each side of the test specimen is comprised of four rigid bars which feature a ..... composite webs surrounded by low-density polyisocyanurate foam (Milliken.
In-Plane Shear Characterization of Sandwich Laminates Using a Picture-Frame Test Configuration Frederick C. Stoll Nathan G. Johnston Paper Number 3211, Presented at American Society of Composites 31st Technical Conference, Williamsburg, VA, September 19-22, 2016
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. 1
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. 2
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
3
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. 4
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 elastic constants [ , , ν ] can be reduced for these laminates to three 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) 5
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 was evaluated between 0.1% and 0.4% strain, and ν was beyond 0.4%, so 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.
6
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.
7
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 1054;,A a layer of resin absorbed into the open cells at the surfaces of the core, 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
8
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. 9
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 10
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.
11
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 12
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 13
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. REFERENCES 1. 2. 3. 4. 5. 6. 7.
8.
9.
Ley, R.P., Lin, W., and Mbanefo, U., “Facesheet Wrinkling in Sandwich Structures,” NASA CR1999-208994, 1999. Stoll, F., “Blade Cost and Weight Reduction using TYCOR W Engineered Cores,” presentation to 2010 Sandia Wind Turbine Blade Workshop, Albuquerque, NM, July 20-22, 2010 Adams, D.F., “The Picture Frame Shear Test method,” High Performance Composites Magazine, November, 2014. Chaterjee, S., Adams, D., and Oplinger, D.W., “Test Methods for Composites - a Status Report Volume III: Shear Test Methods,” Report No. DOT/FAA/CT-93/17, III, 1993. Farley, G.L., Baker, D.J., "In-Plane Shear Test of Thin Panels," Experimental Mechanics, Vol. 23 No. 1, 1983, pp. 81-87. Diel, S., Huber, O., Steinmann, P., and Winter, W., “Design and validation of a new fixture for the shear testing of cellular solids,” Arch Appl Mech (2014), 84:309–321. Trappea, V., Basanb, R., and Grassec, F., “Stiffness and Fracture of Shear Loaded Laminates with Unidirectional and Biaxial Fibre Orientation Investigated with a Picture Frame Test,” Proceedings of ECCM16 – 16th European Conference on Composite Materials, Seville, Spain, 22-26 June 2014. DIN SPEC 4885, “Shear test method using a shear frame for the determination of the in-plane shear stress/shear strain response and shear modulus,” Deutsches Institut für Normung, Berlin, Germany, 2014. ISO 527-4, “Plastics, Determination of Tensile Properties, Part 4: Test Conditions for Isotropic and Orthotropic Fiber-Reinforced Plastic Composites,” International Organization for Standardization, Genève, Switzerland, 1997.
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