Influence of stitching on the out-of-plane behavior of composite materials

Abstract

Fiber-reinforced polymer composites are widely used in the aerospace industry due to their high stiffness and strength-to-weight ratios. However, their applicability can be limited by their relatively low interlaminar properties when compared to metallic alternatives. Through-thickness reinforcement approaches, such as stitching, z-pinning, needling, tufting, and three-dimensional weaving, have been developed in recent decades to enhance the interlaminar properties of composites. Stitching is considered to be an efficient and cost-effective method to reinforce composites in the through-thickness direction. Additionally, stitch parameters (stitch density, linear thread density, thread material, pretension, etc.) highly influence the in-plane and out-of-plane properties. This paper summarizes results from over one hundred papers on the influence of stitch parameters on fracture energy, interlaminar strength, and impact characteristics of stitched composite laminates, sandwich composites, and high-temperature composites. Much of the research on the influence of stitch parameters has focused on thermoset polymer matrix composites (PMCs), while fewer studies have investigated the impact of stitch parameters on high temperature or sandwich composites. Modification of existing and new test methods have been developed to adequately measure the effectiveness of stitching on the out-of-plane behavior of PMC panels. Results demonstrate that out-of-plane properties of PMCs are highly dependent on stitch parameters and can be enhanced by through-thickness stitching.

Keywords

Stitching through-thickness reinforcement fracture stitched composite laminates sandwich composites

Introduction

Lightweight materials are essential in the aerospace industry to reduce fuel consumption and emissions, and to increase aircraft range and payload. In particular, fiber-reinforced composite materials are extensively used due to their tailorability and their high in-plane specific strengths and stiffnesses. However, their applicability can be limited by their low interlaminar shear and interlaminar tensile strengths; this is due to a property mismatch between the composite laminae¹ and the polymer matrix being the primary load carrier for out-of-plane loads. For example, Nie et al.² reported the in-plane tensile and interlaminar shear strengths for a carbon/silicon carbide composite are 234.3 MPa and 29.1 MPa, respectively. Increasing the property mismatch, as a result of increasing the angle between subsequent plies, was shown to lower interlaminar shear strength and delamination resistance under out-of-plane loading.³ These low interlaminar properties in layered composites can result in delaminations, which can be due to impacts or other excessive interlaminar loads during part assembly or service operations. These composite parts typically undergo costly nondestructive evaluations and, depending on the delamination size and location, may need to be replaced.

Novel approaches have been developed to reinforce composite materials in the through-thickness direction to prevent delamination. These methods include z-pinning,^4–7 needling,⁸ three-dimensional (3D) weaving,⁵ stitching,^5,9,10 and tufting. Z-pinning is typically performed by embedding small diameter carbon rods in a polymeric foam preform. An ultrasonic horn is used to drive the pins through the thickness of an uncured prepreg laminate prior to cure.¹¹ The needling process uses downward-barbed needles to re-orient in-plane short fiber mats in the through-thickness direction of dry composite preforms.⁸ Weaving, stitching, and tufting are traditionally performed using an industrial sewing machine or loom to introduce through-thickness reinforcement before resin infusion. Tufting is a form of stitching that uses non-interconnected stitches to reinforce polymer composites in the through-thickness direction. Overall, stitching is a simple 3D reinforcement method that provides similar interlaminar improvements to 3D weaving^5,9 and can be used for any conventional dry composite preform that is commercially available. This capability allows greater design flexibility in the composite layup, making stitching of great interest for improving the performance of composite aerospace structures.

The stitching process involves sewing aramid, carbon, polyester, glass threads, or yarns, into a non-crimped fabric (NCF) or woven dry preform using an industrial or robotic sewing machine. A thread indicates a twisted assemblage of tows, where a tow is an untwisted collection of fiber filaments. Threads can be twisted together to provide enhanced tensile thread strength. NCFs are fabrics that consist of multiple layers of parallel tows that are held together using non-structural (low filament count) threads, typically using a polyester material. Stitching can be performed using prepregs prior to cure, but damage such as in-plane fiber rupture and stitch needle breakage can occur during the stitching process.⁹ The formation of the stitch during the sewing process is referred to as the stitching style. The most common styles of stitches are the modified lock stitch, lock stitch, and chain stitch, as shown in Figure 1(a) to (c), respectively. The lock stitch is an interlock of stitching thread at the mid-plane of the composite laminate and has been shown to create stress-concentrations and reduce in-plane properties.^1,13 Stress concentrations, due to fiber waviness and resin-rich pockets, were reduced by using a modified lock stitch, for which the thread interlock is on the surface of the preform. Both the lock and modified lock stitches require a needle thread and a bobbin thread, which are on opposite sides of the laminate. The bobbin thread is used to interconnect adjacent stitching on both sides of the laminate to enhance the out-of-plane behavior as compared to a non-interconnected stitching style such as tufting. For the chain stitch, a single thread is used to stitch fabric from one side of the composite laminate. Stitches are periodically spaced and characterized by the stitch pitch (P) and spacing (S), as shown in Figure 1(d). The stitch pitch (P) is defined as the distance between two adjacent stitches along the same stitch seam and the stitch spacing (S) is the distance between two adjacent seams of stitching.

