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Experimental investigation of three-dimensional woven composites
Composites: Part A 73 (2015) 242–259
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Composites: Part A journal homepage: www.elsevier.com/locate/compositesa
Experimental investigation of three-dimensional woven composites Kyle C. Warren a,⇑, Roberto A. Lopez-Anido a, Jonathan Goering b a b
University of Maine, Advanced Structures and Composites Center, 35 Flagstaff Road, Orono, ME 04469, United States Albany Engineered Composites, 112 Airport Drive, Rochester, NH 03867, United States
a r t i c l e
i n f o
Article history: Received 2 December 2014 Received in revised form 3 March 2015 Accepted 5 March 2015 Available online 19 March 2015 Keywords: A. Polymer-matrix composites (PMCs) A. 3-Dimensional reinforcement B. Mechanical properties B. Fracture
a b s t r a c t This paper summarizes an extensive experimental study of composites reinforced with threedimensional woven preforms subjected to tensile, compressive and in-plane shear loading. Three innovative three-dimensional woven architectures were examined that utilize large 12 K and 24 K IM7 carbon tows, including two ply to ply angle interlock architectures and one orthogonal architecture. Additionally, a two-dimensional quasi-isotropic woven material was evaluated for comparison. Loads were applied in both the warp and the weft directions for tensile and compressive loading. Digital image correlation was used to investigate full field strains leading up to quasi-static failure. Experimental results including ultimate strengths and moduli are analyzed alongside representative failure modes. The orthogonal woven material was found to have both greater strength and modulus in tension and compression, though a ply to ply woven architecture was found to outperform the remaining three-dimensional architectures. Recommendations are made for improving the manufacturing processes of certain three-dimensional woven architectures. Ó 2015 Elsevier Ltd. All rights reserved.
1. Introduction Three-dimensional (3D) woven carbon composites are currently used in structural aerospace applications including fan blades and fan casings in the next generation CFM International Leading Edge Aviation Propulsion (LEAP) engine as well as the main landing gear brace for Boeing’s 787-8 Dreamliner. In light of these industrial applications, a comprehensive understanding of three innovative 3D woven architectures was sought for expanded use in aerospace structures. This study focused on quasi-static loading of these unique fiber architectures under tensile, compressive, and in-plane shear loads. Compressive and tensile loads were evaluated in both the longitudinal (warp) and transverse (weft) directions. The 3D architectures examined utilize IM7 carbon woven from large 12 K and 24 K tows, representing the current state-of-the-art in industrial aerospace 3D woven composites actively used in industry. Large tow sizes offer a balance between part performance and cost. Using 12 K and 24 K tows allows for scalability of the manufacturing process and consistency within the required equipment as part sizes increase. An analysis of failure modes and stress–strain responses has identified possible sources for material nonlinearity within select 3D fiber architectures. Detailed failure modes and mechanisms for ⇑ Corresponding author. Tel.: +1 207 720 0774. E-mail address: [email protected] (K.C. Warren). http://dx.doi.org/10.1016/j.compositesa.2015.03.011 1359-835X/Ó 2015 Elsevier Ltd. All rights reserved.
two-dimensional (2D) woven composites are available in the literature and will not be discussed in depth in this paper. There is a deficiency in the currently available literature in reliable tensile, compressive and shear data for 3D woven composites. A small number of experimental studies are available in literature and from those, the majority present results from limited sample sets or no longer represent the latest technology and practices in structural three-dimensional preforming. The data collected in this study and discussed in this paper accurately represent materials that are currently used in complex aerospace applications. Additionally, digital image correlation (DIC) was used for full field strain measurement and analysis for all experiments in this study. DIC is a very important and necessary tool to accurately capture strains within 3D woven materials due to large unit cell sizes. The use of appropriate strain measurement techniques for mechanical evaluation of three-dimensional woven materials in tension and compression is limited in the literature. Three-dimensional woven composite materials exhibit many benefits that traditional two-dimensional composites do not. Some of these benefits include improved fracture toughness, higher interlaminar tensile strength, reduced notch sensitivity and improved impact resistance [1,2]. Since three-dimensional woven composites do not consist of laminae as traditional two-dimensional composites do, delamination, or the separation of laminae, cannot occur in the classical sense of the word. A major advantage of using 3D woven materials is near net shape
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preforming; three-dimensional woven composite preforms can be woven into shapes resembling closely the shape of the final part [3]. Reduced fabrication labor, material scrap and part cost are among the advantages realized in industry [4]. While maintaining a constant fiber volume fraction between a two-dimensional and three-dimensional material, through-thickness mechanical characteristics can be enhanced. However, fewer fibers are then found inplane which can result in a reduction of in-plane properties. While this potential reduction of in-plane properties may initially be seen as a limitation, applications of select 3D woven architectures in aerospace structures can be notably more efficient than traditional two-dimensional woven materials. Effective elimination of interfacial failure between layers seen in two-dimensional composites coupled with material nonlinearity and improved fracture toughness allows for designers to use 3D woven composites in new and innovative ways. Tan et al. [5] experimentally evaluated a 3D orthogonal woven carbon fiber composite material. The preform was manufactured from Torayca T-300 (3 K) carbon fiber and molded with Epicote 828 epoxy resin. The resulting panels were characterized in tension with loading in both the warp direction and weft direction. Tensile strength, tensile modulus and Poisson’s ratio were reported for each case. It was found that for the orthogonal weave being studied, the tensile stress–strain response in both loading directions was essentially linear until ultimate failure. Filler tow waviness was found to significantly affect the stiffness and strength of the material. A fracture mechanism and fracture path for the studied orthogonal fiber architecture was identified; interfacial debonding of the z-fiber and the matrix and stuffer yarn pull-out combined with filler yarn/matrix debonding were found to be the primary failure mechanisms for samples loaded in the stuffer direction. Similarly, stuffer