image correlation (DIC) to examine the distribution of strain
over the gage area.
A single sandwich laminate configuration was used for comparing results. The core
material was 25mm thick Diab Divinycell H60 PVC foam. The foam was perforated
with 3mm diameter holes in a 50mm grid pattern for through-infusion. The face sheets
were reinforced with two plies of Saertex 830 g/m2 double bias [-45/+45] E-glass noncrimped knitted fabric, product style U32EX020 00830 01270 264000. The molding
resin was Momentive EPIKOTE™ MGS® RIMR 135 epoxy resin and Epikure MGS®
RIMH1366 curing agent in a 100:30 mass ratio. Vacuum infusion molding process
under high vacuum was used, with initial cure at 50% vacuum. A final cure of 70°C for
eight hours was applied.
Laminate Characterization
Accurate values for elastic properties of the face sheets were sought to provide a
good basis for analytical and experimental comparisons. Instrumented uniaxial tests of
the face laminate material were performed using ISO 527 [9]. Axial strain was measured
using an extensometer clipped to the specimen, and transverse strain was measured
using uniaxial strain gages applied to the two sides of the specimen.
Testing was performed on two different laminate configurations:
1. 0° loading: Three-ply laminates were molded in a [0]3 fabric configuration
providing a fiber angle configuration of [-45/+45]3 with respect to the load
direction. The test configuration prevented twisting of the specimens.
2. 45°loading: Four-ply laminates were molded in a [+45/+45/-45/-45] fabric
configuration providing a fiber angle configuration of [0/90/0/90/90/0/90/0]
with respect to the load direction.
Five specimens of each configuration were tested. The output for each specimen
consisted of a stress versus strain curve from which Young’s modulus was determined,
and a curve for transvers strain versus axial strain for determining the Poisson ratio.
Whether for the 0° or 45° reference axis, the complete set of in-plane orthotropic
,
ν ] can be reduced for these laminates to three
elastic constants [ ,
independent parameters recognizing that
(2)
For a 45° transformation from one reference axis to the other, the transformed
parameters [ , ̅ , ν ] can be obtained with the following equations:
̅
ν
(
ν )
(
ν )
(
+
(3)
−
(4)
ν )
⁄
(1 − ν )
(5)
(6)
(7)
TABLE I. ELASTIC CONSTANTS FOR FACE SHEETS, !" 50%
Measured
Eqn. 5
Eqn. 6
Eqn. 7
Test/
E11
G12
E11
ν12
ν12
Reference
Direction
MPa
MPa
MPa
0 deg
10,830
0.56
10,730
0.60
11,130
45 deg
24,200
0.127
3,480
0.072
23,020
Because
is not directly measured in either reference axis, it must be determined for
each reference axis using Equation 5 applied to the measurements from the other
reference axis.
Both directly-measured and transformation-derived values are summarized in Table
I. Measured values for
correspond to a reference fiber volume fraction (! ) of 50%,
which was within 1% of the actual values based on weight analysis. The stress-strain
response becomes highly nonlinear for the 0° reference direction for tensile strain
beyond 0.4%, so
was evaluated between 0.1% and 0.4% strain, and ν was
evaluated at 0.3% strain.
Some differences can be observed between the directly measured parameters and
the transformation-derived values, though the two Young’s modulus values agree
within 5% for both reference directions. For the in-plane shear loading of interest, the
elastic response is governed largely by laminate stiffness in the fiber axe directions
(±45°). When specifying properties for analysis with respect to the material reference
axis (0°), fiber-dominated behavior is best represented by the Table I parameters within
the bold border. These were used in FEA.