Stitch processing parameters have considerable influence on the mechanical performance in both the in-plane and out-of-plane directions. These processing parameters include the thread material, linear thread density (mass per unit length of the thread), thread finish, stitch density or the number of stitches per unit area (1/(P·S)), stitching distribution or pattern, stitch style, and stitch pretension.¹ Various configurations of these stitch parameters have produced contradictory results regarding their in-plane mechanical behavior. Generally, stitching composites results in approximately a 10% reduction in Young’s modulus and tensile strength.^9,10 The degradation of in-plane properties may not be critical for components that are primarily subjected to interlaminar stresses, where low in-plane stresses result in high margins of safety. However, the reduction of in-plane properties (bulk stiffness and tensile strength) must be considered for stitched structural components with high in-plane stresses. Therefore, careful attention to stitch parameters is needed for an effective design of 3D reinforced composite structures.

The Pultruded Rod Stitched Efficient Unitized Structure (PRSEUS) concept,¹⁴ developed by Boeing, NASA, and the United States Air Force, uses a novel selective stitching approach as a means of joining major structural elements (skin, stringer, and frames) prior to resin infusion and subsequent cure. The reduction of in-plane properties is minimized by stitching structural elements where the delamination due to interlaminar peel stresses at the joined regions are the most critical.¹⁵ Delamination due to interlaminar peel stresses are the primary form of failure in overlapping joints, and stitching has shown to increase the lap joint strength by approximately 60% to 175%.^15–23 Quasi-static structural testing of notched PRSEUS panels has also shown that the selective stitching approach allows unitized structural members to effectively arrest delaminations from barely visible impact damage and translaminar crack growth.¹⁴ As a result of this research, a significant amount of exploration of the out-of-plane behavior of stitched polymer composites was performed in 2007^{15,18,19,24–35} and 2008.^2,20,36–47 Few stitched composite studies are found between 2000-2004, as shown in Figure 2. Prior to 2000, NASA enacted the Affordable Composite Technology (ACT) program to develop a database of composite technology for implementation in production aircraft. A considerable amount of research on stitched composite structures was performed during this time and is summarized in reference.⁴⁸ The in-plane mechanical properties of stitched polymer matrix composites (PMCs) with respect to stitch processing parameters have been well characterized and reviewed.^1,9,10 However, the fracture behavior and associated interlaminar properties have not received as much attention and are not well understood for a variety of composite material systems and stitch parameters. For example, ceramic matrix composites (CMCs) have relatively low in-plane mechanical properties due to their manufacturing processes,⁴⁹ but their interlaminar and flexural strengths have been shown to increase by approximately 20% by adding through-thickness stitching compared to an unstitched structure.² In this study, a review of the fracture behavior (mode I, mode II, and mixed-mode) of stitched composites, and the necessary experiments needed for their characterization are presented. Additionally, a summary of the latest advancements in stitched composites with a focus on the out-of-plane behavior (interlaminar strength and impact) and for select material systems (PMCs, CMCs, carbon/carbon composites and sandwich composites) is provided.

Stitching parameters for fracture characterization

Studies focused on different stitch parameters to evaluate the fracture and interlaminar properties of stitched polymer matrix composites are summarized in Table 1. The percentage of papers in each category was estimated based on the total number of unique papers. In many of these studies, linear thread density (40.6%) and stitch density (56.3%) have been investigated because these two stitch parameters primarily influence the out-of-plane behavior of layered composites. As such, the influence of stitch density and linear thread density on the out-of-plane behavior are the primary stitch parameters discussed in this review. Studies that used various thread and stitch densities are shown in Table 2. This listing includes papers that did not investigate stitch parameters, but evaluated a single set of stitch parameters on the out-of-plane response. Additionally, the ranges represent a standard deviation of stitch density and linear thread density from all referenced papers that evaluated the effectiveness of stitches on the out-of-plane behavior of PMCs. Moderate levels of stitch density (0.0025 stitches/mm² to 0.05 stitches/mm²) have been investigated in a majority (85.4%) of studies. Additionally, most of these studies (90%) investigated linear thread densities that range between 0 denier to 1800 denier. Very few articles report the influence of other stitching parameters such as pretension, stitching style, and thread twist. Pretension has shown to significantly influence out-of-plane properties through finite element modeling⁷² and analytical approaches.⁶² Experimentally quantifying the amount of thread pretension within the cured part is difficult and validation of these modeling approaches requires more attention. Thread tension can be measured during the stitching process,^85,86 but thread relaxation can occur from compaction of the dry preform during the vacuum-assisted resin transfer molding process, which subsequently reduces thread pretension. Several studies have also investigated different stitching thread materials, such as Kevlar™,^19,50,51,74 glass,¹⁹ carbon fibers,^19,50,51,74 Vectran™,⁷⁴ polyamide,⁷⁵ and phenoxy.⁷⁵ Lastly, researchers have also used analytical⁷² and finite element modeling⁷⁶ approaches to investigate the influence of stitch thread stiffness on the delamination resistance of polymer matrix composites.

