EFFECTS OF ANGLED STITCH REINFORCEMENT ON FOAM CORE SANDWICH STRUCTURES. Thien Sok

Transcription

1 EFFECTS OF ANGLED STITCH REINFORCEMENT ON FOAM CORE SANDWICH STRUCTURES by Thien Sok A thesis submitted to the faculty of The University of Utah in partial fulfillment of the requirements for the degree of Master of Science Department of Mechanical Engineering The University of Utah December 2010

2 Copyright Thien Sok 2010 All Rights Reserved

3 The University of Utah Graduate School STATEMENT OF THESIS APPROVAL The thesis of Thien Sok has been approved by the following supervisory committee members: Daniel O. Adams, Chair 07/14/2010 Date Approved K. Larry DeVries, Member 07/14/2010 Date Approved David W. Hoeppner, Member 07/14/2010 Date Approved and by Timothy A. Ameel, Chair of the Department of Mechanical Engineering and by Charles A. Wight, Dean of The Graduate School.

4 ABSTRACT The mechanical effects of varying stitch angles in stitched sandwich structures were investigated. A repeatable fabrication method was developed for stitching Kevlar yarn through sandwich panels of polyurethane foam and dry woven carbon fiber facesheets. Unstitched, 30 o, 45 o, 60 o, and 90 o stitch orientations, measured from the horizontal, were used throughout the study. Mechanical tests of flatwise tension, flatwise compression, and core shear were performed to observe the stiffness and strength effects of the different stitch configurations. With only a measured areal density increase of 2-7%, stitches oriented at 90 o were found to nearly double out-of-plane tensile stiffness and strength while compressive stiffness and strength increased only modestly. Stitches oriented at 30 o and 45 o were found to increase shear stiffness and strength by as much as 70% and 100%, respectively. Analytical models were developed to predict out-of-plane tensile modulus and shear modulus using a rule-of-mixtures approach to account for the separate constituent properties. As expected, the model predicts 90 o as the optimal stitch orientation for flatwise tension stiffness; however, 35.2 o was found to be optimal for shear stiffness. Agreement between model predictions and experimental measurements were within 15% for the core shear model and within 4% for the flatwise tension model when compared to the average. Creep behavior of the varying stitch angle reinforced sandwich structures was also investigated. Portable spring loaded creep test fixtures were

5 specially developed to test reinforced polymer core sandwich panel configurations under flatwise compression and core shear. Traditional creep test methods of using dead weight were not practical for testing multiple specimens simultaneously at high loads (above 3.5 kn in flatwise compression and 5.3 kn for core shear). The required 60 tests at 150 hours made use of servo-hydraulic and electromechanical machines for creep testing not feasible. Low cost creep test fixtures were developed to accurately apply and sustain load, allow for creep measurement, and be compact enough to fit six test fixtures simultaneously in a large convection oven approximately. Use of the developed creep test fixtures resulted in typical creep curves for polymers loaded under compression and shear for the collected creep data with only a 3.8% and 1.2% load loss for flatwise compression creep and core shear creep testing, respectively. Measurements of creep under constant flatwise compressive stress indicate stitch angles of 90 o decreased creep by approximately 75% when compared with unstitched sandwich specimens. Measurements of creep under core shear loading indicate stitch angles 30 o, 45 o, and 60 o decreased creep by approximately 70% when compared to unstitched sandwich specimens. iv

6 TABLE OF CONTENTS ABSTRACT...iii LIST OF TABLES... vii INTRODUCTION... 1 CHAPTERS 1. EFFECTS OF ANGLED STITCH REINFORCEMENT ON STIFFNESS AND STRENGTH PROPERTIES OF SANDWICH STRUCTURES Abstract Introduction Sandwich Panel Fabrication Flatwise Tension Testing Flatwise Compression Testing Core Shear Testing Analytical Modeling Discussion and Conclusions References CREEP OF POLYMER FOAM CORE SANDWICH STRUCTURES WITH ANGLED STITCH REINFORCEMENT Abstract Introduction Sandwich Panel Fabrication Flatwise Compression Creep Testing Core Shear Creep Testing Discussion Conclusions and Recommendations References PORTABLE SPRING LOADED CREEP TEST FIXTURES FOR SANDWICH COMPOSITES Abstract Introduction... 69

7 3.3 Test fixture development Flatwise compression creep testing Through-thickness shear creep test method Creep Testing of Composite Sandwich Specimens Summary References vi

8 LIST OF TABLES Table 1.1 Material selection Comparison of areal densities for varying stitch angles Comparison of flatwise tension tests results for varying stitch angles Comparison of flatwise compression tests results for varying stitch angles Comparison of core shear test results for varying stitch angles Material and geometry properties used in modulus models Material selection Comparison of areal densities for varying stitch angles Results for flatwise compression creep tests Results for core shear creep tests Test fixture load summaries... 75

9 INTRODUCTION Composite sandwich structures are a special class of composite materials which, as the name implies, consists of two thin, stiff facesheets that sandwich a thick, light core. The core is similar to the strut of an I-beam and serves to separate the facesheets which increases the structure s moment of inertia. Facesheets typically consist of metal or fiberbased composite, and the cores are often made of polymeric or metal foam, balsa wood, or honeycomb structures. Composite sandwich panels have been used widely for many years for their high bending stiffness and relatively low weight. To date, composite sandwich panels have found applications in aerospace, automotive components, and building structures. Polymer foams are a desirable core material for their low cost when compared to other traditional cores and their ease of manufacturing as the polymer foam can be molded and fabricated into complex shapes. However, the polymer foams, being the weaker of the two materials in a sandwich structure, are often a source of failure as they are also subjected to out-of-plane tensile and compressive stresses as well as shear stresses. Consequently, through-the-thickness core reinforcements have appeared in many cases in industry and in literature including through-the-thickness stitching. Stitching was found to greatly increase out-of-plane strength and stiffness as well as reduce debonding as a result of impacts [1]. Little work, however, has been performed on the

10 2 topic of varying the angle of stitch reinforcements. It was hypothesized that stitch angles more closely aligned with 45 o measured from the horizontal would increase shear performance of the sandwich structure as it is known that under a state of pure shear stress, the greatest tensile stress is on a plane oriented at 45 o to the original element. Since one primary use of sandwich structures are to carry bending loads, which result in shear forces through the core, varying the stitch angle to reinforce the core could result in more optimal designs. Therefore, knowledge of how varying the stitch angle affects the core behavior of the sandwich panel is important and could aid in design and optimization. The scope of this study is three fold: First, experimentally investigate how different stitch configurations affect stiffness and strength properties of the sandwich structure through mechanical testing. Second, develop a model to predict the out-ofplane tensile stiffness as well as the shear stiffness of the varying stitch angle reinforcement sandwich panels. Third, investigate the creep performance of these same angle-stitch reinforced sandwich panels when placed under constant load at elevated temperatures. The first study presented is an investigation into the stiffness and strength of varying angle stitch reinforcement. To investigate the effect of varying stitch angles through the sandwich structure core, a consistent fabrication technique was developed to stitch all of the sandwich panels. Five configurations of sandwich coupons were fabricated and tested: unstitched, 30 o, 45 o, 60 o, and 90 o when measured from the horizontal. All panels were stitched with Kevlar yarns through six-ply dry woven carbon fiber facesheets surrounding a polyurethane foam core. The different configurations were

11 3 compared through testing under flatwise tension, flatwise compression, and core shear. An analytical model was also developed in this first study to predict the stiffness of the different stitch configurations under flatwise tension and core shear. Since the varying stitch angle reinforcement was hypothesized to increase mechanical properties of the sandwich structures, it was natural to also hypothesize the stitch reinforcement could also reduce creep deformation. Creep has been shown to occur even at room temperatures for certain polymers, and therefore, a knowledge of time-dependent behavior of polymer foam sandwich composites is extremely important. Creep has been studied extensively in the areas of polymers and polymer foams; due to the complexity of the sandwich cores as a result of the reinforcements, however, insufficient information is known regarding their time-dependent behavior. The second study is a creep investigation of the different reinforced polymer foam core sandwich structure configurations. The different stitch reinforcement configurations were creep tested under flatwise compression and core shear. Creep test fixtures were specially developed to creep test the reinforced sandwich panels, as loads were too high and not practical for traditional dead-weight testing. Large numbers of tests at extended time periods were also required making servo-hydraulic and electromechanical machine use not feasible. The third section of this thesis is a detailed description of the spring-based creep test fixtures used in the creep study. The creep test fixtures developed also needed to comply with certain requirements including cost, size, and load monitoring.

12 CHAPTER 1 EFFECTS OF ANGLED STITCH REINFORCEMENT ON STIFFNESS AND STRENGTH PROPERTIES OF SANDWICH STRUCTURES 1.1 Abstract The effects of varying stitch angles on the stiffness and strength properties of stitched sandwich composites were investigated. A fabrication method was developed for stitching Kevlar yarn through sandwich panels of polyurethane foam core and dry woven carbon fiber facesheets. Mechanical tests were performed to observe the effects of stitching. Tests included flatwise tension, flatwise compression, and core shear. Analytical models were developed to predict out-of-plane tensile modulus and shear modulus. Agreement between model predictions and experimental measurements were within 15% for the core shear model and within 4% for the flatwise tension model. Stitches oriented at 90 o were found to greatly increase out-of-plane tensile and compressive modulus as well as tensile strength. Stitches oriented at 45 o and 30 o were found to produce the greatest increase in shear modulus and strength.