Finite Element Analysis
The FEA model used some approximations compared to the physical test
configurations. Symmetry about the mid-plane of the sandwich was used. The test
specimen model is depicted in Figure 4(a). The face sheet/doubler stack at the edges of
the specimen was modelled using shell elements. The model including the frame fixture
is shown in Figure 4(b). The holes, mounting bolts, and pivot pins were not explicitly
modelled, rather multi-point constraints were used to tie together the in-plane
displacements of fixture and specimen nodes corresponding to the central axis of each
mounting bolt and corner pin. The baseline and modified configuration models were
identical except for the location of the corner pivot axes, and the doubler materials and
thickness. The core material and fixture frame were modelled as homogeneous isotropic
solids. The face sheet was modelled in terms of layers of the double bias fabric
reinforcement, with reference to the 0° fabric angle, and the in-plane elastic constants
were those in the bold box of Table I with a ply thickness of 0.643mm. The corner
tension load was applied to one simulated corner pin at a 45° angle, with the opposite
simulated pin fixed with respect to in-plane displacements. Node constraints were
applied to one surface of the fixture bars against out-of-plane displacements. Linear
static analysis was performed. The strain predictions at the center of the gage area under
80 kN applied load are tabulated in Table II.
P
P
Face sheet gage area
Core
Face sheet plus edge
doublers (shell elements)
a) Specimen model (symmetry used)
Fixture
b) Full FEA model with fixture (symmetry used)
Figure 4. Components of the FEA model.
TABLE II. FEA PREDICTIONS OF STRAIN AT THE SPECIMEN CENTER, ,
Configuration
Baseline, no doubler
Baseline, composite doubler
Baseline, steel doubler
Modified
ε (45°)
0.36%
0.32%
0.28%
0.43%
ε (−45°)
-0.34%
-0.31%
-0.27%
-0.43%
80 .
γxy
0.70%
0.63%
0.54%
0.86%
Experiments
A series of test runs was performed on the sandwich design described earlier using
the two fixture configurations. Strain gage instrumentation consisted of pairs of strain
gages mounted near the geometric center on lines of symmetry, one at 45° (tension axis)
and one at -45° (compression axis). Specimens without Digital Image Correlation (DIC)
featured strain gages on both sides. Specimens used with DIC had strain gages only on
the side opposite the DIC preparation. DIC imaging was performed using Correlated
Solutions (Columbia, SC, USA) VIC-3D and VIC-2D systems and software.
Specimens were prepared with a white background and dark speckle pattern. Several
runs were performed, some using 3-D DIC (two cameras), some using 2-D (one
camera). The DIC strains reported for the center of the gage area are average values
over a ~30mm square area.
Multiple test runs were performed on nominally identically specimens, and in some
cases two test runs were performed on an individual specimen. These duplications were
performed while working out problems with the two strain measurement systems, and
verifying unexpected results. Data sets with known problems are omitted, all others are
reported. Results for strain measurements at the specimen center are reported in Table
III. For strain gage instrumentation, shear strain was obtained from the two gage
readings as follows:
&'(
)(45°) − )(−45°)
(8)
When both sides were instrumented, the results from the two sides are averaged. Test
results will be discussed in a following section.
TABLE III. MEASUREMENTS OF STRAIN AT THE SPECIMEN CENTER, ,
Run
1
2
3
4
5
6
7
8
9
10
11
12
Config.
Specimen ID
Baseline,
composite
doubler
DF1007C-1
DF1007C-2
DF1007C-3
DF1265A
DF1327A
DF1327A
DF1370A
AVERAGE
Baseline, steel
doubler
DF1265B
Modified
DF1265C
DF1327C
DF1327C
DF1370C
AVERAGE
Strains at the Specimen Center
Strain Gage
Placement
ε (45°)
ε (−45°)
γxy
2 sides
0.41%
0.37%
0.40%
-0.33%
-0.34%
-0.34%
Invalid
-0.33%
-0.35%
-0.30%
-0.33%
0.74%
0.72%
0.74%
0.39%
0.37%
0.38%
0.39%
1 side
0.72%
0.71%
0.68%
0.72%
Invalid
1 side
0.39%
0.39%
0.40%
0.39%
1 side
Invalid
-0.41%
-0.44%
-0.46%
-0.44%
0.80%
0.83%
0.86%
0.83%
ε (45°)
80 .