Table 1.

Studies on that evaluated influence of stitching parameters on the out-of-plane composite behavior.

Stitch parameter	References	Percent of papers (%)
Linear thread density	^19,50–61	40.6
Stitch density	^{2,50–52,55–58,61–70}	56.3
Stitch pattern	^{28,29,45,58,62,71}	15.6
Pretension	^62,72	6.3
Stitch style	^19,73	6.3
Twist	¹⁹	3.1
Ply orientation	¹⁹	3.1
Stitch material	^{19,50,51,72,74–76}	21.9

Table 2.

Studies on stitch and linear thread densities.

Stitch parameter	Range	References	Percent of papers (%)
Stitch density (stitches/mm²)	0.0025–0.050	^{2,19,25,28,29,40,41,45,50–58,60–70,74,75,77–81}	85.4
	0.051–0.0950	^{36,50–52,58,73,77,82}	19.5
	0.0951–0.140	^{50–52,55,56,58,61,64,69,70,74,76,81,83,84}	36.6
	>0.140	^2,77	4.9
Linear thread density (denier)	0–600	^{19,25,40,41,45,50,51,56,59,61,64,69,70,74,75,79,80}	56.7
	601–1200	^{50,51,53–55,58,59,73,74}	30
	1201–1800	^{28,29,53,55,58,62,65,79,83}	30
	>1800	^55,60,66,67	13.3

Mode I fracture energy

Mode I fracture energy, or the energy required to promote an opening mode of crack growth, can be increased by incorporating through-the-thickness stitching.^{19,41,53,72,76} Under mode I conditions, stitches bridge the opposing crack faces and induce internal traction stresses to resist crack growth. Researchers have noted that stitching does not influence the initial fracture energy when the delamination front is approximately 10 mm from the initial stitch row.⁶³ The fracture energy during propagation and near a stitch seam results in a significantly greater mode I fracture energy than at crack initiation. The relative influence of stitch density on the normalized mode I fracture energy for select linear thread densities and stitching thread materials from published data is shown in Figure 3. The maximum fracture energy of a stitched composite laminate is normalized by the fracture energy of its unstitched counterpart. Low stitch densities correspond to relatively large distances between adjacent stitches, whereas high stitch densities correspond to small distances between adjacent stitches. Overall, increasing the stitch density linearly increases the mode I fracture energy for each material due to higher traction loads from stitch bridging during delamination propagation. During failure of the stitch under mode I delamination, multiple stitch rows have been observed to fail during double-cantilevered beam (DCB) testing.⁵³ Additionally, other studies^28,29 have found that the relative pattern or distribution, while maintaining a constant stitch density, influences the mode I fracture energy. Stitching has also been shown to provide the same delamination resistance for any in-plane fiber orientation.⁴¹The normalized fracture energy can be further improved by increasing the linear thread density of the stitching thread. The influence of linear thread density on the normalized steady-state mode I fracture energy for select stitch densities and stitching thread materials is shown in Figure 4. Overall, increasing the linear thread density proportionally increases the normalized fracture energy. Additionally, the increase in fracture energy appears to be highly dependent on the stitching thread material. For an E-glass thread material (0.04 stitches/mm²), increasing the linear thread density from 600 denier to 1200 denier increases the estimated fracture energy by approximately 174%. A 400 denier carbon thread shows a more significant increase (∼900%) in the normalized fracture energy for stitched densities greater than 0.04/mm². These studies may indicate that there is a stitch parameter interaction between linear thread density, stitch density, and stitch material on the mode I fracture energy.

