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

Copyright © Thien Sok 2010

All Rights Reserved

T h e U n i v e r s i t y o f U t a h G r a d u a t e S c h o o l

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.

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, 30o, 45o, 60o, and 90o 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 90o were found to nearly double out-of-plane tensile stiffness and

strength while compressive stiffness and strength increased only modestly. Stitches

oriented at 30o and 45o 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 90o as the optimal stitch

orientation for flatwise tension stiffness; however, 35.2o 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

iv

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 90o decreased creep

by approximately 75% when compared with unstitched sandwich specimens.

Measurements of creep under core shear loading indicate stitch angles 30o, 45o, and 60o

decreased creep by approximately 70% when compared to unstitched sandwich

specimens.

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 ................................................................ 4

1.1 Abstract ............................................................................................. 4 1.2 Introduction ....................................................................................... 5 1.3 Sandwich Panel Fabrication .............................................................. 6 1.4 Flatwise Tension Testing ................................................................ 14 1.5 Flatwise Compression Testing ........................................................ 19 1.6 Core Shear Testing .......................................................................... 21 1.7 Analytical Modeling ....................................................................... 27 1.8 Discussion and Conclusions ............................................................ 37 1.9 References ....................................................................................... 39

2. CREEP OF POLYMER FOAM CORE SANDWICH STRUCTURES WITH ANGLED STITCH REINFORCEMENT ....................................... 41

2.1 Abstract ........................................................................................... 41 2.2 Introduction ..................................................................................... 41 2.3 Sandwich Panel Fabrication ............................................................ 43 2.4 Flatwise Compression Creep Testing.............................................. 51 2.5 Core Shear Creep Testing ............................................................... 56 2.6 Discussion ....................................................................................... 63 2.7 Conclusions and Recommendations ............................................... 65 2.8 References ....................................................................................... 67

3. PORTABLE SPRING LOADED CREEP TEST FIXTURES FOR SANDWICH COMPOSITES ............................................................ 69

3.1 Abstract ........................................................................................... 69 3.2 Introduction ..................................................................................... 69

vi

3.3 Test fixture development................................................................. 71 3.4 Flatwise compression creep testing ................................................. 73 3.5 Through-thickness shear creep test method .................................... 78 3.6 Creep Testing of Composite Sandwich Specimens ........................ 81 3.7 Summary ......................................................................................... 83 3.8 References ....................................................................................... 84

LIST OF TABLES

Table

1.1 Material selection ......................................................................................................... 7

1.2 Comparison of areal densities for varying stitch angles ............................................. 14

1.3 Comparison of flatwise tension tests results for varying stitch angles ...................... 18

1.4 Comparison of flatwise compression tests results for varying stitch angles ............. 22

1.5 Comparison of core shear test results for varying stitch angles ................................. 26

1.6 Material and geometry properties used in modulus models ...................................... 33

2.1 Material selection ....................................................................................................... 44

2.2 Comparison of areal densities for varying stitch angles ............................................. 51

2.3 Results for flatwise compression creep tests ............................................................. 57

2.4 Results for core shear creep tests ............................................................................... 62

3.1 Test fixture load summaries ........................................................................................ 75

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 fiber-

based 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

2

topic of varying the angle of stitch reinforcements. It was hypothesized that stitch angles

more closely aligned with 45o 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 45o 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-of-

plane 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, 30o, 45o, 60o, and 90o 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

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.

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 90o were found to greatly increase out-of-plane tensile and

compressive modulus as well as tensile strength. Stitches oriented at 45o and 30o were

found to produce the greatest increase in shear modulus and strength.

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

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, 30o, 45o, 60o, and 90o 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 45o 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 45o 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 out-

of-plane tensile modulus and shear modulus.

1.3 Sandwich Panel Fabrication

1.3.1 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/m3 made by General Plastics LAST-A-FOAM FR-6710

7

Table 1.1 Material selection

Configuration Stitching Method

Facesheet Material

Core Material Curing Method

Unstitched N/A

Woven carbon fabric

T300B 3K ([0/90]6T)

[12]

EPON 862 epoxy

EPON 9553 hardener

[13]

6 plies/facesheet

Polyurethane foam (FR-6710) (General Plastics)

[11]

160 kg/m3

12.7 mm thick

Vacuum Assisted

Resin Transfer

Mold (VARTM)

30o, 45o, 60o Hand Stitched 1600 denier

Kevlar 29 yarn

90o

Manual Industrial

Sewing Machine 1600 denier

Kevlar 29 yarn

[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,

30o, 45o, 60o stitch specimens, and 90o stitch specimens, and are described in the next

three sections.