DIC
ε (−45°)
γxy
Not used
0.37%
0.50%
0.35%
0.36%
0.39%
-0.33%
-0.26%
-0.38%
-0.43%
-0.35%
0.70%
0.76%
0.74%
0.79%
0.75%
0.37%
-0.32%
0.69%
0.43%
0.36%
0.40%
0.47%
0.41%
-0.43%
-0.55%
-0.43%
-0.44%
-0.47%
0.86%
0.90%
0.83%
0.91%
0.88%
Analysis of Results
SHEAR STRAINS AT THE CENTER OF THE GAGE AREA
Analytical and experimental results for shear strain at the center of the gage area
under 80 . applied load are analyzed here. Shear strain values are normalized by a
reference (ideal) shear strain, &'(01234 , obtained as follows.
The nominal running shear load on the gage area, .'( , is given by:
5
.'(
√ 47
0.2475;./=
(9)
where , 0.080;.is the applied load and >? 0.2286= is the edge length of the
gage area. Computation of the ideal shear strain takes account of the core shear modulus
and thickness,
20;,A and B
0.025=, respectively; the face sheet shear
modulus, ply thickness, and number of plies per face,
=10,730 MPa , BC4(
0.643 × 10 =, and .C4( 2, respectively; and the shear modulus and thickness of
a layer of resin absorbed into the open cells at the surfaces of the core,
1054;,A
and B
0.4 × 10 =, respectively:
&'(01234
FGH
IJ
IK LK I
FMNH LMNH
0.855%
(10)
Normalized shear strain values at the specimen center from FEA, strain gage, and
DIC are plotted in Figure 5. For the baseline configuration with FRP doublers, the
normalized shear strain values from FEA, strain gage average, and DIC average, were
0.73, 0.84, and 0.87, respectively. For one specimen tested with 6mm thick steel
Normalized shear strain, γxy/γxy ideal
1.1
Key: Config/Doubler, Method
1.0
Baseline/None, FEA
Baseline/FRP, FEA
0.9
Baseline/Steel, FEA
Baseline/Steel, SG
0.8
Baseline/FRP, DIC
Baseline/Steel, DIC
Modified/Steel, FEA
0.7
Modified/Steel, SG
Modified/Steel, DIC
0.6
0
2
4
6
Figure 5. Results for normalized shear strain at the center, ,
80 ., &'( 01234 = 0.855%.
doublers, the normalized shear strain value from DIC was 0.80, significantly less than
the average value of 0.87 for FRP doublers. The doubler design had a large effect on
FEA predictions; the normalized shear strain values of 0.82, 0.73, and 0.64 were
obtained with no doublers, FRP doublers, and steel doublers, respectively. The result
with no doublers agreed best with the experimental values measured with FRP doublers,
which suggests that an unrealistically stiff representation of doubler effectiveness
resulted from the FEA modelling approach, in which doublers were modelled with shell
elements along with the face sheets. Both the FEA and DIC results indicate that for the
baseline configuration, increasing the doubler stiffness reduces the accuracy of the
center shear strain compared to the ideal value.
For the modified configuration, the center shear strain values were very close to
ideal. The normalized values for FEA, strain gage average, and DIC average, were 1.00,
0.97, and 1.03, respectively.
STRAIN DISTRIBUTION IN THE GAGE AREA
Both FEA and DIC were used to visualize the distribution of shear strain over the
gage area. FEA fringe plots of normalized shear strain in the gage area are presented in
Figure 6 for four analysis cases. For the baseline configuration, regardless of the doubler
design, the shear strain was significantly higher at the edges of the gage area than at the
center (Figure 6(a-c)). For the baseline configuration with no doubler, the strain at the
corners was lower than at the center (Figure 6(b)), whereas with doublers the corner
strains were elevated compared to the center. For the baseline configuration, the model
with no doublers provides the best uniformity of shear strain, and closest agreement
with ideal. For the modified configuration, the normalized shear strain was much more
uniform and close to unity over the gage area, with only small regions of slightly lower
shear strain at the corners.
DIC fringe plots for normalized shear strain at 80 kN load are presented in Figure 7
for both the baseline configuration with FRP doublers and the modified configuration.
a)
Baseline config., FRP doubler
b) Baseline config., no doubler
c)
Baseline config., steel doubler
d) Modified config.