yarn/matrix debonding and filler yarn rupture with slight fiber pull-out was identified as the primary failure mechanism when loaded in the filler direction. It was also noted that strain gages would need to be considered carefully prior to use to be sure an adequate representation of the material is captured. Early work by Cox et al. [6] focused primarily on the compressive response of 3D woven angle interlock composites. The findings from this study illustrated that kink band formation in the stuffer tows was the key initiating event that ultimately led to failure. Initial kink bands were found to form near the interaction between misaligned stuffer tows and warp weaver tows. Chou et al. investigated three-dimensional woven composites with reinforcement in three and five directions [7]. Of relevance to the work presented in this paper, the authors examined 12 K orthogonal woven reinforcement with Epikote 828/Epikote 1001 matrix mixture. The tensile response of this material system experienced a material softening phenomenon between 0.5% and 1% tensile strain. Brandt et al. [2] analyzed a novel 3D orthogonal woven composite in tension, bending, interlaminar shear strength, compression, impact and penetration. The study concluded that the new material (orthogonal woven) had greater energy absorption capability and damage tolerance when compared with a comparable two-dimensional woven material. Additionally, the compressive response of the 3D woven was found to be in the same range as two-dimensional woven composites. Brandt et al. conducted a study on the effect of varying the amount of through-thickness reinforcement in multiple 3D woven glass/epoxy composites with respect to tensile, compressive and interlaminar shear strength [8]. The authors found that tensile properties decrease with increasing reinforcement in the through-thickness direction. Compressive strength was found to be greatest when 4% of the fibers were located in the throughthickness direction. When subjected to peel evaluation, the 3D
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materials exhibited improved fracture mechanics over the evaluated 2D architecture. One of the conclusions from this shift in fracture mechanics is an increase in energy absorption capacity and damage tolerance in the 3D materials. Dow Tactix 138 and Shell 1895 resins combined with AS4 carbon and S-2 glass reinforcement were used by Cox et al. [9] to evaluate 3D woven composites. Both orthogonal woven and angle interlock materials were evaluated. The majority of the experimental data reported was for single samples from various materials. Determination of engineering properties for each material was secondary to evaluating the relation between the macroscopic failure modes and microstructural failure mechanisms. It was reported that for compressive and tensile results, tows failed as discrete units. That is, when fibers within a single tow fail, the entire tow section fails. Observations from failure modes suggest that some of the 3D materials were able to redistribute stresses to reach much higher stains after a neighboring tow had failed. Bending failure modes were found to be a combination of both the previously identified compressive and tensile failure modes. Tensile loading was re-examined to capture additional important failure mechanisms [10]. Described in [11,12], layer-to-layer and through-the-thickness interlock woven carbon/epoxy composites with stuffer tow inclusions were investigated. A plastic straightening of the stuffer tows was determined to be the cause of a non-linear softening in the tensile stress–strain response of these composite materials. This straightening was described to occur as misaligned stuffer tow segments began to align with the direction of loading. Stuffer tow rupture and pullout were identified as the primary tensile failure modes. Stuffer tow failure occurred in most if not all stuffers prior to the ultimate failure of the tensile sample for 3D woven interlock composites. Callus et al. performed tensile evaluation and analysis of composites with three-dimensional orthogonal, normal layered interlock and offset layered interlock glass fiber reinforcement [13]. The authors discussed the stress–strain response found in the three materials and noted that, when compared with other experimental investigations in the literature (Cox et al. [10] and Chou et al. [7]), a reduction in tensile modulus was found as tensile strains exceeded 0.4% in both the warp and the weft directions. Using SEM examination, more extensive damage was found in the non-linear region of the tensile response than in the linear region and was attributed to some degree of inelastic tow straightening. This softening response was concluded to be unique to different types of fiber architectures and composite material systems. Leong et al. investigated the response of two similar three-dimensional orthogonal woven composites [14]. Stuffer tow pull out was noted as the primary failure mechanism for the composite when loaded in the warp direction. Extensive longitudinal tow splitting occurred just prior to complete failure. This phenomenon was thought be in part caused by extensive longitudinal matrix cracking from a Poisson’s ratio mismatch between the matrix and longitudinal tows. The compressive response of two orthogonal 3-axis woven composites using 24 K stuffer reinforcement with 12 K x and y (thickness and width) dry tows was examined by Kuo et al. [15]. The first 3-axis material was manufactured with carbon/epoxy stuffer rods and the second with dry tows. For both architectures, the progression of compressive damage was investigated. Kink bands within the stuffer rods and the stuffer tows were identified to be on either the ‘miniscopic’ (tow level) or microscopic (intratow) level. Kuo et al. more recently investigated three different 3-axis orthogonal composites with varying surface patterns in compression [16]. The surface yarn pattern (surface loop density) was found to influence the damage growth within the materials. A
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higher loop density was found to be desirable because it generally resulted in less weaving-induced tow distortion. Stig and Hallstrom evaluated the mechanical performance of a fully interlaced three-dimensional woven composite material [17]. Woven with 12 K T700 carbon tows and manufactured with vacuum injection molding of vinyl ester resin, this composite material experienced a tensile softening response similar to other 3D woven materials. Occurring between 0.5% and 1.0% strain, the softening is described to be indicative of a progressive tensile failure. Damage was found to be more localized in the threedimensional woven material when compared with a twill weave. Mouritz and Cox presented a comparison between stitched, pinned and 3D woven composites [18] from published data. The study mentions that the diameter of woven z-binders, stitches or pins is usually in the 0.1–1.0 mm range, citing that thicker z-binders may result in harmful distortion of the fiber architecture. It is observed that the strength and stiffness of 3D woven and stitched composites and the amount of through-thickness reinforcement