13 5 1.2 Introduction Composite sandwich panels have been widely used for many years. Sandwich panels have been used for their high stiffness properties without large increases in weight. To date, composite sandwich panels have found applications in aerospace, automotive components, and building structures. A sandwich composite is a special class of composite material that is fabricated by attaching two thin, but usually stiff facesheets to a lightweight and relatively thick core. The core material is normally a lower strength material, but its higher thickness provides the sandwich composite with high bending stiffness with overall low density. Closed cell structured foam, balsa wood, and aluminum honeycomb are commonly used core materials. The basic concepts of a sandwich panel are similar to an I-beam. The sandwich panel facesheets, like the flanges on an I-beam, accomodate the tensile and compressive stresses produced from bending loads. The core of sandwich panels, like the web of an I-beam, carry shear stresses under bending loads. Some common failures associated with sandwich structures are debonding between the core and facesheets, core failure in tension or compression, and core shear failure. Consequently, several types of through-the-thickness reinforcements have been investigated including stitching [1-5], strut webs [7], fiber infused pre-form cores [8], grids within the sandwich core [9], and z-pins [10], which have resulted in increases in specific strength, flexural stiffness, flexural strength, out-of-plane tensile strength, core shear strength, and, in many cases reduced debonding. Although there have been several studies of core reinforcements, to date, there has been little work done in investigating and modeling the effect of reinforcements oriented

14 6 at different through-the-thickness angles, specifically stitching at different angles. The focus of this study is a comparison of five stitched sandwich panel configurations: unstitched, 30 o, 45 o, 60 o, and 90 o stitched panels measured from the horizontal, tested in flatwise tension, flatwise compression, and core shear. Because vertical stitches have been used in the past to increase out-of-plane strength and modulus of sandwich structures [1-3], it was hypothesized that stitches at 45 o would more effectively increase shear performance as it is understood that for an element under pure shear, the principal tensile stress should be along its 45 o direction. This paper presents an investigation in using through-the-thickness stitching at various angles in composite sandwich structures. A stitch technique is outlined for fabricating the five stitch panel configurations. Test results are presented and a closed-form analytical model is used to predict the elastic outof-plane tensile modulus and shear modulus. 1.3 Sandwich Panel Fabrication Material selection The first step in comparing the performance of varying stitch angles in stitched sandwich composites was to develop a consistent fabrication technique to manufacture each configuration of stitched sandwich composite tested. Material selection was the first step in the fabrication process. Selection of materials was guided by past work in the areas of sandwich panels and stitching of sandwich panels at the University of Utah. The core, facesheet, and stitching used are presented in Table 1.1. The core used throughout the entire study was selected to be a closed-cell polyurethane foam of 160 kg/m 3 made by General Plastics LAST-A-FOAM FR-6710

15 7 Configuration Unstitched Stitching Method N/A Hand Stitched 30 o, 45 o, 60 o 1600 denier Kevlar 29 yarn 90 o Manual Industrial Sewing Machine 1600 denier Kevlar 29 yarn Table 1.1 Material selection Facesheet Material Woven carbon fabric T300B 3K ([0/90] 6T ) [12] EPON 862 epoxy EPON 9553 hardener [13] 6 plies/facesheet Core Material Polyurethane foam (FR-6710) (General Plastics) [11] 160 kg/m mm thick Curing Method Vacuum Assisted Resin Transfer Mold (VARTM) [11]. This foam was selected for its low density and the foam s ability to be stitched. Since a Vacuum Assisted Resin Transfer Mold (VARTM) technique would be used to infiltrate the panels, open-cell foams could not be used because they would saturate the foam with resin during the infiltration process. A nominal core thickness of 12.7 mm was used for all samples. Facesheets used in all specimens for this study consisted of six ply woven carbon fiber. The layup was [0/90] 6T. The stitching methods varied for unstitched specimens, 30 o, 45 o, 60 o stitch specimens, and 90 o stitch specimens, and are described in the next three sections Unstitched panels and VARTM process Using previously described materials, six plies of woven carbon fiber sheets were used to sandwich a piece of 12.7 mm thick polyurethane closed cell foam. The panel

16 8 assembly was then placed on a flat aluminum tool, vacuum bag sealed, and infiltrated using VARTM processing. VARTM processing utilizes a vacuum and atmospheric pressure to force resin through a bag and mold. Prior to infiltration, porous Teflon-coated fiberglass was used to wrap the panel allowing the release of the panel upon curing. A layer of coarse nylon mesh was placed on the top and bottom sides of the panel to assist transfer of the resin and distribute it to the top and bottom surfaces of the panel. Two tubes were inserted into the vacuum bag, where one tube was attached to a vacuum pump and the other was sealed and placed in a resin container and then opened to allow the transfer of resin. Within the vacuum bag, a breather cloth was placed towards the suction side of the bag to absorb excess resin and allow for any gases to exit the bag. An assembly of the VARTM process is illustrated in Fig o stitched panel fabrication 90 o sandwich panels, like the unstitched panels, also consisted of 6 plies of woven carbon fiber as faces sheets sandwiching a 12.7 mm thick polyurethane closed cell foam with the same [0/90] 6T layup. The facesheets were temporarily taped in place while a 12.7 mm x 12.7 mm grid pattern was also temporarily taped onto the facesheet surface. This grid pattern was used as a template to stitch using a Consew industrial sewing machine, as seen in Fig The opposing side also had paper taped to the surface to prevent the sewing machine feeder from moving the woven carbon pattern. Following the stitching, the paper template was removed. The panels were stitched using 1600 denier Kevlar 29 yarn.

17 9 Figure 1.1 Cross section of an infiltration assembly Figure 1.2 Stitching of sandwich composites using an industrial sewing machine All stitch panels in this study used a modified lockstitch. Traditional lockstitch techniques have equal tension on both the needle and bobbin threads and ideally have the locked portion somewhere near the middle of the thickness. A modified lock stitch was used, as shown in Fig. 1.3, where the upper thread is pulled completely through to the bobbin side. This allowed for easier fabrication as the yarn intersection was made outside the core and the yarns kept continuous through the core.

18 10 Figure 1.3 Diagram of traditional and modified lock stitch o, 45 o,60 o panel fabrication The fabrication process for the angled stitch panels differed as the panels were hand stitched. The layup was exactly the same for the unstitched and 90 o stitched panels. However, the foam had predrilled holes at 30 o, 45 o, and 60 o. The holes were made using a water jet cutting machine with 12.7 mm between stitch rows and 12.7 mm between stitches. The direction of the angled holes alternated from row to row. These panels were also stitched using the same 1600 denier Kevlar 29 yarn. A diagram of an angle stitched panel can be seen in Fig A method was developed to hand-stitch the angled stitch panels with the same modified lockstitch as the 90 o panels. Prior to stitching, needles were modified as shown in Fig The needle eye hole was cut at one side to allow the Kevlar yarn to pass through similar to using the needle as a hook. With approximately 20 needles modified, the panels could then be prepared for stitching. The facesheets were set up as if they were a cover of a book, as shown in Fig. 1.6, exposing the core and revealing the predrilled stitch holes. With the facesheet folded open to expose the core, the precut holes were accessible so needles could be placed and pushed through, as illustrated in Fig. 1.7 (Step 1 and 2). The needles shown in Fig. 1.7 (Step 3) were then pushed through to the

19 11 Figure 1.4 Schematic of a stitched sandwich panel Figure 1.5 Modified needle for stitching 30 o, 45 o, 60 o sandwich panels

20 12 Figure 1.6 Stitch process with core and book-cover-like woven carbon fiber facesheets with needles placed in the core Figure o, 45 o, 60 o, stitching process other side until flush with the core surface using a thimble. Next, the top facesheet shown in Fig. 1.6 was folded closed where the needles were then pushed back up through to be sewn at each stitch, as shown in Fig. 1.7 (Step 4). A picture of the stitch process in Fig. 1.7 (step 4) can be seen in Fig. 1.8 (left). The needle was then threaded and pulled through the core which also pulled the Kevlar yarn through the core. Then, the needle was removed, and the bobbin thread was passed under the new stitch, as shown in Fig. 1.7 (Step 5). The stitch was finally pulled tight, as shown in Fig. 1.7 (Step 6). The

21 13 process is then repeated and can be seen in Fig. 1.7, step A finished panel can be seen in Fig. 1.8 (right). As the needle pushes through the foam core, cells of the closed cell foam are pierced or compressed to rupture. These damaged cells are filled with resin during the infiltration process resulting in a resin column, as shown in Fig In general, the resin column size is proportional to the needle diameter. A 2.9 mm diameter needle is common for industrial sewing machines. However, a needle diameter of 1.9 mm was selected as being large enough to perform the stitching without breaking while reducing the resin column size [1]. The predrilled holes for the 30 o, 45 o, or 60 o stitched sandwich panels, on the other hand, measured 1.3 mm. When the modified needles with Kevlar would pull through, the hole would widen and a comparable resin column resulted for both the 90 o stitch panels and the 30 o, 45 o, and 60 o stitch panels. It is noted that the resin column diameters varied significantly within a stitch and from stitch to stitch. Following VARTM infiltration for all stitched and unstitched panels, areal density measurements were taken for each stitch configuration, shown in Table 1.2. Slight Figure 1.8 Sewing process with Kevlar yarn and needle (right), typical stitched sandwich panel surface (left)

22 14 Figure o (left), 60 o (right), stitched sandwich panel with stitch and resin column Table 1.2 Comparison of areal densities for varying stitch angles Configuration Areal Density g/m^2 Normalized Areal Density Unstitched 6740 ± Degree 7060 ± Degree 6920 ± Degree 6970 ± Degree 7300 ± increases in density were observed for stitched panels. Variations in density are believed to be slight inconsistencies in infiltration where resin did not completely penetrate through the stitches. A 60 o panel with the foam removed can be seen in Fig Flatwise Tension Testing Flatwise tension test method Flatwise tensile testing of the subject reinforced sandwich structures was conducted in accordance with ASTM C297/C297M-04 [14] to investigate improvements in interlaminar tensile strength due to stitching. Fig shows a typical test setup.