1.3.2 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

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. 1.1.

1.3.3 90o stitched panel fabrication

90o 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. 1.2. 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.

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.

10

Figure 1.3 Diagram of traditional and modified lock stitch.

1.3.4 30o, 45o,60o 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 90o stitched panels.

However, the foam had predrilled holes at 30o, 45o, and 60o. 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. 1.4.

A method was developed to hand-stitch the angled stitch panels with the same

modified lockstitch as the 90o panels. Prior to stitching, needles were modified as shown

in Fig. 1.5. 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 pre-

drilled 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

11

Figure 1.4 Schematic of a stitched sandwich panel

Figure 1.5 Modified needle for stitching 30o, 45o, 60o sandwich panels

12

Figure 1.6 Stitch process with core and book-cover-like woven carbon fiber facesheets with needles placed in the core

Figure 1.7 30o, 45o, 60o, 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

13

process is then repeated and can be seen in Fig. 1.7, step 7-10. 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. 1.9. 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 30o, 45o, or 60o 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 90o stitch panels and the 30o, 45o, and 60o 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)

14

Figure 1.9 90o (left), 60o (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±82 1.00

90 Degree 7060±44 1.04

60 Degree 6920±75 1.02

45 Degree 6970±180 1.03

30 Degree 7300±170 1.07

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 60o panel with the foam removed can be seen in Fig. 1.10.

1.4 Flatwise Tension Testing

1.4.1 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. 1.11 shows a typical test setup.

15

Figure 1.10 Typical 60o 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

16

specimens containing 300, 450, and 60o 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.

1.4.2 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 90o coupons resulted in one catastrophic

failure, as shown in Fig. 1.12. Failure of the unstitched specimens resulted in a complete

separation into two halves. 90o stitch panels for most tests exhibited the same failure

17

Figure 1.12 Failure in a flatwise tension test of an unstitched (left), 90o (right) sandwich panel.

mode; stitches and foam would fail simultaneously. For the 30o, 45o, and 60o 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.

1.13 represents a typical 30o panel failure compared with a crosshead versus load plot,

where the first initial load drop results in fracture of the core. As the 30o 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.

1.4.3 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

18

Figure 1.13 Failure progression in flatwise tension test of a 30o 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±0.14 ---- 101±4.3 ----

30o 1.27±0.21 22 112±5.2 11

45o 1.58±0.36 52 133±19 31

60o 1.88±0.14 81 155±14 53

90o 2.23±0.23 110 194±18 91

19

1.2), the stitch contribution is great. The modulus for the 90o stitch panel is nearly double

and the maximum stress attained is over twice that of the unstiched panel.

The 30o, 45o, and 60o stitch panels do not fail completely with the initial load

drop, as seen in Fig. 1.14. 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

1.5.1 Flatwise compression test method

Flatwise compression tests were carried out in accordance with ASTM C365-05

[15]. Because of the similarities in testing between flatwise tension and compression,

similar performances were hypothesized. A flatwise compression test is shown in Fig.

1.15. 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.

1.5.2 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.

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

21

Figure 1.16 Stress versus crosshead displacement plots of flatwise compression test for varying stitch angles

1.5.3 Flatwise compression test results

As predicted, in flatwise compression, the sandwiches with the 90o 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 30o and the 45o stitched

panels, and even a decrease in modulus for those configurations. Having the 30o stitches

actually decreased the modulus by almost 17%.

1.6 Core Shear Testing

1.6.1 Core shear test method

Interlaminar shear tests were conducted in accordance to ASTM C 273-00 [16].

A series of coupons, 51 mm x 203 mm, were prepared from the same five stitch

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±0.03 ---- 45.7±1.1 ----

30o 2.68±0.05 3.4 38.1±2.3 -17

45o 2.81±0.04 8.3 44.3±0.4 -3.2

60o 3.02±0.15 16 47.0±2.6 2.7

90o 4.16±0.18 60 56.2±1.3 23

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. 1.17 (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. 1.17 (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:

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.

1.6.2 Core shear failure modes Two main types of failure occurred for the differing stitch configurations. For the

unstitched and the 90o, failure resulted in a shearing and complete separation of the two

halves, as shown in Fig. 1.18.