are not strongly correlated. This suggests that by taking some of the otherwise in-plane fibers and orienting them through the thickness, benefits of through-thickness reinforced can be exploited without a dramatic sacrifice of in-plane properties. Bogdanovich et al. [19] performed an experimental study focused on joining 3D orthogonal woven composites. Through co-cured and adhesively bonded joints of three-dimensional Eglass preforms, this study concluded that stitching and stapling co-cured preforms prior to infusion greatly increased the singlelap joint strength. Including approximately 2% z-fibers by weight in the preform suppressed delamination during joint failure. Bogdanovich et al. [20] and Lomov et al. [21] have identified damage events during tensile loading of three-dimensional orthogonal woven carbon composites. These events correspond to certain acoustic emission event energy and signal peak frequency. The identified damage events were, in order of occurrence, sporadic cracking of boundary yarns, small transverse matrix cracks and massive transverse matrix cracking combined with tow/matrix debonding at interface boundaries. This paper is focused on exploring the following research objectives: 1. The three-dimensional woven composites in this paper have no biased (off-axis) reinforcement. The goal of this paper is to benchmark the performance of these 3D woven composites with a competitive industry standard two-dimensional woven material with bias reinforcement. 2. Identify material nonlinearities and make design recommendations based on these findings. 3. Characterization of failure modes and fracture paths experienced by the selected three-dimensional woven architectures. 4. Quantify the amount of crimp seen in three-dimensional ply to ply architectures and draw conclusions relating crimp to material properties. 5. Collect reliable data on various three-dimensional woven architectures for use in numerical modeling and analysis. 2. Fiber architectures Four different fiber architectures were examined during this study: a 3D orthogonal preform, two different 3D ply to ply angle interlock preforms, and a traditional two-dimensional weave. Geometric representations of general three-dimensional woven architectures are seen in Fig. 1. All materials were woven at Albany Engineered Composites in Rochester, NH, and molded with a resin transfer molding (RTM) process. IM7 carbon was used in all materials. Tow sizes were either 12 K or 24 K and are explicitly
defined for each material in subsequent sections. Toughened epoxy resins used were either ST-15 or PR-520. ST-15 has a tensile strength of 89.6 MPa and a tensile modulus of 3.31 GPa [22]. Likewise, the PR-520 resin system has a tensile strength of 82.1 MPa and a tensile modulus of 4.0 GPa [23]. IM7 carbon has a tensile strength of 5.6 GPa and a modulus of 276 GPa [24]. The two resin systems have similar physical and mechanical properties and the choice for their use reflects the current trend in industrial aerospace applications. Crimp of a tow is a textile description of the waviness of the tow based on the straightened length as a percentage of the unstretched length. As the waviness of a tow increases, the crimp is said to increase. To quantify crimp in the two ply to ply architectures, measurements were taken along cross sections from digital microscopy images. Crimp is related to tow waviness which is defined in this paper as the amplitude of a tow in a given direction divided by the wavelength of that tow. The tow waviness ratio is the tow waviness in the weft tows divided by the tow waviness in the warp tows. Tow waviness has a significant contribution to the mechanical performance of composites with three-dimensional woven ply to ply fiber reinforcement. 2.1. 3D orthogonal A three-dimensional orthogonal fiber architecture was examined. This preform used 24 K filler tows and 12 K warp weaver and warp stuffer tows. The average total fiber volume fraction for this composite was found to be 51.9%. The preform was molded to a nominal thickness of 4.19 mm with ST-15 resin. Four warp stuffer tows and five weft filler tows are woven as a preform connected by warp weaver tows that travel over two top surface picks and under the next two bottom surface picks. Based on weaving predictions, the warp weaver tows are estimated to be 8.7% of the total architecture by weight. A digital microscopy image illustrates the cross-section of this 3D orthogonal material in Fig. 2. From representative cross sections for this material, a unit cell size of repeating geometry was found to be 10.5 mm in the warp direction and 8.50 mm in the weft direction. The stuffer tows in this fiber architecture in Fig. 2b are notably straighter than the filler tows in Fig. 2a. The inherent waviness of the filler tows is in part a result of the warp weaver tow distorting the outer filler tows during the weaving process. 2.2. 3D ply to ply 12 K/24 K The second three-dimensional fiber architecture subjected to laboratory investigation used 12 K warp tows and 24 K weft tows arranged in a ply to ply angle interlock preform. The preform was molded to a nominal thickness of 4.19 mm with ST-15 resin, resulting in a fiber volume fraction of 50.4%. Fig. 3 depicts cross sections of the fiber architecture showing weft tows (a) and warp tows (b). The amount of crimp in the weft tows is notably greater than the warp tows. Measurements of the unit cell size were made from these representative cross sections and were found to be 10.5 mm in the warp direction and 8.36 mm in the weft direction. The warp tow wavelength was measured to be 10.5 mm with an amplitude of 0.34 mm. The weft tow wavelength was measured to be 8.36 mm with an amplitude of 0.55 mm. With these measurements, the resulting warp tow waviness is 0.0324 and the weft tow waviness is 0.0657. The tow waviness ratio for this architecture is 2.03. 2.3. 3D ply to ply 24 K/24 K The last three-dimensional fiber architecture evaluated was a ply to ply angle interlock with 24 K tows in both the warp and
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(a)
(b)
Fig. 1. Geometric representations of general three-dimensional architectures including (a) orthogonal and (b) ply to ply angle interlock. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
the weft directions. Molded with PR-520 toughened epoxy resin, the nominal thickness of this material was also 4.19 mm. The volume fraction of this material was 59.0%. Similar to the 12 K/24 K ply to ply architecture, the warp tows shown in Fig. 4b are found to have less crimp and are therefore
straighter than the weft tows seen in Fig. 4a. From representative cross-sections, a unit cell size was found to be 10.5 mm in the warp direction and 8.80 mm in the weft direction. The warp tow wavelength was 10.5 mm with an amplitude of 0.383 mm. The weft tow wavelength was 8.80 mm with an amplitude of 0.515 mm. The
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Fig. 2. 3D orthogonal digital microscopy image of fiber architecture. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
warp tow waviness and weft tow waviness are 0.0365 and 0.0585, respectively, resulting in a tow waviness ratio of 1.60 for this 3D architecture.