23 15 Figure 1.10 Typical 60 o stitch panel with foam removed Figure 1.11 Typical flatwise tension test setup with extensometers attached Each coupon tested was 51 mm x 51 mm and was adhesively bonded to steel loading blocks of the same size using Hysol 907 epoxy. All specimens were cut using a water cooled diamond saw. Two specially designed extensometers were used to measure the strain produced on opposite sides of the sandwich specimen during loading. The extensometers were removed prior to failure. A total of five configurations were tested, four with stitches and one unstitched. Each coupon containing stitches contained four rows of centered stitches, each with four stitches in each row. It is important to note that

24 16 specimens containing 30 0, 45 0, and 60 o stitches had only three complete (unservered) stitches per row as the stitch extended beyond the test size. The specimens were tested using a 44 kn electro-mechanical machine. After testing, displacements from the attached extensometers were averaged. Out-of-plane modulus was calculated using: E = (1.1) where P is load, l is the length of the specimen, w is the width of the specimen, and ε is the out-of-plane strain in the z direction calculated from dividing the averaged displacements of the extensometers by the initial thickness. Maximum stress was calculated using: σ, = (1.2) where P, is the maximum load. Six specimens from each configuration were tested Flatwise tension failure modes First, failure is defined as a large instantaneous load drop greater than 10% of current load or as a discontinuous slope found in a stress versus strain or crosshead displacement plot. Failure of unstitched and 90 o coupons resulted in one catastrophic failure, as shown in Fig Failure of the unstitched specimens resulted in a complete separation into two halves. 90 o stitch panels for most tests exhibited the same failure

25 17 Figure 1.12 Failure in a flatwise tension test of an unstitched (left), 90 o (right) sandwich panel. mode; stitches and foam would fail simultaneously. For the 30 o, 45 o, and 60 o stitch angles, the foam would separate first with the stitches still intact, resulting in a major load drop. The two halves of the specimen did not completely separate due to stitching. Fig represents a typical 30 o panel failure compared with a crosshead versus load plot, where the first initial load drop results in fracture of the core. As the 30 o stitch panel is loaded further, residual strength is dependent on the stitching, as shown by an increase in load. Failure progression is noticed as load-carrying stitches fail. At this point, the panel has not completely separated; however, the two halves of sandwich panel can no longer carry higher loads Flatwise tension test results Table 1.3 shows the results obtained from flatwise tensile testing. As expected, the out-of-plane tensile strength and modulus increase as the stitch orientation becomes more aligned with the applied load. It is clear, from the testing, that the angle of stitch is a significant structural parameter. With only slight increases in areal weight (see Table

26 18 Figure 1.13 Failure progression in flatwise tension test of a 30 o stitch panel Table 1.3 Comparison of flatwise tension tests results for varying stitch angles Stitch Angle Flatwise Tensile Strength (MPa) % Improvement from Unstitched Out-of-Plane Modulus of Elasticity (MPa) (Tensile) % Improvement from Unstitched Unstitched 1.04 ± ± o 1.27 ± ± o 1.58 ± ± o 1.88 ± ± o 2.23 ± ±18 91

27 19 1.2), the stitch contribution is great. The modulus for the 90 o stitch panel is nearly double and the maximum stress attained is over twice that of the unstiched panel. The 30 o, 45 o, and 60 o stitch panels do not fail completely with the initial load drop, as seen in Fig Since the stitches do not fail on the initial load drop, the panel, when loaded further, was able to carry approximately half their respective maximum loads. 1.5 Flatwise Compression Testing Flatwise compression test method Flatwise compression tests were carried out in accordance with ASTM C [15]. Because of the similarities in testing between flatwise tension and compression, similar performances were hypothesized. A flatwise compression test is shown in Fig No extensometers were used to measure deflection of the tested coupon. With minimal fixture compliance, the strains for each test were calculated from the deflection of the crosshead using the same equations as the flatwise tension tests, Eq. 1.1 and 1.2. Six specimens from each configuration were tested Flatwise compression failure modes Similar failure modes were observed for all sandwich panel configurations. From the load versus crosshead displacement plots, seen in Fig. 1.16, all configurations had a peak load followed by a small drop in load followed by a plateau. Load was removed after a sufficient plateau was noticed. It was concluded that the foam cores failed due to crushing of the foam cells as little recovery was observed.

28 20 Figure 1.14 Stress versus crosshead displacement plots for flatwise tension tests for varying stitch angles Figure 1.15 Typical flatwise compression test setup

29 21 Figure 1.16 Stress versus crosshead displacement plots of flatwise compression test for varying stitch angles Flatwise compression test results As predicted, in flatwise compression, the sandwiches with the 90 o stitching performed best with approximately a 60% improvement in maximum compressive strength as well as a modest gain of 23% in modulus which can be seen on Table 1.4. Surprisingly however, the there was very little gain for the 30 o and the 45 o stitched panels, and even a decrease in modulus for those configurations. Having the 30 o stitches actually decreased the modulus by almost 17%. 1.6 Core Shear Testing Core shear test method Interlaminar shear tests were conducted in accordance to ASTM C [16]. A series of coupons, 51 mm x 203 mm, were prepared from the same five stitch

30 22 Table 1.4 Comparison of flatwise compression tests results for varying stitch angles Angle Flatwise Compressive Strength (Mpa) % Improvement from Unstitched Out-of-Plane Modulus of Elasticity (Mpa) (Compressive) % Improvement from Unstitched Unstitched 2.60 ± ± o 2.68 ± ± o 2.81 ± ± o 3.02 ± ± o 4.16 ± ± configurations and tested. Two steel plates were adhesively bonded to the face of each coupon using Hysol 907 Epoxy and can be seen in Fig (left) where the line of load passes through opposing diagonal corners. A measure of shear stiffness was made using the same extensometers used for flatwise tensile testing, as seen in Fig (right). These extensometers were attached using aluminum angle pieces bonded to the steel loading plates. Using this configuration, the extensometers measured the relative axial motion occurring between the opposing faces of the sandwich composite specimen. Shear stress was calculated as follows: τ= (1.3) where P is load, L is the length of the specimen, and b is the width of the specimen. Shear modulus was calculated using:

31 23 Figure 1.17 Core shear test schematic with sandwich panel showing load from corner to corner (left), typical core shear test setup with extensometers attached (right) G= (1.4) where S is the slope of the initial portion of the load-deflection curve ( P/ u) where u is the displacement of the loading plates measured using the extensometers and t is the initial thickness of the specimen Core shear failure modes Two main types of failure occurred for the differing stitch configurations. For the unstitched and the 90 o, failure resulted in a shearing and complete separation of the two halves, as shown in Fig Failure of the 30 o, 45 o, and 60 o panels resulted in foam failure with a large drop in load while the two halves remain together, as seen in Fig In the 30 o, 45 o, and 60 o

32 24 Figure 1.18 Failure in a core shear test of an unstitched (left) and a 90 o (right) sandwich panel Figure 1.19 Failure in a core shear test of a 30 o stitch panel stitch panels, the angle direction alternated from row to row resulting in every other row to be loaded in tension, and their alternating rows to be loaded in compression. This can be seen in Fig (left) where the Kevlar yarn is severed near the bottom facesheet and in Fig (right), buckling of the stitches can be seen near the centers of each stitch Core shear test results Note that because each specimen with angled stitching had alternating directions of stitches (+θ to a -θ orientation in adjacent stitch rows), only half the stitches were effectively loaded in tension during shear loading. It is understood for an isotropic element under pure shear, the principle tensile stress should be along its 45 o direction.

33 25 Figure 1.20 Tensile failure (left), buckling failure (right) of stitches under core shear testing (foam removed) Therefore, there was considerable interest to see how the Kevlar stitching would affect the shear behavior of the sandwich structure. Table 1.5 shows the results obtained from interlaminar shear testing. Significant improvements in both shear strength and shear modulus was observed using all four stitch angles investigated. The highest values of shear strength and modulus were recorded for the 30 and 45 stitch orientations, respectively. Plots of the shear stress versus crosshead displacement are found in Fig For the unstitched panel, complete failure was noticed. Stitched panels, on the other hand, have a progression of failures similar to the flatwise tension tests previously described. Core shear testing of the stitched sandwich specimens resulted in failure with significant load drop; however, the stitching would hold the sandwich specimen together. As the specimen is loaded further, some increase in load is noticed with significant crosshead displacement.

34 26 Table 1.5 Comparison of core shear test results for varying stitch angles Stitch Angle Core Shear Strength (MPa) % Improvement from Unstitched Core Shear Modulus (MPa) % Improvement from Unstitched Unstitched 0.84 ± ± o 1.68 ± ± o 1.62 ± ± o 1.54 ± ± o 1.51 ± ± Figure 1.21 Shear stress versus crosshead displacement plots of core shear tests for varying stitch angles

35 Analytical Modeling A model, adapted from an analysis of strut webs [7], was developed to predict the elastic behavior of the varying stitched sandwich configurations. Using a mechanics-ofmaterials-based approach, an analytical solution was developed to predict out-of-plane modulus as well as shear modulus. Using a rule-of-mixtures (springs in parallel) approach, the stitch contribution can be accounted for by its area fraction. The rule-ofmixtures approach assumes relative strain as equal for the foam and stitch. An easy to use analytical solution was developed to predict out-of-plane modulus and shear modulus of the sandwich structure as functions of the foam modulus, stitch modulus, stitch angle, stitch area, stitch spacing, and the stitch row spacing Out-of-plane tensile modulus An analytical model has been developed to predict the out-of-plane tensile modulus. A diagram of a model element is shown in Fig To model the out-ofplane tensile modulus, a mechanics-of-materials-based approach to model the stitch was taken. A rule-of-mixtures was used to account for the foam and stitch contribution to the modulus by way of area fraction. Starting from a global coordinate system (refer to Fig. 1.22) and using a rule-ofmixtures, the stitch and foam contribution to the modulus is given by the relationship: E =E α+e (1 α) (1.5)

36 28 Figure 1.22 Out-of-plane tensile modulus model with stitch where E, E, and α are the out-of-plane stitch contribution to the modulus, the out-of-plane foam modulus, and area fraction of the stitch, respectively. First the stitch is assumed to be a two-force member with pinned-pinned boundary conditions at the top and bottom of the stitch. E can be related to stress and strain of the stitch given as: E = (1.6) where, σ sinθ, is the vertical stress component of the stitch, where σ and θ, and ε are the axial stress in the stitch, the stitch angle with respect to the horizontal, and the out-ofplane strain, respectively. The lateral forces are assumed to be balanced due to the alternating stitch pattern. The axial stress in the stitch is then defined as:

37 29 σ =ε E (1.7) where ε and E are the axial stitch strain and axial stitch modulus, respectively. The axial strain in the stitch can then be related to the global out-of-plane strain by way of strain transformation equations: ε =±γ cosθ sinθ +ε sin θ (1.8) where γ is the shear strain in the subject element. It is assumed that for a uni-axial load, or stress only in the z-direction, the shear strain is zero or γ = 0. After substitution of variables for axial strain and simplifying, the stitch contribution to the global out-of-plane modulus results in the relation: E =sin θ E (1.9) Finally, area fractions are defined as: α= (1.10) where A, a, and b are the stitch cross-sectional area, stitch spacing, and stitch row spacing, respectively. After substituting into the original out-of-plane modulus equation, Eq. 1.5, the final equation results in the relation:

38 30 E =sin θ E + E 1 (1.11) Shear modulus An analytical shear model was also developed using a mechanics-of-materialsbased approach. A rule-of-mixtures was again used to account for the addition of stitches shown in Fig Using area fractions to account for the stitch contribution to the shear modulus, the model starts with the relation: G =G α+g (1 α) (1.12) where G, G, and α are the effective shear modulus contribution from the stitch, the shear modulus of the foam, and the stitch area fraction, respectively. The stitch is still assumed to be a two-force member with pinned-pinned boundary conditions at the top and bottom of the stitch. G can be related to stress and strain of the stitch given in Eq. 1.13: G = (1.13) where, σ cosθ, is now the horizontal stress component of the stitch, where σ and θ, and γ are the axial stress in the stitch, the stitch angle with respect to the horizontal, and the global shear strain, respectively. Axial stitch stress is defined as:

39 31 Figure 1.23 Shear modulus model with stitch σ = ε E (1.14) where ε and E are the axial stitch strain and axial stitch modulus, respectively. The axial strain of the stitch can be related to the global shear strains by way of strain transformation equations: ε = γ cosθ sinθ + ε sin θ (1.15) where γ and ε are the global shear strain for the stitch and the global out-of-plane strain for the stitch, respectively. For a pure shear loading condition, the out-of-plane strain is assumed to be zero or ε = 0. After variable substitution of variables and simplifying, the resulting global shear modulus for the stitch results in the relationship:

40 32 G = cos θ sinθ E (1.16) Thus, after variable substitution, replacing the area fractions, α, with that found in Eq. 1.12, and simplifying, the global shear modulus for the stitched sandwich structure results in the model written as: G =cos θ sinθ E + G 1 (1.17) Material properties for model To completely model the out-of-plane modulus and the shear modulus of the different stitch sandwich configurations, three material properties and four stitch geometries were required. The material properties required included: E, G, and E, which are the out-of-plane modulus of the foam, shear modulus of the foam, and stitch modulus, respectively. The stitch geometries needed included: θ, a,b, and A which are the stitch angle, stitch spacing, row spacing, and the stitch cross sectional area, respectively. Material and geometry parameters used for modeling are shown in Table 1.6. E and G, were obtained from experimental data collected from the previously described flatwise tension and core shear tests of unstitched panels. E, the modulus of the stitch, was obtained from previous work performed at the University of Utah [1], where a rule of mixtures was also used to determine the modulus of the stitch. E was calculated as:

41 33 Table 1.6 Material and geometry properties used in modulus models E G θ a s b s A s Foam 101 Mpa 47.8 Mpa Stitch 10.1 GPa varying 12.7 mm 12.7 mm 1.47 mm 2 E =E V +E V (1.18) where the diameter of the Kevlar and resin were measured using vernier calipers to calculate the volume fractions and the Kevlar/resin moduli were determined from references [17, 18]. E was found to be 10.1 GPa. Stitch geometries, stitch spacing, row spacing, and stitch angle were all previously defined. However, in using the stitch as a lumped parameter instead of Kelvar and resin separately, an in-situ approach was taken to obtain an effect area of the stitch using the previously described model. The effective stitch area was found using the previously described flatwise tensile modulus model: Eq. 1.11, 90 o stitch specimen test data, and unstitched specimen data. The effective out-of-plane modulus for the 90 o stitch panel was assumed as 194 MPa taken from experimental data. E was found to be 10.1 MPa from previous work performed at the University of Utah. E was measured as 101MPa from unstitched specimen test data. The geometries θ, a, and b are defined as 90 o, 12.7 mm, and 12.7 mm, respectively. Finally, solving for A, 1.47 mm 2 is the resulting effective stitch area. With this area, the effective diameter was found to be approximately 1.37 mm. In short, the model was curve fit to the 90 o experimental data point.

42 34 The needle diameter used for the 90 o stitched panels was 1.9 mm (refer to section 2.3) which is larger than the in-situ diameter found. This may be due to a varying diameter along the stitch as well as variation from the reported manufacturer material properties used in obtaining the stitch modulus. Measurements taken of the resin columns varied over the length of the stitch as well as from stitch to stitch. This in-situ approach is, in effect, an average stitch area. This effective in-situ area of the stitch found, using the out-of-plane modulus model, was also used in the shear modulus model. With all model parameters defined, predictions can be made at varying stitch angles Model results Agreement between model predictions and experimental measurements were within 15% for the core shear model and within 4% for the flatwise tensile model when compared to the average. The predicted and measured moduli values for the flatwise tension tests are presented in Fig The out-of-plane modulus model presented above used material properties from the unstitched and the 90 o, so an exact match between the experimental and predicted values at those two points is expected. The predicted values at the remaining angles matched extremely well when compared to the average, within 4%. The model was extremely accurate for the out-of-plane tensile predictions even though the 30, 45, and 60 panels had only three complete stitches per row (refer to section 3.1), where the fourth stitch was severed due to the size of the coupon whereas the 90 0 stitched coupon had four complete stitches.

43 35 Figure 1.24 Comparison between predicted and measured out-of-plane modulus (tension) The predicted and measured values for shear modulus are presented in Fig The shear model used core shear experimental test data from the unstitched panels. All other properties were kept the same. It is known that an isotropic material loaded under pure shear will have maximum tensile and compressive forces along its 45 o, but 35.2 o was found to be the mathematical maximum prediction. The results of the core shear test do not indicate a clear maximum as the 30 o and 45 o stitch panels performed statistically similar. The model also does not take into account the alternating stitch direction. It is known from the flatwise compression tests that the behavior of stitching in some cases, at low angles of 30 o, and 45 o (refer to section 4.3), may decrease the effective modulus of the sandwich core under compression. Due to the alternating stitch pattern of the 30 o, 45 o and 60 o, under load, the core shear tests effectively apply tensile loads to only half the

44 36 Figure 1.25 Comparison between predicted and measured shear modulus stitches, while the other half are loaded under compression previously described in the core shear test section and seen in Fig The stitches loaded under compression do contribute to the overall shear modulus; however, the lower contribution from these stitches may be a factor in the differences seen in Fig for the 30 o and 60 o panels between the predicted and measured shear modulus values. Due to the accuracy at 45 o, it is hypothesized that these 45 o stitches loaded in compression contribute relatively more, as seen in Table 1.4. The 90 o stitch specimens were shown to increase the modulus by approximately 23% when compared to the unstitched specimens, while the 60 o specimens only had an increase in modulus of 2.7% and the 30 o and the 45 o stitch specimens resulted in a decrease in modulus. In other words, the 45 o stitch reinforcement under compression found in a core shear test

45 37 contribute relatively more when compared to the other stitch angles under compression due to flatwise compression results. 1.8 Discussion and Conclusions A repeatable stitch method was developed and presented for the fabrication of 30, 45, 60, and 90 o stitch panels so that mechanical performances of these configurations could be compared. Up to a 7% increase in areal density was recorded. Out-of-plane tensile strength, tensile modulus, compressive strength, compressive modulus, shear strength, and shear modulus have been investigated for the five configurations of stitch reinforced polyurethane sandwich structures. As expected, the out-of-plane tensile strength and modulus increases as the stitch orientation becomes more aligned with the applied load with the highest load and modulus reported in the 90 o stitch panel. Whereas stitch failure occurred initially in the 90 stitched specimens, an initial foam failure occurred in the three angled stitch configurations: 30, 45, and 60 keeping the test coupon together during testing. Also, as predicted, the out-of-plane compressive strength favored the 90 o stitching angle as well as a modest gain of 23% in modulus. Surprisingly however, there was very little gain for the 30 o and the 45 o stitched panels in terms of compressive strength, and even a decrease in compressive modulus. Significant improvements in both shear strength and shear modulus was observed using all four stitch angles investigated. The highest values of shear strength and modulus were recorded for the 30 and 45 stitch orientations, respectively. Increases in areal density and cost of stitching are tradeoffs to these increases in stiffness and strength. Stitch

46 38 reinforcements may be optimized for use limiting their increases in weight and cost while significantly improving performance. Analytical models were also developed for out-of-plane tensile modulus as well as interlaminar shear modulus. Material properties used in the models were obtained from experimental data. An effective in-situ area was also obtained from the data. Agreement between model predictions and experimental measurements were within 15% for the core shear model and within 4% for the flatwise tension model when compared to the average. The discrepancies between the shear model and measured values were thought to have come from the alternating stitch pattern which resulted in compressive stress for half the angled stitches. Compressive performance is quite low for stitches not aligned to the loading direction. Alternating stitch patterns during core shear testing loaded half their respective stitches in tension and half in compression. The model assumed properties for all stitches were in tension. The model predicted the 45 o stitch specimens more accurately leading to a hypothesis that the 45 o stitches loaded under compression contribute relatively more than any other angle. The area fraction parameters assume the contribution of the stitches to be linear. One reservation to these models is that the limits of this assumption may need to be validated at extreme area fractions. Extremely high or low area fraction percentages may yield unexpected behavior. Overall, the closed form models presented predicted the out-of-plane and shear modulus within 15% and include parameters for stitch angle, stitch spacing, and row spacing which could be easily implemented in design. Using stitching to reinforce sandwich core performance is feasible. Adding stitch reinforcement is a tradeoff due to increases in areal density and cost. The addition of