Failure of the 30o, 45o, and 60o panels resulted in foam failure with a large drop in

load while the two halves remain together, as seen in Fig. 1.19. In the 30o, 45o, and 60o

24

Figure 1.18 Failure in a core shear test of an unstitched (left) and a 90o (right) sandwich panel

Figure 1.19 Failure in a core shear test of a 30o 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. 1.20 (left) where the Kevlar yarn is severed near the bottom facesheet and

in Fig. 1.20 (right), buckling of the stitches can be seen near the centers of each stitch.

1.6.3 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 45o direction.

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. 1.21. 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.

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±0.03 ---- 47.8±6.2 ----

30o 1.68±0.03 100 75.8±4.1 58

45o 1.62±0.09 93 81.1±8.8 70

60o 1.54±0.06 83 59.3±6.2 24

90o 1.51±0.002 79 52.6±6.7 10

Figure 1.21 Shear stress versus crosshead displacement plots of core shear tests for varying stitch angles

27

1.7 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-of-

materials-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-of-

mixtures 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.

1.7.1 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. 1.22. To model the out-of-

plane 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-of-

mixtures, the stitch and foam contribution to the modulus is given by the relationship:

𝐸 = 𝐸 𝛼 + 𝐸 (1 − 𝛼) (1.5)

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-of-

plane 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:

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:

30

𝐸 = 𝑠𝑖𝑛 𝜃 𝐸 + 𝐸 1 − (1.11)

1.7.2 Shear modulus

An analytical shear model was also developed using a mechanics-of-materials-

based approach. A rule-of-mixtures was again used to account for the addition of stitches

shown in Fig. 1.23. 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:

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:

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)

1.7.3 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:

33

Table 1.6 Material and geometry properties used in modulus models

E G θ as bs As

Foam 101 Mpa 47.8 Mpa ----- ----- ----- -----

Stitch 10.1 GPa ----- varying 12.7 mm 12.7 mm 1.47 mm2 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, 90o stitch specimen test data, and

unstitched specimen data. The effective out-of-plane modulus for the 90o 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

90o, 12.7 mm, and 12.7 mm, respectively. Finally, solving for 𝐴 , 1.47 mm2 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 90o experimental data

point.

34

The needle diameter used for the 90o 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.

1.7.4 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. 1.24. The out-of-plane modulus model presented above used material

properties from the unstitched and the 90o, 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 900

stitched coupon had four complete stitches.

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. 1.25.

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 45o, but 35.2o

was found to be the mathematical maximum prediction. The results of the core shear test

do not indicate a clear maximum as the 30o and 45o 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 30o, and 45o (refer to section 4.3), may decrease the effective modulus of

the sandwich core under compression. Due to the alternating stitch pattern of the 30o, 45o

and 60o, under load, the core shear tests effectively apply tensile loads to only half the

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. 1.20. 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. 1.25 for the 30o and 60o panels

between the predicted and measured shear modulus values.

Due to the accuracy at 45o, it is hypothesized that these 45o stitches loaded in

compression contribute relatively more, as seen in Table 1.4. The 90o stitch specimens

were shown to increase the modulus by approximately 23% when compared to the

unstitched specimens, while the 60o specimens only had an increase in modulus of 2.7%

and the 30o and the 45o stitch specimens resulted in a decrease in modulus. In other

words, the 45o stitch reinforcement under compression found in a core shear test

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 90o 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 90o

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 90o stitching angle as well as a modest gain of 23% in

modulus. Surprisingly however, there was very little gain for the 30o and the 45o 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

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 45o stitch

specimens more accurately leading to a hypothesis that the 45o 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

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, 1385-1398. 4. 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 30 1445–1461. 7. 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 16-20. 9. Muthyala, V. D., 2007, “Composite Sandwich Structure with Grid Stiffened Core,” Master’s thesis, Osmania University, Hyderabad, India.

40

10. Mouritz, A. P., 2006, “Compression Properties of Z-Pinned Sandwich Composites,” Journal of Material Science, Melbourne, Australia, 41:5771-5774. 11. 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, 0-8031-4228-5. 15. American Society for Testing and Materials, “Standard Test Method for Flatwise Compressive Properties of Sandwich Cores,” West Conshohocken, PA, ASTM 365/C 365M-05, 0-8031-4228-5. 16. American Society for Testing and Materials, “Standard Test Method for Shear Properties of Sandwich Core Materials,” West Conshohocken, PA, ASTM C 273-00, 0-8031-4228-5. 17. Barbero, E. J., 1998, Introduction to Composite Materials Design, Taylor & Francis, Inc, Philadelphia, PA, pp. 5-85. 18. Swanson, S. R., 1997, Introduction to Design and Analysis with Advanced Composite Materials, Prentic Hall, Upper Saddle River, NJ. Chap 2-4

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 90o stitch angles decrease creep deformation by

approximately 75% when compared to unstitched sandwich specimens. Under shear

loading, stitch angles of 30o, 45o, and 60o 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

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-the-

thickness 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.