Increasing the linear thread density can result in a reduction in the in-plane mechanical properties due to increased fiber waviness and the formation of resin pockets near the stitch.⁹ For example, Heß et al.³⁰ reported a 10% to 14% reduction in the in-plane properties for carbon/epoxy laminates using 612 denier and 1224 Denier E-glass stitching. Also, untwisted carbon fiber threads within woven carbon fabric have shown to increase mode I interlaminar fracture energy without impacting in-plane properties.¹⁹ A uniform distribution of untwisted filaments within the displaced region of in-plane fibers is developed, thereby decreasing the resin-rich pockets near the stitching regions. Heb et al.¹⁹ reported that the steady-state mode I fracture energy is primarily controlled by the thread diameter or linear thread density. However, this finding is somewhat contrary to other reported data, as shown in Figures 3 and 4. The mode I fracture energy also appears to be highly dependent on thread stiffness and stitch tensile strength.^72,76 For example, mode I fracture energy is increased by a factor of 2 and 15 when using 612 denier E-glass thread and 756 denier carbon thread, respectively. These results agree with several computational studies that have focused on the influence of thread stiffness, failure load, and crack length.^72,76 In particular, Glaessgen et al.⁷⁶ used a virtual crack closure technique to investigate the influence of thread stiffness on the mode I strain energy release rate for double cantilever beam (DCB) specimens. Increasing the thread stiffness caused an increase in the stitch failure force.⁷⁶

Mode II fracture energy

Under mode II fracture conditions, delaminations develop due to internal shear stresses between the composite laminae and result in a sliding action between two opposing crack faces. During a mode II delamination, the stitches resist the crack front through a “plowing” action, in which the stitches deform the surrounding matrix near the delaminated interface. The additional energy expended to resist crack growth is primarily due to the deformation of the surrounding matrix and not the failure of the through-thickness stitching.⁸⁷ This plowing action is a result of a snubbing phenomenon, first identified by Cox⁸⁸ and Cartié et al.⁸⁹ Snubbing refers to the significantly large and non-uniform shear stresses at the delaminated interface when a bridged through-thickness reinforcement laterally deflects under mode II conditions. The influence of the snubbing effects can also be induced under mode I conditions when specimens are subjected to large displacements. After deformation of the matrix due to plowing, the through-thickness reinforcement fails primarily due to shear plasticity, internal splitting, and frictional pullout.⁸⁹

Similar to mode I conditions, increasing the stitch density can increase the normalized mode II fracture energy. The influence of stitch density on the normalized mode II fracture energy for select linear thread densities and stitching thread materials is shown in Figure 5. Increasing the stitch density increases the mode II fracture energy by up to 330% when compared to its unstitched composite counterpart for carbon and Kevlar stitching materials.⁵¹ Unlike mode I fracture behavior, neither the thread material nor thread strength seems to have a significant impact on mode II fracture energy.⁵³ For example, 675 denier polyester thread at a low stitch density (0.02 stitches/mm²) has the same relative performance as that of 756 denier carbon thread with a greater stitch density (0.08 stitches/mm²). This may be greatly influenced by the relative deformation of the matrix as stated in the first paragraph. Additionally, Jain et al.⁵¹ noted that stitching does not significantly influence mode II fracture energy at crack initiation. Wood et al.²⁹ reported that the number of stitches along the crack front initially improved the mode II fracture energy, but this improvement was not apparent for long crack growth with significant stitch bridging zone lengths. Furthermore, stitching in carbon/epoxy laminates also reduces unstable crack propagation that is normally associated with end-notch flexure (ENF) testing.⁵¹ Prior to failure, the crack front wraps around the stitch as the stitches bridge the crack plane.²⁹ The influence of linear thread density on the steady-state mode II fracture energy for select stitch densities is shown in Figure 6. Increasing the linear thread density of Kevlar thread increases the fracture energy of PMCs, whereas the polyester thread does not improve the fracture energy for moderate (675 denier) to large (1350 denier) linear thread densities. Additionally, relatively high strength and stiffness stitching threads, such as carbon thread, do not appear to significantly increase the mode II fracture energy as compared to a polyester stitching thread. Further increases in the mode II fracture energy may be attained by increasing the stitch pre-tension. Jain et al.⁵¹ analytically showed that the mode II fracture energy increases at a greater rate due to larger surface tractions near the crack tip generated by greater stitch pretensions. The steady-state fracture energy can also be improved by altering the through-thickness orientation of the stitches. Stitches that are diagonally oriented against the direction of crack growth, as shown in Figure 7, have been shown to have twice the mode II fracture energy as compared to stitches oriented with the direction of crack growth.⁵³ This behavior is primarily attributed to an increase in the shear stiffness along the delamination plane.⁵³ Stitches that are oriented with the direction of crack growth result in lower fracture energy due to stitch frictional sliding, but greater sustained load at the delaminated interface. Stitches that are oriented against the direction of crack growth will fail due to high shear stresses that result in unstable crack growth, but greater crack-growth resistance.