47 39 strong, lightweight fibers, such as Kevlar, only slightly increases density by as little as 2-7%, while performance gains are potentially large with more than doubling the stiffness and strength in some cases. The ability to use stitches at different angles enables designs to be optimized for improved load handling. Optimization has always been important to the class of composite materials and sandwich structures. The approaches presented above may provide additional tools for the designer to reinforce sandwich cores. 1.9 References 1. Stanley, L.E., 2001, Development and Evaluation of Stitched Sandwich Panels, Master s thesis, University of Utah, Salt Lake City, UT. 2. Gharpure, S.S., 2006, Failure Mechanisms of Stitched Sandwich Composite Under Interlaminar Loading, Master s thesis, University of Utah, Salt Lake City, UT. 3. Lascoup, B., et al., 2006, On the Mechanical Effect of Stitch Addition in Sandwich Panels, Compiegne Cedex : Composites Science and Technologies, 66, Raju, I.S. and Glaessgen, E.H., 2001, Effect of Stitching on Debonding in Composite Structural Elements, Analytical and Computational Methods Branch Langley, Hampton. 5. Mouritz, A. P., 2003, Fracture and Tensile Fatigue Properties of Stitched Fiberglass Composites, Institute of Mechanical Engineers Materials: Design and Applications, Melbourne, 218 L. 6. Mouritza, A. P., et al., 1999, Review of Applications for Advanced Three- Dimensional Fibre Textile, Composites, Melbourne, A Stoll, F. and Banerjee, R., 2001, Measurement and Analysis of Fiber-Composite- Reinforced-Foam Sandwich Core Material Properties, International SAMPE Symposium and Exhibition, Long Beach, CA. 8.. Stoll, F., et al, 2004, High-Performance, Low-Cost Infusion Cores for Structural Sandwich Panels, Proceedings of SAMPE, Long Beach, CA, May Muthyala, V. D., 2007, Composite Sandwich Structure with Grid Stiffened Core, Master s thesis, Osmania University, Hyderabad, India.

48 Mouritz, A. P., 2006, Compression Properties of Z-Pinned Sandwich Composites, Journal of Material Science, Melbourne, Australia, 41: LAST-A-FOAM FR-6710 Polyurethane Foam, General Plastics Manufacturing Company, Tacoma, WA. 12. Synthetic Carbon Fiber Fabric I.O.P. Woven, Class 70, BGF Industries, Inc., Los Angeles CA. 13. BPF 862 Epon Resin, RSC 9553 Curing Agent, E.V. Roberts Corporation, Carson, CA. 14. American Society for Testing and Materials, 2004, Standard Test Method for Flatwise Tensile Strength of Sandwhich Constructions, West Conshohocken, PA, ASTM C 297/C 297M-04, American Society for Testing and Materials, Standard Test Method for Flatwise Compressive Properties of Sandwich Cores, West Conshohocken, PA, ASTM 365/C 365M-05, American Society for Testing and Materials, Standard Test Method for Shear Properties of Sandwich Core Materials, West Conshohocken, PA, ASTM C , Barbero, E. J., 1998, Introduction to Composite Materials Design, Taylor & Francis, Inc, Philadelphia, PA, pp Swanson, S. R., 1997, Introduction to Design and Analysis with Advanced Composite Materials, Prentic Hall, Upper Saddle River, NJ. Chap 2-4

49 CHAPTER 2 CREEP OF POLYMER FOAM CORE SANDWICH STRUCTURES WITH ANGLED STITCH REINFORCEMENT 2.1 Abstract Creep behavior of polyurethane foam/cfrp sandwich structures with varying stitch angle reinforcement was investigated. Stitch reinforcement was hypothesized to limit creep deformation in polymer core sandwich structures. A stitch method was developed to fabricate sandwich structures with varying stitch angles. Portable creep test fixtures were developed to test the subject sandwich specimens under flatwise compression and core shear loading at elevated temperatures. Results from flatwise compression loading indicate that 90 o stitch angles decrease creep deformation by approximately 75% when compared to unstitched sandwich specimens. Under shear loading, stitch angles of 30 o, 45 o, and 60 o decreased creep deformation by approximately 70% when compared to unstitched sandwich specimens. 2.2 Introduction Polymer foams are widely used as core materials of sandwich structures. Polymer foam cores in sandwich panels, in addition to creating a relatively high bending stiffness

50 42 per unit weight, may be easily mass-produced. Polymer foam core sandwich structures are increasingly being considered for load bearing components. However, such polymers have been known to creep and exhibit large irrecoverable strains, even at room temperature, limiting their use and life in structural applications. Creep deformation is important to consider, not only under high temperature loadings, but also in designs that use polymers at a large range of temperatures and stress levels. Much work has been performed on the topic of polymer creep [1-3] resulting in empirical models, mechanical analogs, and constitutive models. These models have been extended to the creep of polymer foams [4-5], where linear viscoelasticity theory is adapted for high-strain creep. Similar polymer creep models have also been used to predict the creep behavior of sandwich structures with foam cores in various loading conditions [6-7]. In addition to the use of monotonic foam cores in sandwich panels, through-thethickness reinforcements, such as stitching and other inserts, have been used in the past to improve sandwich core strength, stiffness, and facesheet debonding performance [8-12]. Stitching has been found to significantly increase the out-of-plane tensile strength and core shear strength [8-9]. The fiber direction also affects the creep resistance of carbon fiber/epoxy laminates, reducing the creep deformation of the polymer, provided the fiber direction is aligned in the direction of loading [13]. Therefore, it has been hypothesized that through-the-thickness stitching at angles aligned with the loading direction could reduce creep deformation. To date, no published research investigations have focused on investigating the creep behavior of stitch-reinforced foam core sandwich structures.

51 43 In this investigation, fabrication procedures were developed to introduce stitching at 30 o, 45 o, 60 o, and 90 o angles as measured from the horizontal through polyurethane foam core sandwich specimens. Creep testing was performed using specially designed fixtures that produced flatwise compression and core shear loading. These tests were used to assess the affect of stitch angle orientation on the creep behavior of sandwich structures. 2.3 Sandwich Panel Fabrication Five sandwich configurations were investigated in this study: unstitched, 30 o, 45 o, 60 o, and 90 o stitched configurations. The same materials were used in all configurations, as shown in Table 2.1. The facesheets consisted of six plies per facesheet ([0/90] 6T ) of woven carbon fabric T300B 3K [14]. The core material used was a closed cell polyurethane foam (FR-6710 from General Plastics) [15], with a density of 160 kg/m 3, and a nominal thickness of 12.7 mm. The closed cell foam was selected for its low density and the foam s ability to be stitched. Open-cell foams could not be used because resin would saturate the foam during the infiltration process. All specimens were cured using a Vacuum Assisted Resin Transfer Mold (VARTM) technique, as shown in Fig o stitched panel fabrication All stitch specimens in this study used 1600 denier Kevlar 29 yarn with a modified lockstitch. Traditional lockstitch techniques have equal tension on both the needle and bobbin threads and ideally have the locked portion somewhere near the middle of the thickness. A modified lock stitch was used, where the upper thread is

52 44 Configuration Unstitched Stitching Method N/A Hand Stitched 30 o, 45 o, 60 o 1600 denier Kevlar 29 yarn 90 o Manual Industrial Sewing Machine 1600 denier Kevlar 29 yarn Table 2.1 Material selection Facesheet Material Woven carbon fabric T300B 3K ([0/90] 6T ) [12] EPON 862 epoxy EPON 9553 hardener [13] 6 plies/facesheet Core Material Polyurethane foam (FR-6710) (General Plastics) [11] 160 kg/m mm thick Curing Method Vacuum Assisted Resin Transfer Mold (VARTM) Figure 2.1 Cross section of a VARTM infiltration assembly

53 45 pulled completely through to the bobbin side, as shown in Fig This allowed for easier fabrication as the yarn intersection was made outside the core and the yarns kept continuous through the core. The 90 o stitch specimens were stitched using a Consew industrial sewing machine for convenience and speed. A needle diameter of 2.9 mm is common for industrial sewing machines. However, previous work performed at the University of Utah [8] suggests an optimal needle diameter of 1.9 mm being large enough to perform the stitching operation without breaking while minimizing the diameter of the resin column surrounding the stitch yarn. To stitch, facesheets were temporarily taped in place while a 13 mm x 13 mm grid pattern was also temporarily taped on top of the facesheets to provide a stitch template, as shown in Fig The opposing side also had paper taped to the surface to prevent the feeder teeth from damaging the dry woven carbon fibers. Following stitching, the paper template was removed o, 45 o,60 o panel fabrication The fabrication process for the angled stitch specimens differed slightly as the stitches were hand stitched. The layup was exactly the same for the unstitched and 90 o stitched specimens. However, holes were premachined at 30 o, 45 o, and 60 o using a water jet cutting machine. All stitch rows were spaced 12.7 mm from each other with 12.7 mm spacing between stitches. The direction of the angled holes alternated from row to row. A diagram of an angle stitched panel can be seen in Fig. 2.4.

54 46 Figure 2.2 Diagram of traditional and modified lock stitch. Figure 2.3 Stitching of sandwich composites using an industrial sewing machine.