43

In this investigation, fabrication procedures were developed to introduce stitching

at 30o, 45o, 60o, and 90o 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, 30o, 45o,

60o, and 90o 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/m3,

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.

2.1.

2.3.1 90o 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

44

Table 2.1 Material selection

Configuration Stitching Method

Facesheet Material

Core Material Curing Method

Unstitched N/A

Woven carbon fabric

T300B 3K ([0/90]6T)

[12]

EPON 862 epoxy

EPON 9553 hardener

[13]

6 plies/facesheet

Polyurethane foam (FR-6710) (General Plastics)

[11]

160 kg/m3

12.7 mm thick

Vacuum Assisted

Resin Transfer

Mold (VARTM)

30o, 45o, 60o Hand Stitched 1600 denier

Kevlar 29 yarn

90o

Manual Industrial

Sewing Machine 1600 denier

Kevlar 29 yarn

Figure 2.1 Cross section of a VARTM infiltration assembly

45

pulled completely through to the bobbin side, as shown in Fig. 2.2. This allowed for

easier fabrication as the yarn intersection was made outside the core and the yarns kept

continuous through the core.

The 90o 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. 2.3. 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.

2.3.2 30o, 45o,60o 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 90o

stitched specimens. However, holes were premachined at 30o, 45o, and 60o 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.

46

Figure 2.2 Diagram of traditional and modified lock stitch.

Figure 2.3 Stitching of sandwich composites using an industrial sewing machine.

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 90o specimens. To stitch, the facesheets were set up as if they

were a cover for a book, as shown in Fig. 2.5. 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. 2.6. Approximately 20 needles were used to stitch larger panels row by

row. A step-by-step diagram is illustrated in Fig. 2.7. 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

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 30o, 45o, 60o sandwich specimens

49

Figure 2.7 30o, 45o, 60o 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

7-10. 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. 2.9. 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)

50

Figure 2.9 90o left, 60o right, stitched sandwich panel with stitch and resin column

diameter of 1.9 mm was selected to stitch the 90o specimens using an industrial sewing

machine [1]. The predrilled holes for the 30o, 45o, or 60o 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 90o stitch specimens and the 30o, 45o, and 60o

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 60o

panel with the foam removed can be seen in Fig. 2.10.

51

Table 2.2 Comparison of areal densities for varying stitch angles

Configuration Areal Density g/m^2 Normalized Areal Density

Unstitched 6740±82 1.00

90 Degree 7060±44 1.04

60 Degree 6920±75 1.02

45 Degree 6970±180 1.03

30 Degree 7300±170 1.07

Figure 2.10 Typical 60o stitch panel with foam removed

2.4 Flatwise Compression Creep Testing

2.4.1 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 C365-05 [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

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 82oC was deemed to be appropriate

for creep testing given the glass transition temperature (138oC) 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.

2.4.2 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. 2.11. The compression

spring, capable of a maximum load of 5.3 kN and load rate of 36 N/mm, uses a centrally

53

Figure 2.11 Compression creep test fixture diagram

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 0.025 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. 2.12.

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 www.mcmaster.com. Placement of the spring

between the specimen and the hinge resulted in a mechanical disadvantage of

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 ±45o 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.

56

2.4.3 Flatwise compression creep test results

Flatwise compression creep test results are presented in Fig. 2.13 and Table 2.3.

Of all sandwich configurations investigated, the 90o stitched specimens yielded the least

creep deformation, followed by the 60o , 45o, and 30o stitch orientations. The 90o stitches

were found to decrease total creep deformation by approximately 75% when compared to

the unstitched specimens.

Each configuration, except for the 30o 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 90o 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

2.5.1 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 273-00 [18]. Stress was calculated as shown in

Eq. 2.1:

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±0.2 6.0±0.3 4.7±0.1 1.5±0.4 3.7

30o 1.4±0.3 4.3±0.3 2.9±0.3 1.6±0.3 3.0

45o 1.2±0.2 3.3±0.4 2.1±0.2 1.5±0.2 1.7

60o 1.2±0.1 2.2±0.1 1.1±0.1 1.1±0.2 1.1

90o 1.0±0.2 2.0±0.2 1.0±0.1 1.1±0.2 0.9

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 273-00 [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 82oC 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

59

constant temperature of 82°C where time lapse images could be taken through a modified

door.