in-plane material properties and improves out-of-plane properties (e.g., inter-laminar shear strength, and impact toughness [18]). 3. Experimental method
2.4. 2D woven 3.1. Sample preparation In addition to three-dimensional preforms, a two-dimensional woven fiber architecture was also investigated. Eight layers of [0/ 90] were molded using the same RTM process described for the 3D fiber architectures. A [45, 0, 45, 90] symmetric layup was utilized, with layer orientations corresponding to the warp fiber orientations, to represent a quasi-isotropic material (in-plane). Each layer was woven from 12 K warp, 12 K weft tows in a plain weave. The nominal thickness was 3.30 mm. The fiber volume fraction was found to be 49.5%. This conventional two-dimensional woven configuration has no through-the-thickness fiber reinforcement and is therefore susceptible to interlaminar failure modes including delamination. 2.5. Fiber volume fraction When making comparisons between two-dimensional and three-dimensional composites of the same fiber volume fraction, one must consider the relative amount of in-plane fiber reinforcement. Two-dimensional composites maintain roughly all fibers inplane; traditional composites have no through-the-thickness fiber reinforcement. A significant percentage of the fibers within the 3D woven architectures are oriented through the thickness. Taking some of these fibers out of plane in 3D architectures reduces
To ensure the highest sample quality and process repeatability, all materials were cut using a waterjet with garnet abrasive. After samples were cut to the appropriate nominal dimensions, they were conditioned at 23 °C, 50% relative humidity for at least 24 h prior to measuring sample dimensions. 3.2. Digital image correlation With the relatively large unit cell size characteristic of the evaluated 3D woven materials, traditional strain gages have been found to exhibit large variation in measured strain. Studies focused on 3D woven materials have found that strain gauges need to be carefully considered to be sure that the data captured is representative of the material [5]. For this reason, all experimental work utilized the ARAMIS system for digital image correlation for accurate measurement of full field strains. Digital image correlation utilizes a specialized paint spatter pattern on the surface of samples to track points on that surface through time. Of particular interest is the size of the point tracked in space, called facets, and the separation distance between each facet, called the step. A facet is a collection of pixels in a neighborhood that represent a discrete point in a set of points called the
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Fig. 3. 3D ply to ply 12 K/24 K digital microscopy image of fiber architecture. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
field. By first establishing a strain-free state, a deformation gradient is initialized and updated as the camera system accurately tracks the movement of facets through a calibrated control volume in space. As the step between facets becomes smaller, the amount of information extracted from a sample in space increases. Previous work has identified a facet size of 15 pixels and a step of 13 pixels to produce accurate results for similar control volumes [25,26]; all strain fields were analyzed with these settings. 4. Procedure Laboratory procedures are described in this section for tensile, compressive and in-plane shear evaluation. Samples from each material were evaluated with loading in the warp direction and the weft direction for tension and compression. Additionally, samples were subjected to in-plane shear loading. A minimum of five samples were considered for each data set. Standard laboratory atmosphere, 23 °C, 50% relative humidity, was maintained within the laboratory during all mechanical evaluation. 4.1. Tension The procedure for evaluating the tensile response of the four materials was in accordance with ASTM D3039, Tensile Properties of Polymer Matrix Composite Materials. Experiments were run with displacement control at a rate of 1.27 mm/min on a 500 kN Instron servo-hydraulic frame. A 250 kN load cell was used with 250 kN hydraulic wedge grips. Samples were centered between the wedge grips prior to gripping at 20 MPa. The ARAMIS system for DIC was calibrated with a
100 mm 80 mm 80 mm control volume. The nominal width of tensile samples measured 25.4 mm. The corresponding gage area is dependent upon the actual material thickness as listed previously. A macro script was created within the ARAMIS system to ensure that the area in the gage used for calculation of strain was held constant from one sample to another. This calculation area measured approximately 19 mm wide and 100 mm long. Using this method, numerous unit cells were captured for an average strain calculation. The tensile modulus of elasticity was calculated by assuming a linear elastic region between 1000 and 3000 microstrain. 4.2. Compression The combined loading compression (CLC) fixture was used to evaluate materials in compression in accordance with ASTM D6641. Experiments were conducted in displacement control. The standard stipulates that any given experiment should last between 1 and 10 min to be considered quasi-static. At the recommended initial load head displacement rate of 1.27 mm/min, experiments did not last for one minute, prompting an adjustment of the displacement rate allowing experiments to last between 2 and 3 min, corresponding to a rate of 0.254 mm/min. The nominal width of compressive samples measured 12.7 mm. The gage area is dependent upon each of the material thicknesses as listed previously. Evaluation of the four materials was performed in both the longitudinal (warp loading) and the transverse (weft loading) directions. The ARAMIS system was calibrated to a 35 mm 28 mm 28 mm control volume for full field strain analysis. A 100 kN Instron frame was used with a 100 kN load cell.
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Fig. 4. 3D ply to ply 24 K/24 K digital microscopy image of fiber architecture. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
Plates were used in the grip wedges to create flat surfaces for transmitting loads to the CLC fixture. Calculation of compressive strains for each sample using the ARAMIS system involved creating a macro script for repeatable selection of the calculation area. The calculation area was selected to be approximately 10 mm 10 mm, centered in the gage, to include as much of the gage area as possible. The compressive modulus of elasticity was calculated by assuming the linear elastic region to be between 1000 and 3000 microstrain. The sample thickness was sufficient to prevent buckling of the material during compressive loading. 4.3. In-plane shear Qin et al. [27] evaluated a three-dimensional woven material with an orthogonal architecture. Digital image correlation was used with the Iosipescu shear fixture. The investigation found that, given the small unit cell size of the particular orthogonal architecture, measuring approximately 2.4 mm, the use of the Iosipescu shear fixture was appropriate and allowed for multiple repeating unit cells along the gage length. The unit cell of the 24 K warp, 24 K weft 3D woven ply to ply architecture in this paper is roughly 9.8 mm in the warp direction and 8.1 mm in the weft. The other three-dimensional materials have unit cells of similar sizes. To capture the response of three-dimensional preforms woven with 12 K and 24 K tows, a gage area that captures more repeating geometries within a material is preferable. The gage length used with the Rail Shear Fixture from ASTM D7078 nominally measures 30.5 mm, whereas the gage length of samples utilizing the Iosipescu shear fixture is 12 mm. In-plane shear response was therefore evaluated in accordance
with ASTM D7078, Shear Properties of Composite Materials by VNotched Rail Shear Method. The gage areas for each material are dependent upon their respective thicknesses. Samples were placed in the Rail Shear Fixture and bolts in the fixture were torqued 54 Nm while the sample was held square to the fixture and centered. The fixture was placed in a 100 kN Instron frame with a 25 kN load cell. The ARAMIS system was used for shear strain measurement with a 65 mm 52 mm 52 mm calibrated control volume. A macro script was created to generate consistent calculation areas for shear angle from the ARAMIS strain field. The calculation area used for all samples evaluated for inplane shear response was 6 mm wide (perpendicular to the gage length) and 29 mm long, centered between the notches.