55 47 Figure 2.4 Schematic of a stitched sandwich panel A method was developed to hand-stitch angled stitch specimens with the same modified lockstitch as the 90 o specimens. To stitch, the facesheets were set up as if they were a cover for a book, as shown in Fig Modified needles, produced by cutting the needle eye hole at one side, were placed in the predrilled holes. These modified needles allowed the Kevlar yarn to pass through allowing the needle to be used as a hook, as shown in Fig Approximately 20 needles were used to stitch larger panels row by row. A step-by-step diagram is illustrated in Fig With the facesheet folded open, as shown in Fig. 2.5, needles were placed in each premade hole (Fig. 2.7 Step 1 and 2). The needles were then pushed through the other side using a thimble (Fig. 2.7 Step 3). With the needles pushed through, the cover, seen in Fig. 2.5, was then folded back over the core and the needles were pushed back through the dry facesheets, ready to be stitched at each needle (Fig. 2.7, Step 4). A picture of the stitch process at step 4 can be seen in Fig. 2.8 (left). The needle was threaded and pulled through, then removed and the

56 48 Figure 2.5 Stitch process with core and book-cover-like woven carbon fiber facesheets with needles placed in the core Figure 2.6 Modified needle for stitching 30 o, 45 o, 60 o sandwich specimens

57 49 Figure o, 45 o, 60 o sandwich specimen stitching process bobbin thread passed under the new stitch, (Fig. 2.7 Step 5). The stitch was finally pulled tight, (Fig. 2.7 Step 6). The process was then repeated and can be seen in Fig. 2.7, Step This process was repeated row by row for the entire panel. A finished panel can be seen in Fig. 2.8 (right). As the needle pushes through the foam core, cells of the closed cell foam are pierced or compressed to rupture. These damaged cells are filled with resin during the infiltration process resulting in a resin column as shown in Fig In general, the resin column diameter is proportional to the needle diameter. As stated previously, a needle Figure 2.8 Stitch process with Kevlar yarn and needle (right), typical stitched sandwich panel surface (left)

58 50 Figure o left, 60 o right, stitched sandwich panel with stitch and resin column diameter of 1.9 mm was selected to stitch the 90 o specimens using an industrial sewing machine [1]. The predrilled holes for the 30 o, 45 o, or 60 o stitched sandwich specimens, on the other hand, measured 1.3 mm. When the modified needles with Kevlar were pulled through the core by hand for the angled stitches, the hole would expand and a comparable resin column resulted for the 90 o stitch specimens and the 30 o, 45 o, and 60 o stitch specimens. It is noted that the resin column diameters varied significantly within a stitch and from stitch to stitch. Areal density measurements were taken for each stitch configuration shown in Table 2.2. Areal density measurements reveal slight increases in density for stitched specimens. It was expected that the stitched specimens would have higher areal densities as the Kevlar and resin are denser than the polyurethane foam core. Variations in density between the stitched specimens are believed to be from inconsistencies in VARTM infiltration where air could possibly have been trapped at some stitch locations. A 60 o panel with the foam removed can be seen in Fig

59 51 Table 2.2 Comparison of areal densities for varying stitch angles Configuration Areal Density g/m^2 Normalized Areal Density Unstitched 6740 ± Degree 7060 ± Degree 6920 ± Degree 6970 ± Degree 7300 ± Figure 2.10 Typical 60 o stitch panel with foam removed 2.4 Flatwise Compression Creep Testing Flatwise compression creep test method The previously described five configurations of stitched sandwich structures were creep tested under flatwise compression loading in accordance with a quasi-static flatwise compressive test method, ASTM C [17]. A test specimen size of 51 mm x 51 mm was selected to include stitch reinforcement within the sandwich specimen. Specimens were cut using a water cooled diamond saw. Since all five configurations were to be

60 52 creep tested at the same stress level, preliminary tests were conducted and it was determined that approximately, a compression stress level of 1.4 MPa was well-suited for creep testing. Thus, a compressive force of approximately 3.6 kn of force was required for creep testing. Further, an elevated temperature of 82 o C was deemed to be appropriate for creep testing given the glass transition temperature (138 o C) and the desired amount of creep. The creep tests were run for 150 hours. A total of six specimens were tested from each of the five sandwich configurations. A Blue-M convection oven was used to maintain the test specimens at a constant temperature of 82 C. It was desired that six specimens be tested simultaneously. With the required load level of 3.6 kn, however, a traditional creep test method using deadweight testing was deemed impractical. Thus a compact spring-based testing apparatus was designed for use in the available oven space Flatwise compression creep test fixture A spring-loaded creep test fixture was specifically developed to test the reinforced stitched sandwich specimens at the required high loads. In addition to providing the relatively high loads, the creep test fixtures needed to be relatively low cost while allowing for load to be applied accurately and with minimal change during creep testing. Finally, the fixtures needed to be compact enough such that multiple fixtures could be fit into the available convection oven. Several possible approaches were considered. The design finally selected for use is based on a spring loaded lever arm arrangement, shown in Fig The compression spring, capable of a maximum load of 5.3 kn and load rate of 36 N/mm, uses a centrally

61 Figure 2.11 Compression creep test fixture diagram 53

62 54 threaded rod and nuts to apply and maintain a compression load on the specimen. The entire load frame was sized to be as small as possible, as six test fixtures were to be fabricated and fit inside a 900 mm wide, 1200 mm tall, and 1500 mm deep laboratory oven. Some details of the test fixture are as follows (refer to Fig. 2.11). Two box beams, 38 mm x 100 mm and 3.2 mm thick, were compressed using a threaded rod and compression spring. Translation of a hex nut placed between the two box beams shifted the load from the threaded rod to compressing the two box beams and onto the test specimen. Dial indicators of mm resolution were used to monitor the out-of-plane displacement of the specimen. A modified oven door was created to view the dial gages during testing. A high resolution camera and intervalometer (a timer device used for time lapse photography) were used to capture images of the dial gages every hour. A still frame taken of a test in progress is shown in Fig An inherent difficulty with the outlined approach is that any creep response of the specimen will cause the load applied by the spring to change. One way to reduce load change is to use a spring with the lowest spring constant possible, such that changes to the spring deflection due to specimen creep will result in minimal changes in the applied load. Hence, a relatively low stiffness spring of 29 N/mm with adequate maximum load of 5.3 kn was selected for use. The springs selected were approximately 100 mm in diameter and approximately 300 mm tall compressing to less than half of its nominal height. All springs were purchased from Placement of the spring between the specimen and the hinge resulted in a mechanical disadvantage of

63 55 Figure 2.12 Time lapse frame of compression creep test approximately 0.8, in effect decreasing the spring rate. Thus, for an applied specimen load of 3.6 kn, a 4.5 kn load was required from the spring. Due to small load change that accompanies specimen creep, it was desired to monitor load during creep testing. To accurately monitor the load applied by the spring on the specimen, two biaxial ±45 o strain gages were applied to a coil on the spring. The strain gage placement compensated for any off-axis loading of the spring. Off-axis loads would cause one side of the spring to compress more; however, loads were calculated from measurements from the sum of the two gages which results in a more accurate total load applied by the spring. A full Wheatstone bridge wiring setup of the gages also compensated for thermal expansion of the spring and gages. Each spring was individually calibrated using a 44 kn electromechanical testing machine. The spring calibration exhibited extremely linear behavior and resulted in a resolution of approximately 2 N or 0.05% of the total load.

64 Flatwise compression creep test results Flatwise compression creep test results are presented in Fig and Table 2.3. Of all sandwich configurations investigated, the 90 o stitched specimens yielded the least creep deformation, followed by the 60 o, 45 o, and 30 o stitch orientations. The 90 o stitches were found to decrease total creep deformation by approximately 75% when compared to the unstitched specimens. Each configuration, except for the 30 o stitch orientation, exhibited an initial linear strain versus time response, followed by a secondary region characterized by a lower slope. Following testing, the maximum spring load loss recorded was 3.7%. The recovery column in Table 2.3 represents the specimen thickness measured several months after testing and the specimen thickness at the peak of creep deformation. Therefore, the recovery column represents both the elastic and viscoelastic components of deformation recovered after testing. From the recovery presented, the 90 o panel recovers approximately 50% its total displacement, whereas the unstitched specimens recovered only 25% of its total displacement. Real-time recovery was not measured due to difficulties in quickly removing all loads from the test fixtures and monitoring their displacements with time. 2.5 Core Shear Creep Testing Core shear creep test method Core shear creep testing was performed in a similar manner to that specified in the quasi-static shear test method, ASTM C [18]. Stress was calculated as shown in Eq. 2.1:

65 57 Figure 2.13 Strain versus time for flatwise compression creep tests Table 2.3 Results for flatwise compression creep tests Initial Static Displacement (mm) Displacement after 150 hours (mm) Creep Displacement (mm) Total Recovery (mm) Test Fixture Spring Load Loss % Unstitched 1.3 ± ± ± ± o 1.4 ± ± ± ± o 1.2 ± ± ± ± o 1.2 ± ± ± ± o 1.0 ± ± ± ±

66 58 τ= (2.1) where P is the load, L is the length of the specimen, and b is the width of the specimen. Preliminary testing indicated a shear stress of approximately 690 kp was well suited for core shear creep testing. Core shear creep tests were conducted with one variation in the length requirement from the quasi-static core shear standard. ASTM C [18] requires the length of the sandwich specimen to be a minimum of 12 times the thickness. With a sandwich thickness of approximately 15 mm, a minimum length of 183 mm was needed to meet the standard. Creep testing specimens that were 183 mm x 51 mm resulted in springs loads greater than 8 kn which was deemed too high for commercially available springs. The specimen length was therefore shortened to 152 mm. This new size required spring loads of approximately 6.8 kn which was sufficiently low for commercially available springs with relatively low spring constants. This decreased load also allowed for the use of existing fixture materials, the same used in the flatwise compression creep test fixtures. Specimens were cut to the specified dimensions using a water cooled diamond saw. Similar to the flatwise compression creep tests, five configurations of stitched sandwich structures previously described were tested under core shear creep at an elevated temperature and duration of 82 o C and 150 hours, respectively. Six tests were performed from each of the five configurations; six specimens were tested simultaneously. A Blue-M laboratory oven was used to maintain the test specimens at a

67 59 constant temperature of 82 C where time lapse images could be taken through a modified door Core shear creep test fixture development The flatwise compression creep test fixture design was modified for core shear testing. A larger spring, with a maximum load of 7.7 kn, was moved between the two box beams, changing the load on the test specimen from compression to tension, as shown in Fig The spring, with centrally threaded rod and nuts, was used to apply and maintain tension load on the specimen. The load rate of this spring was approximately 58 N/mm which is higher than that of the flatwise compression spring. However, lower deflections were also expected, and thus, load loss due to specimen deflection was minimized. The same biaxial strain gage placement was used and a calibration of each spring was also performed on a 44 kn load frame. Specimens were prepared and adhesively bonded to each shear plate using Hysol 9394 high temperature epoxy. To measure displacement during the test, dial gages of mm resolution were placed on magnetic bases and measured the relative axial motion occurring between the opposing faces of the sandwich composite specimen. A high resolution time lapse set-up was used to monitor the dial indicators where pictures were taken every hour. Dial readings were then recorded after testing. As shown in Fig. 2.15, all six creep fixtures were placed in a large convection oven with a modified door for viewing.