Figure 6. FEA fringe plots of normalized shear strain, &'( /&'( 01234 , in the gage area.
a) Baseline configuration, FRP doublers
(DF1370A)
b) Modified configuration (DF1370C)
Figure 7. DIC fringe plots for normalized shear strain, &'( /&'( 01234 , for , = 80 . .
(Note that the fringe plots do not cover the entire gage area because the frame interior
is 213mm across compared to the 229mm gage dimension, and the plots do not span the
entire visible gage area.) The baseline configuration has a fringe scale range of 0.47,
almost twice that of the modified configuration (0.27). For the baseline configuration
the edges of the gage area show elevated strain compared to the center, whereas the
corners show reduced strain. This is qualitatively similar to the FEA fringe plot for the
baseline configuration with no doubler (Figure 6(b)), further suggesting that the FEA
representation of the edge doublers may be overly stiff compared to reality.
The modified configuration shows an unexpected gradient of shear strain from left
to right. This was observed with both Specimens DF1327C and DF1370C. This gradient
is not consistent with the lines of symmetry in the test configuration, therefore it does
not imply a shortcoming of the basic configuration, but rather a deficiency in the
execution, such as rotation of the specimen during loading, or shortcomings in the test
fixture (see Details of the Modified Test Fixture Design, below). Despite the apparent
gradient, the indicated strain is within ±10% of nominal over most of the gage area, and
is much more uniform and closer to the ideal than the baseline configuration.
STRENGTH MEASUREMENTS
Test results for ultimate strength are presented here for four sandwich designs. Two
different core types were used:
• PVC foam, consisting of 50mm thick 60 kg/m3 PVC foam (either Diab
Divinycell H60 or 3A Composites Airex C70.55)
• Web core, consisting of 50mm thick sandwich core with discrete fiberglass
composite webs surrounded by low-density polyisocyanurate foam (Milliken
TYCOR W0.1-RB core).
The face sheet reinforcement, epoxy molding resin, and processing were the same as
described above for the experimental program. For each of the core types, two different
sandwich laminates were fabricated, featuring 1-ply and 2-ply face laminates,
respectively. Three specimens from each design were tested in each test configuration.
The results for average ultimate applied load are plotted in Figure 8. Example
ultimate failure modes are shown in Figure 9 for the modified test configuration. Failure
propagated very rapidly over the height of the gage section, so the initial location and
mode of failure are not known, although analytical work not reported here suggests that
face sheet stability was near critical, if not critical, at the failure loads.
For 1-ply and 2-ply web core panels, the ultimate load decreased 15% to 17%
between the baseline and modified configurations. This is consistent with the higher
ratio of shear strain (and stress) to applied load in the center region of the gage area for
the latter configuration compared to the former. For the PVC foam core with 1-ply face
sheets, average ultimate load increased 13% between baseline and modified
configurations. A possible explanation is that is that local stress concentrations present
in the baseline configuration may have initiated early failure compared to the modified
configuration in which they were not present.
The advantage of the modified configuration, based on the investigation reported
here, is that the ultimate load values can be converted with confidence to values for the
ultimate panel shear strength, .'( O4L , or the ultimate face shear strength, P'( O4L , for
use in the structural design process.
250
-15%
-2%
Ultimate load P, kN
200
150
+13%
100
-17%
Baseline
config.
50
Modified
config.
0
PVC foam core TYCOR W0.1- PVC foam core TYCOR W0.1RB core
RB core
1-ply face
1-ply face
2-ply face
2-ply face
Figure 8. Measurements of three-specimen-average ultimate load for sandwich specimens, showing
the percent change due to the modified configuration.
a) PVC core, 1ply faces
b) PVC core, 2ply faces
c) Web core, 1-ply
faces
d) Web core, 2-ply
faces
Figure 9. Example ultimate failure modes with modified configuration.