5. Failure modes 5.1. Tension Three-dimensional woven ply to ply composites demonstrate tensile failure mechanisms that are sensitive to the amount of crimp in the warp direction relative to the weft direction. When evaluated in the warp direction, warp tow rupture and withdrawal from surrounding weft tows was evident. Failure planes were found to be repeatable between numerous samples and located between the weft tow/matrix interfaces. The fracture path began on one surface of the sample and propagated directly through the thickness, and normal to the direction of loading. Matrix cracking was also observed on the surface of the samples, adjacent to the warp tows and parallel to loading. Representative failure modes for both ply to ply architectures, 12 K/24 K and 24 K/24 K, are seen in
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Fig. 5. 3D ply to ply failure modes for (a) 12 K/24 K architecture and (b) 24 K/24 K architecture. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
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Fig. 6. 3D ply to ply inclined failure mode for (a) 12 K/24 K and (b) 24 K/24 K fiber architectures. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
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Fig. 5a and b, respectively. This observed failure mode was consistent between both ply to ply architectures when loaded in the warp direction. Evaluation with loading in the weft direction revealed different failure modes. For the ply to ply architecture with 12 K tows in the warp and 24 K tows in the weft direction, the failure path is inclined with respect to the weft tows. A representative failed sample is seen in Fig. 6a. The strain to failure of the composite was found to be 1.46% with a 5.54% coefficient of variation (COV) when loaded in the warp direction and 2.13% (14.8% COV) when loaded in the weft direction. This discrepancy in strain to failure can be partially explained by the amount of crimp in the weft tows vs. the warp tows. Weft tows within the studied ply to ply architectures have more crimp than warp tows. The crimp ratios for the 12 K/24 K and 24 K/24 K three-dimensional architectures described previously are 2.03 and 1.60, respectively. This characteristic is thought to allow the material to experience higher strain to failure when loaded in the weft direction as the weft tows begin to straighten. This failure mode was seen in all samples of this architecture subjected to the same experimental conditions. The trend in the strain to failure for the 24 K/24 K ply to ply architecture is similar to that of the 12 K/24 K ply to ply architecture in that the strain to failure when loaded in the warp direction was 1.85% (2.86% COV) as compared with 2.67% (6.24% COV) when loaded in the weft direction. The observed failure path, however, was found to also be inclined with respect to the weft tows but in the opposite direction, as seen in Fig. 6b. This discrepancy in a
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matrix-dominated onset of ultimate failure between two distinctly different three-dimensional woven ply to ply architectures may partially be caused by the different tow sizes found in the warp and the weft directions as well as the mismatch in matrix between the architectures. Fig. 7 illustrates the typical stress–strain responses observed for both ply to ply fiber architectures. A bilinear response or material softening can be seen in each of the curves. The trend is more pronounced for each of the two materials when loading is oriented in the weft direction. Weft samples experience not only lower strengths, but also reduced linear elastic properties and greater strain to failure. The trend in weft and warp loaded bilinearity for both materials can be in part attributed to each of the materials having more crimp in the weft direction. The crimp in a tow tends to reduce when the tow is subjected to strain. This reduction imparts strain on the surrounding matrix. The greater the amount of crimp found in a tow, the stronger this tendency is. The onset of this phenomenon in warp tows (similar to ‘plastic tow straightening’ seen by others [7,10,13]) results in localized matrix damage, corresponding to the transition region between the linear-elastic region and the second region seen in Fig. 7. Unlike the work by Cox et al. [10], the evaluated ply to ply architectures did not include any stuffer tows in the warp direction and therefore the associated tow straightening is experienced in the weaver tows. Fig. 8 shows the tensile strains developed in the loading direction for the 24 K/24 K ply to ply architecture for loading in both the warp (8a and b) and weft directions (8c and d). Each of these
Fig. 7. Tensile stress–strain response for 3D ply to ply architectures loaded in the warp and the weft directions. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
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direction from the warp to the through-thickness direction. A typical tensile failure can be seen in Fig. 9a, with a magnification highlighting stuffer tow and warp weaver rupture in Fig. 9b. When subjected to loading in the weft direction, a strain to failure of 1.43% (7.14% COV) was experienced. Filler tow rupture was found to be the primary failure mechanism, with matrix cracking along the warp weavers and parallel to the stuffer tows. This suggests further agreement with Tan et al. [5] where a weft loaded orthogonal architecture was also found to fail adjacent to stuffer tows. 5.2. Compression
Fig. 8. Tensile strain in the loading (vertical) direction for 3D ply to ply 24 K/24 K sample loaded in the warp direction (a and b) and weft direction (c and d). Strain fields (a) and (c) correspond to a state just prior to the bilinear transition region whereas strain fields (b) and (d) correspond to a strain state just after this transition region. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
figures represent strain fields from just prior to the bilinear transition region (8a and c) and just after the transition (8b and d). Areas of high surface strains were found in the matrix material with notably higher strain concentrations in the weft loaded direction. This localized strain and corresponding damage experienced by the matrix is a contributor to the bilinear stress–strain behavior seen in both ply to ply architectures. Observed failure modes in the 3D orthogonal material coincide with a previous study by Tan et al. [5]. Warp loaded tension samples were found to exhibit a linear stress–strain response to failure. With a 1.61% (4.50% COV) strain to failure when loaded in the warp stuffer direction, fracture planes were identified along the filler tows accompanied by matrix/filler tow debonding. Additionally, stuffer tow withdrawal was seen with rupture found in the stuffer tows. Leong et al. noted a similar stuffer tow withdrawal that was accompanied by longitudinal (stuffer tow) cracking prior to failure [14]. Longitudinal stuffer tow cracking was not found in any of the evaluated samples in this experimental investigation. Matrix cracking was present in the surrounding composite parallel and adjacent to the filler tows. The warp weaver tows were found to have ruptured near the global failure when the tow changes
Different failure modes were observed in the warp and the weft directions when subjected to compressive loading conditions for the 3D orthogonal fiber architecture. A typical sample failure for warp-direction loading is seen in Fig. 10a. Warp stuffer tows are seen to be relatively straight running in the warp (left–right) direction. This straightness is a result of the weaving process for this orthogonal preform. When on the loom, the warp stuffer tows are kept in tension throughout the weaving process. The high degree of stuffer tow alignment has led to a mild brooming-type failure mixed with a through-thickness angled failure caused by macroscopic compressive stuffer tow failures and kink band formation. Unlike warp stuffer tows, weft filler tows are not kept in tension during the weaving process for this preform. Fig. 10b depicts a weft-loaded failure of the same material. It is seen that the weft filler tows, now running from the left to the right in the figure, do not maintain the same level of fiber straightness through the gage area of the sample, leading to an angled through-thickness failure from fiber microbuckling. The stress–strain response from this orthogonal woven material was essentially linear until failure for both the warp and weft loaded cases. Callus et al. [13] noted a softening in the stress–strain response for both warp and weft loaded compression of an orthogonal woven composite that was not seen in these samples. Compressive failure modes were found to be largely a concurrent debonding between the matrix and the surrounding tows coupled with tow microbuckling and isolated kink band formation for both the 3D ply to ply architectures. Failure modes and fracture paths found in samples loaded in the weft direction were similar to those found in warp direction loading. Typical warp-loaded sample failures are found in Fig. 11a and b for the 12 K/24 K ply to ply and 24 K/24 K ply to ply architectures, respectively. For comparison, a typical 2D woven quasi-isotropic failure is seen in Fig. 11c. Kink band formation as a primary compressive failure mode was found to coincide with previous work by Kuo et al. [15] and Cox et al. [6]. 5.3. In-plane shear In-plane shear samples were oriented in the warp/weft plane with the warp tows perpendicular to the gage length as defined between the notches in the samples. Failure modes observed during shear evaluation were consistent between both 3D ply to ply architectures. Shear strains were found to be greatest in the matrix material following the interface between one weft tow and the adjacent weft tow and were isolated to the gage area between the two notches. Shear loads are initially carried by the warp and weft tows in the ply to ply architectures. As shear strains build in the matrix it begins to degrade. This degradation is thought to transfer the initial shear loads from the tows to the weakening matrix. Concurrently, the weft tows begin to angle with increased shear loading. The combination of these two observed failure mechanisms leads to the softening seen in the shear stress–strain response in Fig. 12.