68 Figure 2.14 Core shear test fixture diagram 60

69 61 Figure 2.15 Time lapse frame of core shear test Core shear creep test results Core shear creep test results are presented in Fig and Table 2.4. It is noted that one dial gage was defective for the 45 o, 60 o, and 90 o tests, and thus, only five test results are reported for these configurations. The strain versus time plots in Fig show a higher initial creep rate followed by a lower secondary creep rate for all configurations tested. As expected, the unstitched specimens followed by the 90 o stitch specimens resulted in the greatest creep deformation. However, the 30 o, 45 o, and 60 o stitch sandwich specimens all deformed similarly. This result was not expected, as the 45 o stitch angle under shear loading would produce the maximum tensile and compressive stresses on planes oriented at 45 o. However, both the 45 o and 60 o stitch orientation reduced creep by approximately 70% when compared to the unstitched

70 62 Figure 2.16 Strain versus time for core shear creep tests Table 2.4 Results for core shear creep tests Initial Static Strain Total strain After 150 hours Creep (Strain) Test Fixture Spring Load Loss % Unstitched 0.05 ± ± ± Degree 0.02 ± ± ± Degree 0.02 ± ± ± Degree 0.02 ± ± ± Degree 0.03 ± ± ±

71 63 specimens. The unstitched specimens experienced the maximum load loss observed; however, the maximum decrease in load was only 1.2%. No creep recovery was measured due to the difficulties in unloading the test specimen. 2.6 Discussion Generally accepted linear viscoelastic models are said to be valid for loads below 50% of the materials compressive strength [19]. According to Huang and Gibson [4], viscoelastic models have been found to be increasingly nonlinear above 50% of the compressive strength, as other deformation mechanisms may be occurring. In this investigation, creep tests were performed at approximately 54% of compressive strength for the unstitched sandwich configurations. These loads were initially selected to result in a measureable amount of creep in the most creep resistant stitch configuration, the 90 o stitch panel. The resulting amounts of creep deformation would generally be considered high for actual applications. The strains resulting from creep reach nearly 50% strain for the unstitched configurations. The highest load loss recorded from the unstitched panel was 3.7% at approximately 6.1 mm of deformation. To decrease load loss and remain in the valid linear viscoelastic model stress limits, lower stresses are recommended for future tests as it would also resemble more realistic applications where creep deformations may be a concern. Real-time creep recovery was not measured due to the difficulty in completely removing load and accurately measuring recovery in a timely fashion. Total recovery, however, was measured several months after testing. Recovery data indicated that the 90 o stitch specimens recovered almost 50% of total displacement whereas the unstitched specimens recovered only 25% of total creep displacement. The lower recovery

72 64 displayed by the 45 o, followed by the 30 o, and unstitched configurations may be an indication that other deformation mechanisms such as cell crushing may be involved in the total deformation measured. Core shear creep test results revealed that the 30 o, 45 o, and 60 o stitch orientations performed similarly and were most resistant to creep deformation. It is noted that due to the alternating stitch pattern used, only half the stitch reinforcement is effectively loaded in tension and half is loaded in compression. Shear strains resulting from the core shear creep tests were greater than 12% shear strain. The shear stress applied for testing was approximately 82% of the unstitched core shear strength. At these stress levels, linear viscoelastic models are also not considered accurate. These load levels were selected to result in measureable amounts of creep deformation in the most creep resistant stitch configuration (45 o ). It is recommended that future core shear creep tests be performed at lower stress levels which would allow linear viscoelastic models to be more accurate, lower load loss, and better simulate conditions found in actual applications. In designing for creep considerations, the most important parameters are stress, temperature, and time. In this study, additional considerations that may have affected creep results include quality of stitch infiltration, stitch tension, and stitch resin column diameter variation. Stitch infiltration was not explored, and would require harvesting stitches and performing density determinations. Automated stitching methods may have remove most stitch tension variation as well as the resin column diameter variation as compared to a hand stitching technique.

73 Conclusions and Recommendations This investigation included the development of stitching and fabrication methods to produce 30 o, 45 o, 60 o, and 90 o angled stitch orientations in foam core sandwich panels. Stitching at 90 o was found to greatly reduce creep under flatwise compression creep testing by approximately 75% when compared to unstitched specimens. The 30 o, 45 o, and 60 o stitch sandwich specimens performed comparably under core shear creep testing and reduced creep by approximately 70% when compared to the unstitched sandwich specimens. These large creep reductions were recorded with only up to 7% increase in areal density for the addition of stitch reinforcement. Stitch reinforcements may be further optimized for use in specific applications, further reducing increases in weight and cost. Creep test methods were developed using specially designed test fixtures for flatwise compression creep and core shear loading. The spring-loaded test fixtures were able to sustain the required loads with minimal load loss and permitted six creep tests to be performed simultaneously in a large convection oven. The spring elements were instrumented with strain gages to allow for accurate load application as well as load loss monitoring. The results of this investigation suggest that stitch reinforcement may be used to greatly reduce creep deformation in foam core sandwich panels. Depending on the design and loading conditions expected, stitch reinforcement may be optimized according to angle and stitch density to improve creep performance and limit weight addition. Stitching could be limited to creep-prone locations or combinations of stitch angles could be implemented for complex loading situations.

74 66 Creep behavior of stitched sandwich composites is related to the geometry of the stitch angle as well as the properties of the constituents. Traditional creep models like the Findley power law [21] have been shown to provide accurate predictions for linear viscoelastic materials. For future work, a model may be developed relating creep, stress, and temperature based on linear viscoelasticity theory. It is often impractical to test long-term behavior of materials directly with experiment because of the time required. Thus, predicting long-term creep using shortterm tests may be extremely useful. One of the most common techniques, Time Temperature SuperPosition (TTSP), has been used for polymers by shifting the curves from tests at different temperatures horizontally along a logarithmic time axis to generate a single master curve. Thus, a long-term test may be replaced by several short-term tests. Future work in this area would be of interest, as it is hypothesized that the creep of each constituent material would be different at different temperatures. The interactions of the constituents may lead to nonlinear behavior. Finally, creep fatigue interactions would be an interesting topic of study for stitch reinforced polymer core sandwich structures. As these sandwich structures are used in more diverse applications, including those subjected to fatigue loading, knowledge of their time-dependent behavior becomes increasingly important. It is known that cyclic loading may increase creep of certain polymers [20]. Further, stitch reinforcements may be viewed as stress concentrations, and thus may affect the performance of sandwich structures under fatigue loading.

75 References 1. ASM International, 2003, Characterization and Failure Analysis of Plastics Library of Congress Catalog-in-Publication Data, s.l. 2. American Society for Testing and Materials, 2009, Standard Test Methods for Tensile, Compressive, and Flexural Creep and Creep-Rupture of Plastics, West Conshohocken, PA, ASTM D Wineman, A. S., 2000, Mechanical Response of Polymers, Cambridge University Press, Cambridge, UK. 4. Huang, J. S. and Gibson, L. J., 1991, Creep of Polymer Foams, Journal of Materials Science, 26, Zhu, H. X. and Mills, N. J., 1999, Modelling the Creep of Open-Cell Polymer Foams, Journal of the Mechanics and Physics of Solids, 47, 7, (21). 6. Huang, J. S. and Gibson, L. J., 1990, Creep of Sandwich Beams with Polymer Foam Cores, Journal of Materials in Civil Engineering. 7. Shenoi, R. A., Allen, H. G. and Clark, S. D., 1997, Cyclic Creep and Creep-Fatigue Interaction in Sandwich Beams, Journal of Strain Analysis 8. Stanley, L.E., 2001, Development and Evaluation of Stitched Sandwich Panels, Master s thesis, University of Utah, Salt Lake City, UT. 9. Lascoup, B., et al., 2006, On the Mechanical Effect of Stitch Addition in Sandwich Panels, Compiegne Cedex : Composites Science and Technologies, 66, Gharpure, S.S., 2006, Failure Mechanisms of Stitched Sandwich Composite Under Interlaminar Loading, Master s thesis, University of Utah, Salt Lake City, UT. 11. Mouritz, A. P., 2006, Compression Properties of z-pinned Sandwich Composites, Journal of Material Science, Melbourne, Australia, 41: Raju, I.S. and Glaessgen, E.H., 2001, Effect of Stitching on Debonding in Composite Structural Elements, Analytical and Computational Methods Branch Langley, Hampton. 13. Pang, F., Wang, C. H. and Bathgate, R. G., 1997, Creep Response of Woven-Fibre Composites and the Effect of Stitching, Composites Science and Technology, Geelong, Synthetic Carbon Fiber Fabric I.O.P. Woven, Class 70, BGF Industries, Inc., Los Angeles CA.

76 LAST-A-FOAM FR-6710 Polyurethane Foam, General Plastics Manufacturing Company, Tacoma, WA. 16. BPF 862 Epon Resin, RSC 9553 Curing Agent, E.V. Roberts Corporation, Carson, CA. 17. American Society for Testing and Materials, Standard Test Method for Flatwise Compressive Properties of Sandwich Cores, West Conshohocken, PA, ASTM 365/C 365M-05, American Society for Testing and Materials, Standard Test Method for Shear Properties of Sandwich Core Materials, West Conshohocken, PA, ASTM C , Yourd, R. A., 1996, Compression Creep and Long-Term Dimensional Stability in Appliance Rigid Foam, Journal of Cellular Plastics, Pittsburgh, PA, Vinogradov, A. M., 2003, Creep-Fatigue Interaction in Polymers, American Society of Civil Engineers 16th Engineering Mechanics Conference Seattle. 21. Findley, W. N., 1976, Creep and Relaxation of Nonlinear Viscoelastic Materials, North-Holland Publishing Company Toronto, Xu, Y., 2009, Creep Behavior of Natural Fiber Reinforced Polymer Composites, Ph. D. thesis, Louisiana State University.