Mixed-mode fracture energy

In many practical stiuations, structural components are subjected to combined loads during service operation. As a result, the development of delamination is due to a combination of tensile and shear loading near the crack front. The normalized mode I and mode II fracture energy with respect to the modal ratio is shown in Figure 8.⁹¹

The modal ratio is the mode I fracture energy divided by the mode II fracture energy. A 0% and 100% modal ratio corresponds to a pure mode II and mode I fracture energy, respectively. For relatively low modal ratios (<30%), Trabelsi et al.⁹⁰ experimentally observed that stitch failure does not occur in a woven carbon/epoxy composite material system. Increasing modal ratio from 30% to 70% increases the normalized mode I and mode II fracture energys by a factor of 4.0. The increase in the overall fracture energy under mixed-mode conditions has also been observed by other researchers.⁹¹ Further increases in the modal ratio results in a decrease in the mode II fracture energy as the modal ratio approaches 100% (pure mode I fracture).

Interlaminar strength

The primary test method to evaluate the interlaminar shear strength of stitched CMCs is the compressive interlaminar shear test per ASTM standard C1292-12.⁹² Figure 9 shows the influence of stitch density (0.01–0.16 stitches/mm²) on the normalized interlaminar shear strength of CMCs for two selected sizes (1K and 3K) of carbon fiber. Increasing the stitch density is shown to increase interlaminar shear strength by approximately 10%–30% using 1K carbon fiber tows as stitching thread materials in CMCs. Similar increases in the interlaminar shear strength has also been observed for polymer matrix composites.¹⁹ High strength carbon fiber stitching tows (T300 3K) resulted in lower interlaminar shear strength as compared to T300 1K carbon fiber tows. This behavior is due to the ineffectiveness of the chemical vapor infiltration process to fill the displaced volume of the in-plane fibers near large carbon fiber tows.⁷⁷ Therefore, large voids developed near the stitches. The primary failure modes observed in stitched CMCs were delamination, stitch fiber filament failure, and secondary ceramic matrix microcracks that were deflected by the nearby stitch filaments.² The stitching fibers completely ruptured after the silicon-carbide delaminated during the interlaminar shear test, unlike traditional PMCs where a plowing action can occur for ductile matrices.² Nie et al.² also noted that stitching improved the in-plane tensile strength by approximately 27% with no influence on the flexural strength for high stitch densities (SD ≥0.04 stitches/mm²).

Measurements of the interlaminar tensile strength of stitched composites are not presently available in the literature. This is typically because the effectiveness of the through-thickness reinforcement is primarily observed after the stitch has bridged the crack. It is difficult to estimate the tensile strength due to a significant reduction in area after the bulk material has delaminated. However, the interlaminar or flatwise tensile test has been used to characterize the load-displacement response (or traction-separation response) after delamination has occurred. A more detailed discussion is provided in the test methods and analysis portion of this manuscript.

Impact behavior

During low-velocity impact events, delamination is initiated by the development of translaminar microcracks. These microcracks initially develop due to a property mismatch between the matrix and reinforcement material; however, they can also occur due to particle inclusions, resin-rich regions, or residual stresses. The translaminar microcracks grow to nearby plies and are halted by nearby plies of different orientation as the impact energy approaches the threshold energy at which delamination occurs.⁹³ Above the threshold energy, large out-of-plane normal stresses (through the thickness) result in mode I delamination between adjacent plies as a result of nearby microcracks.⁹⁴ Further away from the origin of impact, the laminate develops high interlaminar shear stresses due to local bending. This deformation results in the formation of mode II delaminations. Large delamination zones can occur during impact and are typically not visible at the surface.

Due to the inherent discretization of the stitching process, the initiation of delamination from microcracks due to impact is not halted by the presence of stitching. However, delamination growth can be significantly arrested or minimized by the through-thickness reinforcement, resulting in a reduction in the delamination area.^40,61,69,70 Researchers have also reported that the improved impact damage resistance was only observed for quasi-isotropic laminates with a thickness greater than approximately 1.9 mm.⁸⁴ The same results were not obtained when using a cross-ply configuration of a similar thickness.²⁵ The normalized delamination area with respect to impact energy for different linear thread densities, stitch densities, and layup configuration is shown in Figure 10. Overall, the normalized damage area is relatively independent of the impact energy for select stitch densities and linear thread densities. Very little differences are observed in the damaged area between a cross-ply and a quasi-isotropic laminate at relatively low stitch densities (<0.04 stitches/mm²). This similarity suggests that damage associated with an increase in the interlaminar stresses between laminae, due to a greater property mismatch in the cross-ply laminate, is relatively contained between adjacent stitching regions. From a damage tolerance perspective, this characteristic may increase the tailorability of composite designs that were otherwise unachievable without through-thickness reinforcements.