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Fig. 9. Typical (a) 3D orthogonal tensile failure in warp loading with (b) magnification of 3D orthogonal stuffer tow and warp weaver tow rupture. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
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Fig. 10. Typical (a) longitudinal and (b) transverse 3D orthogonal compressive failures. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
The 3D orthogonal woven architecture experienced a similar response and failure mechanism to the two 3D woven ply to ply architectures. Shear angles were once again found to be isolated to the area between the notches. Seen in Fig. 13 is a shear angle field distribution of the orthogonal architecture in a representative sample prior to failure. Highlighted as areas of greater shear strain are the matrix pockets found between filler tows along the gauge. The discontinuities in these matrix rich regions between filler tows are warp weaver tows on the surface of the material. Similar to the ply to ply architecture, significant matrix cracking is observed in the regions just described. 6. Results and discussion 6.1. Tension Tensile experimental results are presented in Table 1. Poisson ratios are also presented as calculated in the linear elastic region used for modulus calculation. The 3D orthogonal woven material
outperformed all other evaluated materials in tension for both strength and tensile modulus, with the exception of the two-dimensional woven material in transverse loading. The 3D orthogonal preform is made from five warp stuffer tows woven to four weft filler tows with orthogonal through-thickness reinforcement. The reduction of strength and stiffness seen in weft vs. warp loading conditions is attributed to the mismatch in inplane fibers in the warp and the weft directions. The weft-loaded modulus was found to be 84.9% of the warp-loaded modulus, reflecting the 4:5 weft to warp tow ratio. Weft loaded strength was found to be 61.7% of the warp loaded strength. Deviations from the 4:5 weft to warp tow ratio are likely a result of tow waviness and the addition of warp weaver tows. The three-dimensional ply to ply architectures, when compared with one another, reveal that the 24 K warp, 24 K weft preform has greater strength and stiffness for both warp and weft direction loading conditions. The transverse (weft-loaded) tensile strength of the 24 K/24 K ply to ply material is 54% of the warp loaded strength. For the 12 K warp, 24 K weft ply to ply architecture, the
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Fig. 11. Typical longitudinal compressive failure for (a) ply to ply 12 K/24 K, (b) ply to ply 24 K/24 K and (c) quasi-isotropic 2D woven fiber architectures. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
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Fig. 12. Shear stress–strain response for 3D architectures. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
tensile strength when loaded in the weft direction is 46% of the tensile strength when loaded in the warp direction. This suggests that despite volume fraction differences between architectures, a mismatch of 12 K warp, 24 K weft tows combined with greater crimp in the weft tows yields a proportionally lower strength when comparing the weft strength to the warp strength. Although the resin systems used for each of these two 3D architectures were different, this is thought to have only a minor contribution to the mechanical properties reported. Trends in material nonlinearity were common between both material systems. 6.2. Compression Compressive results are presented in Table 2. Of the 3D fiber architectures examined, the orthogonal woven was found to have the highest strengths when loaded in both the warp and the weft directions. This is attributed to the straightness of the warp stuffer and weft filler tows. Both ply to ply architectures lack the inclusion of ‘straight’ fibers within the respective preforms, resulting in noticeably lower compressive strengths. A trend seen in the compressive behavior for each of the 3D architectures is lower strength and stiffness when loaded in the weft direction as compared with the warp direction. This trend can once again be explained by the increase in crimp in the weft direction as compared with crimp in the warp direction for both ply to ply architectures. This induced waviness transfers stresses from tows to the matrix, resulting in the observed lower strengths and stiffnesses.