77 CHAPTER 3 PORTABLE SPRING LOADED CREEP TEST FIXTURES FOR SANDWICH COMPOSITES 3.1 Abstract Portable spring-loaded creep test fixtures were developed to test sandwich composite specimens under both flatwise compression and through-thickness shear loading. These low-cost creep test fixtures were able to sustain required loads with minimal load loss, and permitted six creep tests to be performed simultaneously in a large convection oven. Accurate load and creep deformation measurements were possible through the use of strain gages and dial gages, respectively. Successful use of the test fixtures was demonstrated through creep testing of polymer foam core sandwich composites reinforced with Kevlar stitching. 3.2 Introduction Polymer foams are widely used as core materials in sandwich structures due to their relatively low cost, low density, and ease of manufacture. However, polymer foams used as cores of sandwich composites are especially vulnerable to creep when subjected to through-the-thickness tension, compression, or shear loading. Creep deformations

78 70 usually results from sustained loading at elevated temperatures. For polymers, however, creep can occur at relatively low temperatures [1]. Time-dependent creep behavior of polymeric composites has been studied for many years [3]. As composites become more widely used for safety critical components, knowledge of long-term loading behavior also becomes increasingly important. In recent years, core reinforcements have been developed and used to increase sandwich panel strength and modulus under out-of-plane tension, compression, and shear loading. An example of one such core reinforcement, angled stitching, is shown in Fig Insufficient information is known about the added resistance to creep deformation produced from such core reinforcements. Thus, an investigation was performed to investigate the creep performance of stitch reinforced sandwich composites under different loading conditions. As a result, it was necessary to develop creep test fixtures to produce flatwise compression and through-thickness shear loading of composite sandwich specimens at elevated temperatures. The test fixtures were required to sustain the applied load with minimum inherent load loss due to specimen creep, be cost effective, and be compact such that multiple fixtures could be placed into an environmental test chamber for simultaneous testing. Figure o stitch reinforcement in polymer core sandwich structure

79 Test fixture development Numerous creep test methods currently exist. Creep test methods that require the application of relatively high loads commonly are performed using servo-hydraulic or electromechanical test machines equipped with environmental test chambers. Additionally, creep tests are often performed using dead weight loading; however, such tests are limited to a practical volume of weights that can be used. Devices with mechanical advantage and dead weights have also been used within environmental chambers with some increases in maximum load due to a lever arm, but are limited by space [4]. For both flatwise compression and through-thickness shear testing of sandwich composites, standard test methods exist for quasi-static loading, complete with prescribed specimen dimensions and loading methodologies. Since no test methods exist for creep loading, the existing specimen dimensions and loading methods were used. The use of such specimen dimensions led to required load levels that were deemed too large for deadweight loading. Thus, compact spring-based creep testers were developed for use in a large convection oven. Since creep tests usually run for long periods of time, it is desirable to test multiple specimens simultaneously. A review of the literature revealed that specialized spring-loaded fixtures had been developed in several cases. An ASTM standard creep test method for use with concrete [6] consists of two springs on either side of a concrete specimen mounted on a large frame. Awal used a modified version of this test method [7] which consisted of a large load frame, disc springs, and a hydraulic jack to maintain load on the concrete specimens. This setup is not suitable for elevated temperature

80 72 testing due to the hydraulic jack. Spring-loaded fixtures have also been developed for use with composite laminates where Henshaw et al. [8] developed a fixture which consisted of a tension spring and fulcrum, compressing laminates between two plates mounted on linear bearings. This approach uses a lever arm which increases the effective spring rate which inherently increases load loss due to creep. Tuttle et al. [9] also developed a spring and fulcrum test device for use applying tensile stresses on composite laminates. Similar to the Henshaw et al. design, a lever arm also increased mechanical advantage effectively increasing spring rate. Also, the placement of the specimen required a relatively long and slender specimen, not suitable for out-of-plane tests of sandwich specimens. Finally, a spring-loaded test apparatus was developed to measure degradation and creep of polymers [10]. This design consisted of a base and springs with centrally threaded rod which applied tensile and/or compressive stresses on the specimen. The base was specially developed to contain liquid for degradation of the polymer. No mechanical advantage was utilized which results in load loss directly proportional to spring rate. Since these fixture designs were specialized for their respective circumstances, development of a new, spring-loaded fixture design was required to creep test polymer core sandwich specimens. An inherent difficulty with using spring-loaded test fixtures for creep testing is that any creep deformation produced in the specimen will cause the load applied by the spring to change. However, it is desired that the fixture sustain a constant load during the test. Thus, while needing to be cost effective, and compact, the test fixture design also required that a less than 5% load change occur during creep testing of sandwich composites.

81 Flatwise compression creep testing It was desired to perform flatwise compression creep testing in a similar manner to that used for quasi-static testing, as specified in ASTM C 365 [11]. This test method specifies that the minimum specimen cross section should be 25 mm x 25 mm. To investigate the effects of stitch reinforcements, however, a 50 mm x 50 mm specimen size was required. Based on preliminary testing performed using this specimen size, a compression load of approximately 3.6 kn was deemed necessary for flatwise compression creep testing. Further, simultaneous testing of multiple testing was desired at elevated temperature using a large convection oven. These requirements led to the development of the spring-loaded test fixture shown in Fig and Fig This fixture consisted of two 38 mm x 102 mm rectangular steel box beams connected at one end with a hinge, as shown in Fig The specimen was placed at the opposite end of the 610 mm long steel box beams. The compression spring selected was capable of a maximum load of 5.3 kn and had a spring constant of 36 N/mm. A central threaded rod and nuts were used to compress the spring and apply and maintain a compression load on the specimen. To reduce the amount of spring displacement and thus load drop associated with specimen creep deformation during testing, the spring was placed off to one side of the center of the fixture. This positioning resulted in a mechanical disadvantage, such that a force equal to 0.81 times the spring force was applied to the specimen. The spring positioning was selected to ensure that the maximum required specimen load could be provided using the selected coil spring. A summary of the fixture loading capabilities is provided in Table 3.1. Spacers were used at the hinged ends of the beams to ensure that the beams

82 Figure 3.2 Compression creep test fixture diagram 74

83 75 Figure 3.3 Flatwise compression creep test fixture Table 3.1 Test fixture load summaries Maximum load capacity of coil spring Mechanical disadvantage produced by test fixture Maximum specimen force provided by coil spring Flatwise Compression Creep Test Fixture Through-Thickness Shear Test Fixture 5.3 kn 7.7 kn kn 6.8 kn Spring constant of coil spring 36 N/mm 58 N/mm Effective spring constant produced at specimen 29 N/mm 45 N/mm Maximum load loss experienced 1.6% 0.96%

84 76 were parallel for a given thickness of sandwich specimen. Dial indicators with a mm resolution were used to measure the out-of-plane displacement of the specimen during testing. To accurately monitor the load applied by the spring onto the specimen, two biaxial ±45 degree strain gages were applied to coils of the spring. The strain gage placement compensated for any off-axis loading of the spring. A full Wheatstone bridge wiring setup was used which provided for temperature compensation. Each spring was calibrated individually using a 44 kn electromechanical testing machine, as shown in Fig. 3.4 (left). The spring calibration exhibited linear behavior in the spring up to peak operational load of 4.5 kn, as shown in Fig. 3.4 (right). The resulting strain resolution obtained during testing was approximately 2 N or 0.05% of the total load. Figure 3.4 Spring calibration in 44 kn load frame (left), Load versus strain results (right)

85 77 As creep deformation occurs in the specimen during the test, a small angle is produced between the two box beams of the test fixture as well as a reduction in the compression force applied by the coil spring. However, due to the length of the beams, the maximum angle change is small (less than 0.2 o for 2.5 mm of specimen displacement). Due to the spring placement between the hinge and the specimen and the resulting mechanical disadvantage, the effective spring constant produced by the fixture at the location of the specimen is decreased. As a result, the load drop resulting from specimen deformation is reduced. The maximum load loss recorded during any flatwise compression creep testing performed to date was 1.6% of the total load. Prior to testing, all springs were loaded against their central threaded rod and preheated for at least 24 hours. From preliminary testing, it was determined that approximately 90 N of additional load was needed to be loaded on the spring to account for fixture compliance at the start of testing. To apply the load to the specimen at the start of the test, one nut was simply loosened from the threaded rod, transferring load from the threaded rod to the box beams. To monitor displacement measurements from the dial gages, a high resolution digital camera was set up. Using an intervalometer, time lapse images were taken every hour. Dial readings were then recorded after testing from the time lapse images. As shown in Fig. 3.5, all six creep fixtures were placed in a Blue- M laboratory convection oven with a modified door for viewing the dial indicators during the test.

86 78 Figure 3.5 Time lapse frame of six simultaneous compression creep tests 3.5 Through-thickness shear creep test method Similar to flatwise compression loading, the creep test method for throughthickness shear loading was patterned after that used for quasi-static testing. ASTM C 273 [12] describes a through-thickness shear test method used for sandwich composites and core materials. The length of the specimen specified in the standard is at least 12 times the specimen thickness and the width not less than 50 mm. This relatively large specimen would require a significantly higher applied load than possible using a coil spring similar to that used for flatwise compression loading. As a result, a shorter specimen length of 152 mm (10 times the specimen length) was adopted for shear creep testing. The specimen width was retained at 51 mm. The flatwise compression creep test fixture design was modified for throughthickness shear testing, as shown in Fig. 3.6 and 3.7. The spacing between the box beams

87 Figure 3.6 Through-thickness shear test fixture diagram 79

88 80 Figure 3.7 Through-thickness shear creep test fixture was widened to accommodate the different orientation of the longer test specimen. Two hinges connected by a steel plate were used to join the steel tubes and provide the required movement at the shear specimen. A larger spring, with a load capacity of 7.7 kn, was moved to a position between the box beams, changing the load on the test specimen from compression to tension. Additionally, larger threaded rods were selected as higher loads were expected. Although the spring constant for this spring, 58 N/mm, was higher than that used for flatwise compression, lower deflections were also expected. The same biaxial strain gage placement was used, and each spring was calibrated individually. A summary of loading capabilities for the through-thickness shear test fixture is provided in Table 3.1. Prior to testing, specimens were adhesively bonded to steel shear loading plates using Hysol 9394 high temperature epoxy. To measure the shear deformation produced