Stitch density is the primary stitch parameter that arrests and delays delamination for quasi-isotropic (thickness = 4, 6 mm) and cross-ply (thickness = 2 mm) laminates with a carbon/epoxy material system, as shown in Figure 10. The impactors used in these studies were steel hemispherical tips with diameters ranging from 12.7 mm to 15.9 mm (mass = 2.28 kg to 6.26 kg).^25,70 Increasing the stitch density from 0.028 stitches/mm² to 0.111 stitches/mm² decreases the normalized delamination area from 93% to 53% when compared to its unstitched counterpart using Vectran thread with a carbon/epoxy laminate. This decrease in the delamination area did not result from increasing the linear thread density of the stitching thread, but rather increasing the linear thread density from 200 denier to 400 denier decreased the damaged area by approximately 10% for relatively low stitch densities (≤0.028 stitches/mm²). Stitched laminates subjected to impacts showed no significant difference with respect to peak force and absorbed energy as compared to unstitched laminates.²⁵ Greater reductions in the delamination area may be achieved by increasing the linear thread density or other stitch parameters such as thread stiffness or thread pretension. Additionally, there were no observed differences in the damaged area for high stitch densities (≥0.111 stitches/mm²) for a quasi-isotropic carbon/epoxy material system with 200–400 denier Vectran threads. X-ray CT images of the damage distributions in the thickness direction for unstitched and stitched composite laminates are shown in Figure 11. The damage grew in a conical fashion from the point of impact, where the greatest amount of delamination was observed near the opposing face of the composite laminate.⁴⁰ Damage appeared to be more uniform in the thickness direction for the stitched composite laminates where additional delamination and microcracking near the point of impact were observed. The occurrence of additional delamination and microcracking is attributed to the stress-concentrations that occurred at the resin-rich areas located around the through-the-thickness reinforcements.^61,70 Increasing the linear thread density (> 400 Denier) increases the fiber waviness near the stitched zones and thus may accentuate additional microcracking and delamination.^64,95 Tan et al.⁶⁴ performed quasi-static indentation tests and found that the incipient damage load for laminates with high linear thread densities is lower than in their unstitched counterpart; this difference was primarily due to matrix cracks near the stitching loops.^95,96 The formation of microcracks also reduces the fatigue life,^94,97 increases the development of delaminations associated with impact,^60,66,98 and increases gas permeability^39,99–102 during service operations.

Significant attention has also been given to reinforcing sandwich composites with stitching to enhance the core-to-facesheet separation resistance under low-velocity impact. Internal cores such as polymeric foams have been primarily used since the through-thickness stitching produces resin pathways that form stitch-resin columns during the resin transfer molding process. Stitching increases the weight by approximately 1%.⁶⁰ Furthermore, stitches that are oriented at 45° have been shown to enhance the in-plane flexural rigidity, in-plane shear, and out-of-plane compressive strength.^66,98 Under impact, traditional damage modes^103,104 of foam core sandwich composites subjected to impact appear to be absent when the sandwich composite is reinforced with through-thickness stitching.⁶⁷ Specifically, the bottom surface of the impacted sandwich composite does not develop delaminations as typically observed in sandwich composites without through-thickness reinforcement. In stitched sandwich composites subjected to low-velocity impact, the primary form of failure is stitch-matrix column buckling and delamination of the top-most surface. A high density of stitches has also been shown to stiffen and strengthen composite structures, allowing for greater energy absorption and facesheet delamination suppression during the impact of orthogonally-stitched⁶⁵ and obliquely-stitched⁶⁷ sandwich composites. However, incipient failure during impact may occur earlier in the stitched component due to the weak interface between the stitching and facesheets.^67,98 Stitched regions subjected to impact have also been reported to undergo larger regions of core cracking than unstitched panels.⁶⁷

Test methods and analysis to characterize stitched composites

Test methods to evaluate the fracture, interlaminar, and impact characteristics of PMCs, sandwich composites, and high-temperature composites (carbon/carbon composites and CMCs) are shown in Table 3. Generally, PMCs have been used to evaluate the influence of stitching on their interlaminar and fracture behavior. Few studies have experimentally investigated the influence of stitching using sandwich composite and high-temperature composite material systems. The DCB,¹⁰⁵ ENF,¹¹⁴ and low-velocity impact test methods¹³² have been the principal experimental methods to evaluate the effectiveness of stitching in a PMC. The DCB and ENF test methods are used to estimate the mode I and mode II fracture energy, respectively, whereas low-velocity impact test methods are used to estimate the energy absorption characteristics.

Table 3.

Summary of interlaminar strength and fracture test methods.