Compressive failures seen in ply to ply architectures with debonding between matrix and surrounding tows oriented in the loading direction were isolated when compared with two-dimensional woven failures. Two-dimensional failure modes were observed to contain extensive material brooming protruding beyond the bounds of the original thickness. Through-thickness reinforcement found in the 3D woven architectures contains the failure to closely within the original thickness. By doing this, certain failure modes are minimized or eliminated, including delamination. The ply to ply architecture with 24 K warp, 24 K weft tows has a higher volume fraction than the other materials, which likely contributes to the increase in strength and stiffness from the other ply to ply fiber architecture. Additionally, the 12 K/24 K ply to ply architecture has 27% more tow waviness than the 24 K/24 K ply to ply architecture. Compressive strength and stiffness from weft loading normalized to the compressive strength and stiffness from warp loading is one possible comparison to evaluate the performance of a material when subjected to a range of loading conditions. Normalized strengths and stiffnesses can also be used as a measure of the three-dimensional weaving process with regards to weaving-induced crimp and other tow nonlinearities. A visual comparison between all evaluated materials is seen in Fig. 14. The ply to ply 24 K/24 K architecture has greater relative stiffness ratios from a weft-loaded to warp-loaded orientation in compression as compared with the 12 K/24 K ply to ply architecture. The strength ratio,
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Fig. 13. Shear angle along the gauge area in the matrix for the 3D orthogonal fiber architecture. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
Table 1 Experimental tensile results. Fiber architecture
Tensile strength MPa (COV) 0°
3D 3D 3D 2D
orthogonal ply to ply 12 K/24 K ply to ply 24 K/24 K woven
1150 703 855 765
Tensile modulus GPa (COV)
90°
(4.83%) (7.45%) (7.82%) (3.41%)
710 323 461 717
0°
(7.81%) (3.44%) (3.84%) (3.79%)
69.6 53.4 62.5 49.9
m12 (COV)
90°
(2.61%) (5.56%) (2.88%) (6.06%)
59.1 33.0 43.0 46.6
(2.08%) (3.78%) (2.61%) (4.97%)
0.066 0.257 0.244 0.330
(13.8%) (4.93%) (11.3%) (5.15%)
Table 2 Experimental compressive results. Fiber architecture
Compressive strength MPa (COV) 0°
3D 3D 3D 2D
orthogonal ply to ply 12 K/24 K ply to ply 24 K/24 K woven
494 270 301 483
(3.23%) (6.15%) (7.07%) (8.02%)
however, was found to be slightly smaller in the 24 K/24 K ply to ply material as compared with the 12 K/24 K material. The compressive weft to warp stiffness ratio for the 3D orthogonal fiber architecture was found to be greater than all other relative weft to warp stiffness or strength ratios for all evaluated materials. 6.3. In-plane shear The examined architectures have no biased plies; there is no fiber reinforcement found in the three-dimensional woven
Compressive modulus GPa (COV)
90° 328 203 217 435
(2.81%) (3.78%) (6.76%) (3.33%)
0° 55.6 47.9 59.3 45.5
(10.3%) (7.58%) (6.67%) (4.98%)
90° 53.3 31.2 41.2 41.4
(7.68%) (2.20%) (4.16%) (2.08%)
materials outside of the orthogonal warp, weft and through-thickness directions. An observed nonlinearity and softening in the shear stress–strain response was found in all 3D architectures. This nonlinearity is not specific to 3D woven architectures. Plain woven [0/90] carbon composites of similar fiber volume fraction have exhibited similar trends in shear stress–strain nonlinearity [28]. The primary difference seen in 3D woven composites, including the three architectures examined in this paper, is that the strain to failure is much greater for 3D woven composites than for 2D. Experimental in-plane shear results are presented in
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Fig. 14. Weft-loaded compressive strength and stiffness normalized to compressive warp-loaded strength and stiffness. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
Table 3 Experimental shear results. Fiber architecture 3D orthogonal 3D ply to ply 12 K/24 K 3D ply to ply 24 K/24 K
Shear strength MPa (COV)
Shear modulus GPa (COV)
84.8 (3.07%) 89.6 (3.29%) 106 (3.68%)
3.94 (4.68%) 4.10 (5.07%) 5.00 (4.16%)
Table 3. The increased shear performance seen in the 24 K warp, 24 K weft ply to ply architecture over the 12 K/24 K can be partially attributed to the higher fiber volume fraction found in the 24 K/ 24 K material. 7. Conclusions Comprehensive mechanical evaluation of four composite materials has been discussed. Composites molded with three different 3D fiber architectures were evaluated and compared with a baseline 2D woven composite. Tensile, compressive and in-plane shear properties were examined. One of the main benefits of a composite molded from a 3D woven preform is that the fiber architecture can be tailored to achieve specific types of performance based on the application. Comparing the response of an orthogonal woven composite to a ply to ply woven composite demonstrates this benefit. Recognizing that other three dimensional woven
architectures are possible, this paper highlights only three specific architectures. As a result of the experimental investigation presented, the most important conclusions are: 1. Material nonlinearity is seen in the tensile response of 3D woven composites with ply to ply architectures when loaded in the weft direction. This nonlinearity has been explained and is attributed to the degree of crimp in the warp direction relative to the weft direction. Nonlinearity is also observed in orthogonal architectures under shear and select tensile loading. As a result of these significant nonlinearities, strain-based design should be considered with 3D woven composites. 2. Failure mechanisms for 3D woven composites in tension, compression and in-plane shear have been identified and characterized. One common initial failure mechanism witnessed in all evaluated samples was debonding of the matrix from tow inclusions for both resin systems. 3. The ratios between compressive warp-loaded strength and stiffness normalized to compressive weft-loaded strength and stiffness have been presented. This architecture-dependent relationship was used to compare the effects of crimp in compressive loading for each 3D woven material. Similar comparisons can be made in samples loaded in tension. The weaving process can be tailored to adjust the relative crimp between the warp tows and the weft tows. If a similar material response
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between warp loading and weft loading is desired, it is recommended that tension be controlled in weft tows during the weaving process to result in a tow waviness ratio closer to 1.0. 4. A reliable set of quasi-static data from three novel 3D woven architectures is now available. This data can be utilized for modeling of 3D woven materials as well as benchmarking of new architectures to these existing materials that are actively used in the aerospace industry.