2.5.2 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. 2.14. 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

0.0025 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.

60

Figure 2.14 Core shear test fixture diagram

61

Figure 2.15 Time lapse frame of core shear test

2.5.3 Core shear creep test results

Core shear creep test results are presented in Fig. 2.16 and Table 2.4. It is noted

that one dial gage was defective for the 45o, 60o, and 90o tests, and thus, only five test

results are reported for these configurations. The strain versus time plots in Fig. 2.16

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 90o stitch

specimens resulted in the greatest creep deformation. However, the 30o, 45o, and 60o

stitch sandwich specimens all deformed similarly. This result was not expected, as the

45o stitch angle under shear loading would produce the maximum tensile and

compressive stresses on planes oriented at 45o. However, both the 45o and 60o stitch

orientation reduced creep by approximately 70% when compared to the unstitched

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±0.007 0.13±0.010 0.08±0.004 1.2

30 Degree 0.02±0.004 0.05±0.006 0.02±0.006 0.6

45 Degree 0.02±0.004 0.04±0.007 0.02±0.007 0.4

60 Degree 0.02±0.004 0.04±0.003 0.02±0.003 0.5

90 Degree 0.03±0.002 0.07±0.004 0.04±0.006 1.0

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 90o

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

90o stitch specimens recovered almost 50% of total displacement whereas the unstitched

specimens recovered only 25% of total creep displacement. The lower recovery

64

displayed by the 45o, followed by the 30o, 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 30o, 45o, and 60o 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 (45o). 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.

65

2.7 Conclusions and Recommendations

This investigation included the development of stitching and fabrication methods

to produce 30o, 45o, 60o, and 90o angled stitch orientations in foam core sandwich panels.

Stitching at 90o was found to greatly reduce creep under flatwise compression creep

testing by approximately 75% when compared to unstitched specimens. The 30o, 45o,

and 60o 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.

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 short-

term 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.

67

2.8 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 D2990-09. 3. 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, 637-647. 5. 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, 1437-1457(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, 1385-1398. 10. 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, 41:5771-5774. 12. 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, 57. 14. Synthetic Carbon Fiber Fabric I.O.P. Woven, Class 70, BGF Industries, Inc., Los Angeles CA.

68

15. 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, 0-8031-4228-5. 18. American Society for Testing and Materials, “Standard Test Method for Shear Properties of Sandwich Core Materials,” West Conshohocken, PA, ASTM C 273-00, 0-8031-4228-5. 19. Yourd, R. A., 1996, “Compression Creep and Long-Term Dimensional Stability in Appliance Rigid Foam,” Journal of Cellular Plastics, Pittsburgh, PA, 32. 20. 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, 0-486-66016-8. 22. Xu, Y., 2009, “Creep Behavior of Natural Fiber Reinforced Polymer Composites,” Ph. D. thesis, Louisiana State University.

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

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. 3.1.

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 3.1 60o stitch reinforcement in polymer core sandwich structure

71

3.3 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

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.

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3.4 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. 3.2. and Fig. 3.3. 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. 3.3. 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

74

Figure 3.2 Compression creep test fixture diagram

75

Figure 3.3 Flatwise compression creep test fixture

Table 3.1 Test fixture load summaries

Flatwise Compression Creep

Test Fixture Through-Thickness Shear Test

Fixture

Maximum load capacity of coil spring

5.3 kN 7.7 kN

Mechanical disadvantage produced by test fixture

0.81 0.78

Maximum specimen force provided by coil spring

4.5 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%

76

were parallel for a given thickness of sandwich specimen. Dial indicators with a 0.025

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)

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 0.2o 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.

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 through-

thickness 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 through-

thickness shear testing, as shown in Fig. 3.6 and 3.7. The spacing between the box beams

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Figure 3.6 Through-thickness shear test fixture diagram

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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

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during the test, dial gages with 0.0025 mm resolution were placed on magnetic bases to

measure the relative axial motion occurring between the opposing faces of the sandwich

composite specimen. The fixture springs were preloaded on threaded rods, heated to the

test temperature, and held for at least 24 hours prior to testing. From preliminary testing,

an additional 440 N of load was needed to be placed on the spring to account for fixture

compliance at the start of testing.