Test type	Property	Stitched			Unstitched
Test type	Property	Polymer composite laminate	Sandwich composite	High-temperature composite	High-temperature composites
Double cantilevered beam¹⁰⁵	Mode I fracture energy	^{6,19,28,41,50,53–58,63,71,73–76,78,83,106}	^107–109	–	^110–113
End-notch flexure¹¹⁴	Mode II fracture energy	^{29,51,53,55,57,72,75,82,106}	–	–	¹¹⁰
Mixed-mode fracture ¹¹⁵	Mixed-mode fracture energy	^90,91,106	–	–	–
Flatwise or interlaminar tensile^{6,59,116–118}	Traction-separation response	^{6,54,56,59,63,74}	⁶⁵	–	¹¹²
Short beam shear¹¹⁹	Interlaminar shear strength	^19,45,75	–	–	–
Compact tension, single-edge-notch tension¹²⁰	Fatigue crack growth and translaminar fracture toughness	⁷⁹	–	–	^121–123
Interlaminar shear^124,125	Traction-separation response, interlaminar shear strength	^–	–	^2,77	¹²⁶
Low velocity impact¹²⁷	Energy absorption	^{25,40,61,69,70,80,84}	^60,67,68,128	–	^129–131

Estimating the fracture energy of stitched composites has resulted in the development of new test methods^83,106 and the modification of existing standards.¹⁰⁷ Current DCB test standards use an embedded insert at the specimen midplane to act as a delamination or debond site. Loading blocks or piano hinges are used to apply an external load above and below the debond region, resulting in flexural bending loads that promote crack growth. The propagation of the debond occurs when the internal tensile stresses at the crack front exceed the interfacial strength between the plies. As the delamination front approaches a stitched seam, the internal tensile stresses must exceed the tensile strength of the stitched seam before further crack propagation can occur. As a result, a high-rotational constraint is developed at the crack front due to the through-thickness reinforcement. This constraint subjects the delaminated arms to greater bending stresses and induces flexural failure. Guenon et al.¹³³ also observed that crack growth can directionally deviate, or branch to adjacent plies, instead of propagating in a self-similar manner. Significant stitch bridging can also occur during fracture, which may invalidate the small plastic zone assumption within linear elastic fracture mechanics that is used to develop these test methods. Therefore, current test standards are not considered suitable to evaluate stitch composite laminates. The influence of crack length relative to the stitch location has also not been investigated, although a significant amount of fracture data is available for stitched PMCs. Alternatives to the DCB, ENF, and mixed-mode bending (MMB) fracture tests have been developed to address the high rotational constraint by inducing additional tension to delay failure.^83,106 Other researchers have also reinforced the DCB and ENF fracture specimens with aluminum, steel, or composite doubler plates to prevent failure of the delaminated arms.^28,29,50,83 The thickness of the doublers used in these studies range from 2 mm to 10 mm in thickness. Lastly, adding doublers to the specimens has shown to develop stable crack growth as compared to specimens without doublers.⁹⁰ Further research on the sizing of the doublers for select stitch densities and linear thread densities needs to be performed.

The addition of the doubler plates for the DCB or ENF test specimens alters flexural rigidity and the location of the neutral axis of the delaminated arms that are subjected to bending. Modified beam theory may not provide an accurate estimate of the fracture energy due to the homogeneous specimen stiffness assumption. Reeder¹³⁴ used a strength of materials approach to estimate the shift in the neutral axis to correct the flexural rigidity of the delaminated arms with doubler plates to estimate the fracture energy. Other methods, such as the modified compliance calibration method, do not require this correction as the bending stiffness of the specimen is measured directly. The analytical approach proposed by Reeder was validated using MMB fracture tests, and similar results have been obtained from another study.⁹⁰ Furthermore, the correction proposed by Reeder is not valid for delaminations that occur in bi-materials, such as core-to-facesheet separation in sandwich composites. Due to an interfacial property mismatch between the facesheet and core, sandwich composites do not develop a pure mode I delamination response during modal fracture tests.^135,136 Although doubler-reinforced DCB tests have been used to evaluate the effectiveness of stitching on the mode I fracture energy of sandwich composites,¹⁰⁷ the unreinforced single cantilever beam test appears to be the most suitable standardized test method to evaluate the mode I fracture energy for stitched sandwich composites.^108,137,138 This difference is due to a dependency on the thickness of the doubler in estimating the fracture energy.¹³⁹ Therefore, it is not recommended to use doublers for bi-materials to evaluate the effectiveness of stitching without further investigation.