DETAILS OF THE MODIFIED TEST FIXTURE DESIGN
The components of the test fixture and associated hardware are shown in Figures
3(c-d) and 10. The doubler plates feature blind threaded holes (Figure 10(a)) matching
the pattern in the fixture bars (Figure 10(b)), and are used with 8mm (5/16”) diameter
shoulder bolts for assembly. This feature reduces specimen preparation effort by
eliminating the machining of a large number of holes. It was found that with the two
independent frames on each side of the specimen, spurious stresses sometimes caused
failure of the core in flatwise tension. In response, the doubler plates were modified with
two through holes each (Figure 10(a)), allowing eight shoulder bolts to span the full
thickness of the assembly to lightly compress the specimen and prevent the tension
failure. After testing, doubler plates bonded with epoxy can be conveniently released
for reuse by heating in an oven to 175°C.
One of the tension loading fixtures is shown in Figure 3(d). The two studs at the
bottom engage with the corner holes of each frame to apply the load. The components
are steel and are intentionally massive for stiffness to minimize deflection, as each stud
carries its load in cantilever fashion. Locking collars are used to keep the two arms
Blind threaded holes
Two through holes
a)
c)
Steel doubler plate
b) Fixture bars
Bonding alignment fixture
d) Prepared specimen with doublers
Figure 10. Components of modified picture-frame test configuration.
against the assembled specimen and frame. At each of the non-loaded frame corners, a
stud in one bar engages the hole in the overlapping bar (Figure 10b)). To accurately
locate all of the doublers in proper relative position during bonding, an alignment fixture
is used (Figure 10(c)). A prepared specimen is shown in Figure 10(d). The test
configuration and hardware have been found to work well, although the specimen-plusframe assembly is heavy for manual handling, >30 kg (>70 lb). Mechanical assistance
for lifting and positioning is desirable.
The vertical arms of the loading fixtures (Figure 3(d)) were machined for sliding fits
with respect to both the 38mm diameter cylindrical bar at the top and the 25mm diameter
stud at the bottom. Although the diameter clearance is only about 0.1mm (0.004”), there
was an undesirable backlash which could cause minor misalignments in the fixture
under load. Shims were used during experiments and testing to reduce the backlash, but
this is not an ideal solution. A new design for the vertical arms will feature a press fit
for the studs at the bottom of the figure, and a split fixture with clamping bolts for the
top.
CONCLUSIONS
An experimental and analytical investigation of two picture-frame test
configurations for in-plane shear strength of sandwich constructions was conducted.
Both configurations used specimens with edge doublers to provide a 229mm square
interior gage area. The baseline configuration used corner pins spaced at 267mm,
extending between the frames on the two sides of the specimen in the conventional
manner, and requiring corner cutouts in the specimen. The modified configuration
featured frame corner pins spaced at 229mm (centered on the corners of the gage area)
which were discontinuous between the two sides of the specimen.
Instrumented tests were conducted on a sandwich configuration with foam core and
E-glass/epoxy face sheets featuring double-bias [-45/+45]2 fiber architecture. The shear
strain in the gage area was studied analytically with finite element analysis (FEA), and
experimentally with strain gages and digital image correlation (DIC). The following
conclusions were drawn:
1. The modified test configuration is superior to the baseline both in the uniformity
of the shear strain field, and in the agreement of measured strains with the ideal
value (the nominal value assuming uniform shear stress over the gage area).
2. Shear strain at the specimen center for the modified configuration agreed very
well with the ideal value in both FEA and experimental measurements, and good
uniformity of the strain over the gage area was demonstrated.
3. For the baseline configuration with composite edge doublers, the measured
shear strain at the center of the gage area was approximately 15% below ideal.
DIC measurements showed significant nonuniformity of the shear strain, with
higher strains at the edges of the gage area, and lower strains near the corners.
4. With the baseline configuration, FEA results indicate that increasing the
stiffness of the edge doublers has a negative effect on both the uniformity of the
shear strain field, and on the agreement with the ideal value.
5. The hardware implementation of the modified configuration worked well and
appears suitable for repeated use in standard testing. A minor modification to
the first-generation test fixture design was recommended.
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