To start the test, one nut was again loosened, transferring load from the threaded

rod to the box beams. To monitor the dial gages, the same high resolution digital camera

and intervalometer were used to obtain time lapse images every hour. A total of six creep

fixtures were placed in a large Blue-M laboratory convection oven with a modified door

for viewing during testing.

3.6 Creep Testing of Composite Sandwich Specimens

The flatwise compression and through-thickness shear test fixtures described

above were used to evaluate the use of Kevlar stitching to reduce creep deformations in

polymer foam core sandwich composites. A series of tests were performed using

specimens fabricated with stitches at different orientations as well as unstitched

specimens. All testing was performed at an elevated temperature of 82oC for a total of

150 hours. A total of six specimens were tested simultaneously for each sandwich

configuration.

Typical test results, presented for the case of 60o stitch reinforcement (as shown

previously in Fig. 3.1) are shown in Fig. 3.8 and Fig 3.9 for flatwise compression and

through-thickness shear, respectively. For both loading methods, specimens exhibited an

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Figure 3.8 Strain versus time results for flatwise compression creep testing

Figure 3.9 Strain versus time results from through-thickness shear creep testing

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initial, linear strain versus time response, followed by a secondary region characterized

by a lower slope as expected. After testing, specimens were removed from the fixtures

and the recovery of creep deformation was measured for flatwise compression tests

following several months. Real-time recovery was not measured due to difficulties in

quickly removing all loads from the test fixtures and monitoring their displacements with

time. Core shear recovery was not measured as the loading plates were removed for use

in testing more sandwich configurations.

The load levels used in these tests were selected to produce a measureable amount

of creep in the most creep-resistant stitch configuration for each type of loading. The

resulting levels of creep deformation would generally be considered high for actual

applications. The strains resulting from creep reached nearly 50% normal strain for

flatwise compression loading 12% shear strain for through-thickness shear loading.

Lower stresses are recommended for future tests as it would also resemble more realistic

applications where creep deformations may be a concern.

3.7 Summary

Portable and compact, spring-loaded creep test fixtures were developed for testing

polymer foam core sandwich composite specimens under flatwise compression and

through-thickness 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

inherent load loss due to creep deformation was limited to 1.6% for flatwise compression

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loading and 1.0% for through-thickness shear loading. Successful use of the test fixtures

was demonstrated through creep testing of polymer foam core sandwich composites

reinforced with Kevlar stitching.

3.8 References

1. ASM International, 2003, Characterization and Failure Analysis of Plastics, Library of Congress Catalog-in-Publication Data. 2. Wineman, A. S., 2000, Mechanical Response of Polymers, Cambridge University Press, Cambridge, UK. 3. Scott, David W., Lai, James S. and Zureick, Abdul-Hamid, 1995, “Creep Behavior of Fiber-Reinforced Polymeric Composites: A Review of the Technical Literature,” Journal of Reinforced Plastics and Composites. 4. Muller, Werner, et al., 2007, “Long-term shear strength of geosynthetic clay liners,” Geotextiles and Geomembranes, Berlin, 26. 5. 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 D2990-09. 6. American Society for Testing and Materials, 2002, “Standard Test Method for Creep of Concrete in Compression,” West Conshohocken, PA, ASTM C512-02. 7. A.S.M. and Awal, Abdul, 2002, “A Study on Shrinkage and Creep of Prepacked Aggregate Concrete,” Advances in Mechanics of Structures and Materials, Bangladesh Agricultural University, Bangladesh, India. 8. Henshaw, John M. and Pierson, Ben, 2005, “Annual Progress Report, Automotive Lightweight Materials,Low-Cost Test Methods for Advanced Automotive Composite Materials; Creep Compression Fixture,” US Department of Energy, pp. 243-247. 9. Lupkes, K. R. and Tuttle, M. E., 1995, “Mass Loss and Residual Stiffness and Strength of IM7/5260 Composites Subjected to Elevated Loads and Temperatures,” University of Washington, Department of Mechanical Engineering, Seattle, WA. 10. Fan, Yu-Bo, et al., 2008, “Effects of Mechanical Load on the Degradation of Poly (D, L-Lactic acid) Foam,” Polymer Degradation and Stability, Beijing, China.

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11. American Society for Testing and Materials, “Standard Test Method for Flatwise Compressive Properties of Sandwich Cores,” West Conshohocken, PA, ASTM 365/C 365M-05, 0-8031-4228-5 12. American Society for Testing and Materials, “Standard Test Method for Shear Properties of Sandwich Core Materials,” West Conshohocken, PA, ASTM C 273-00, 0-8031-4228-5.