To improve the predictive capability of crack progression in stitched composites, experimental tests such as interlaminar tension tests^{6,54,56,59,63,124} and interlaminar shear tests¹²⁴ are often used to characterize the traction-separation response of the through-thickness reinforcement. Although, several analytical micromechanical models have been developed by Jain and Mai,^140–142 Cox,^88,143,144 and Plain¹⁴⁵ to predict the traction-separation response of a through-thickness reinforcement. These analytical methods do not include unstable pullout behavior after the failure of the through-thickness reinforcement. Tan et al.⁵⁹ used interlaminar tension tests to identify essential features of the fracture process of through-thickness stitching. The primary failure mechanisms during mode I fracture were identified as 1) debonding of the thread/matrix interface, 2) slack absorption, 3) thread failure, and 4) pullout friction.⁵⁹ Thread failure can occur at the interlocked region near the bobbin thread or near the crack interface, which can lead to variation in the measured load at which frictional sliding occurs.⁵⁴ Overall, the stitch traction-separation responses to simulate mode I delaminations were incorporated as a material model for nonlinear spring elements in a cohesive zone finite element model. Excellent correlation of the load-displacement responses between the experimental measurements and the finite element modeling predictions were achieved.⁵⁶

A lack of experimental data to verify analytical and computational approaches exists for unique material systems such as sandwich composites and high-temperature composites (carbon-carbon composites and CMCs), as shown in Table 3. High-temperature composites have received very little attention with regard to their interlaminar and fracture properties, which may be the result of the restricted access due to governmental regulations. Current available literature shows that in-plane mechanical properties and interlaminar shear strengths can be improved by incorporating stitches in a high-temperature composite.^2,9,52 The only test that has been used to evaluate the effectiveness of stitching in high-temperature composites is the compressive interlaminar shear test.¹²⁵ Corresponding test methods to evaluate unstitched high-temperature composites are the DCB and single-edge notch bend tests. These studies show that complications can occur during testing due to the complex woven architecture of high temperature composites. These complications include tow bridging,¹⁴⁶ material nonlinearity from pre-existing voids,¹⁴⁷ and energy-releasing mechanisms outside of the initial delamination plane.¹¹⁰

Summary

A literature review on the out-of-plane behavior of stitched composites is presented. Studies reveal that stitching generally improves the out-of-plane properties of polymer matrix composites. It should be noted that through-the-thickness reinforcements influence the in-plane properties (bulk moduli and tensile strength) as well as the out-of-plane properties (interlaminar properties and modal fracture energies). For aerospace structures, in-plane properties are the most critical in establishing the structural design. Reviews of in-plane properties of stitched composites can be found.^{5,9,10,37,148} Understanding the out-of-plane behavior is necessary from a damage-tolerant design perspective. In particular, mode I and mode II steady-state fracture energies of stitched polymer matrix composites have shown to be dependent on stitch parameters and may increase by a factor up to 15 and 3, respectively, for Vectran, carbon, and Kevlar stitching materials. Stitch density and linear thread density are the two primary stitch parameters that have shown to improve the out-of-plane properties. Under mode I delamination, stitches behave primarily as a bridging mechanism to resist crack growth. Stitches have also been observed to mainly deform and plow through the adjacent matrix during mode II loading conditions.

Current test standards do not appropriately address how to effectively determine the modal fracture energies of stitched composites. This is primarily due to the high rotational constraint developed by the through-thickness stitching that leads to facesheet failure and possible significant stitch bridging, which may lead to inaccuracies in the calculated fracture energy. Doublers and new test methods have been developed to prevent failure of the delaminated arms during testing. Standard procedures for determining doubler sizing and estimating appropriate initial crack lengths relative to the stitch location need to be developed.

Additional investigation is needed to understand the influence of unique stitch parameters, such as stitch pretension, stitch twist, and stitching style. Lastly, the influence of stitch parameters on high-temperature composites, such as carbon-carbon composites, CMCs, and sandwich composites, are not well studied. Research indicates that the out-of-plane properties of CMCs can be improved by approximately 30% without impacting in-plane properties by using untwisted low-filament count (1K) carbon fiber tows for stitching thread. Overall, proper selection of stitch parameters such as stitch density and linear thread density can further improve the out-of-plane properties of polymer matrix composites.

Footnotes

Declaration of Conflicting Interests

The author(s) declared no potential conflicts of interest with respect to the research, authorship, and/or publication of this article.

Funding

The author(s) disclosed receipt of the following financial support for the research, authorship, and/or publication of this article: This work was supported by the Air Force Research Laboratory (Award #: FA8650-19-2-2211). The authors received the support from Raspet Flight Research Laboratory’s Marvin B. Dow Stitched Composites Development Center and Dr. James Ratcliffe of NASA Langley Research Center.

ORCID iD

Rani W Sullivan

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Influence of stitching on the out-of-plane behavior of composite materials – A mechanistic review

Abstract

Keywords

Introduction

Stitching parameters for fracture characterization

Mode I fracture energy

Mode II fracture energy

Mixed-mode fracture energy

Interlaminar strength

Impact behavior

Test methods and analysis to characterize stitched composites

Summary

Footnotes

Declaration of Conflicting Interests

Funding

ORCID iD

References