Acknowledgments The research presented in this paper was supported by Albany Engineered Composites. The first author was supported in part by the Sakellaris Graduate Fellowship through the College of Engineering at the University of Maine. References [1] Bogdanovich AE. Multi-scale modeling, stress and failure analyses of 3-D woven composites. J Mater Sci 2006;41(20):6547–90. ISSN 0022-2461. [2] Brandt J, Drechsler K. Manufacture and performance of carbon/epoxy 3-D woven composites. In: 37th International SAMPE symposium; 1992, p. 864– 77. [3] Mouritz A, Bannister M, Falzon P, Leong K. Review of applications for advanced three-dimensional fibre textile composites. Compos Part A: Appl Sci Manuf 1999;30(12):1445–61. ISSN 1359-835X. [4] McClain M, Goering J. Overview of recent developments in 3D structures. In: SME composites manufacturing; 2012, p. 1–12. [5] Tan P, Tong L, Steven G, Ishikawa T. Behavior of 3D orthogonal woven CFRP composites. Part I. Experimental investigation. Compos Part A: Appl Sci Manuf 2000;31(3):259–71. ISSN 1359-835X. [6] Cox B, Dadkhah M, Inman R, Morris W, Zupon J. Mechanisms of compressive failure in 3D composites. Acta Metallurgica et Materialia 1992;40(12): 3285–98. ISSN 0956-7151. [7] Chou S, Chen H-C, Chen H-E. Effect of weave structure on mechanical fracture behavior of three-dimensional carbon fiber fabric reinforced epoxy resin composites. Compos Sci Technol 1992;45(1):23–35. ISSN 0266-3538. [8] Brandt J, Drechsler K, Arendts F-J. Mechanical performance of composites based on various three-dimensional woven-fibre preforms. Compos Sci Technol 1996;56(3):381–6. ISSN 0266-3538. [9] Cox B, Dadkhah M, Morris W, Flintoff J. Failure mechanisms of 3D woven composites in tension, compression, and bending. Acta Metallurgica et Materialia 1994;42(12):3967–84. ISSN 0956-7151. [10] Cox BN, Dadkhah MS, Morris W. On the tensile failure of 3D woven composites. Compos Part A: Appl Sci Manufact 1996;27(6):447–58. ISSN 1359-835X.
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[11] Cox B, Dadkhah M. The macroscopic elasticity of 3D woven composites. J Compos Mater 1995;29(6):785–819. ISSN 0021-9983. [12] Xu J, Cox BN, McGlockton M, Carter W. A binary model of textile composites – II. The elastic regime. Acta Metallurgica et Materialia 1995;43(9):3511–24. ISSN 0956-7151. [13] Callus P, Mouritz A, Bannister M, Leong K. Tensile properties and failure mechanisms of 3D woven GRP composites. Compos Part A: Appl Sci Manuf 1999;30(11):1277–87. ISSN 1359-835X. [14] Leong K, Lee B, Herszberg I, Bannister M. The effect of binder path on the tensile properties and failure of multilayer woven CFRP composites. Compos Sci Technol 2000;60(1):149–56. ISSN 0266-3538. [15] Kuo W-S, Ko T-H. Compressive damage in 3-axis orthogonal fabric composites. Compos Part A: Appl Sci Manuf 2000;31(10):1091–105. ISSN 1359-835X. [16] Kuo W-S, Ko T-H, Chen C-P. Effect of weaving processes on compressive behavior of 3D woven composites. Compos Part A: Appl Sci Manuf 2007;38(2):555–65. ISSN 1359-835X. [17] Stig F, Hallstrom S. Assessment of the mechanical properties of a new 3D woven fibre composite material. Compos Sci Technol 2009;69(11):1686–92. ISSN 0266-3538. [18] Mouritz A, Cox B. A mechanistic interpretation of the comparative in-plane mechanical properties of 3D woven, stitched and pinned composites. Compos Part A: Appl Sci Manuf 2010;41(6):709–28. ISSN 1359-835X. [19] Bogdanovich AE, Dannemann M, Dll J, Leschik T, Singletary JN, Hufenbach WA. Experimental study of joining thick composites reinforced with non-crimp 3D orthogonal woven E-glass fabrics. Compos Part A: Appl Sci Manuf 2011;42(8):896–905. ISSN 1359-835X. [20] Bogdanovich AE, Karahan M, Lomov SV, Verpoest I. Quasi-static tensile behavior and damage of carbon/epoxy composite reinforced with 3D noncrimp orthogonal woven fabric. Mechan Mater 2013;62:14–31. ISSN 01676636. [21] Lomov S, Karahan M, Bogdanovich A, Verpoest I. Monitoring of acoustic emission damage during tensile loading of 3D woven carbon/epoxy composites. Textile Res J 2014. 0040517513519510ISSN 0040-5175. [22] ST15 RTM Resin, Hexcel; 2013. [23] CYCOM PR 520 RTM Resin System, Cytec Engineered Materials; 2012. [24] HexTow IM7 Carbon Fiber, Hexcel; 2014. [25] Berube KA. Integration of process parameter control and digital image correlation methods in an investigation of the variability of marine polymer matrix composite material properties, PhD dissertation. University of Maine; 2012. [26] Melrose PT, Lopez-Anido R, Muszyski L. Elastic properties of sandwich composite panels using 3-D digital image correlation with the hydromat test system, MS thesis, University of Maine; 2004. [27] Qin L, Zhang Z, Li X, Yang X, Feng Z, Wang Y, et al. Full-field analysis of shear test on 3D orthogonal woven C/C composites. Compos Part A: Appl Sci Manuf 2012;43(2):310–6. ISSN 1359-835X. [28] Aly-Hassan MS, Hatta H, Wakayama S, Watanabe M, Miyagawa K. Comparison of 2D and 3D carbon/carbon composites with respect to damage and fracture resistance. Carbon 2003;41(5):1069–78. ISSN 00086223.
Contents lists available at ScienceDirect
Composites: Part A journal homepage: www.elsevier.com/locate/compositesa
Experimental investigation of three-dimensional woven composites Kyle C. Warren a,⇑, Roberto A. Lopez-Anido a, Jonathan Goering b a b
University of Maine, Advanced Structures and Composites Center, 35 Flagstaff Road, Orono, ME 04469, United States Albany Engineered Composites, 112 Airport Drive, Rochester, NH 03867, United States
a r t i c l e
i n f o
Article history: Received 2 December 2014 Received in revised form 3 March 2015 Accepted 5 March 2015 Available online 19 March 2015 Keywords: A. Polymer-matrix composites (PMCs) A. 3-Dimensional reinforcement B. Mechanical properties B. Fracture
a b s t r a c t This paper summarizes an extensive experimental study of composites reinforced with threedimensional woven preforms subjected to tensile, compressive and in-plane shear loading. Three innovative three-dimensional woven architectures were examined that utilize large 12 K and 24 K IM7 carbon tows, including two ply to ply angle interlock architectures and one orthogonal architecture. Additionally, a two-dimensional quasi-isotropic woven material was evaluated for comparison. Loads were applied in both the warp and the weft directions for tensile and compressive loading. Digital image correlation was used to investigate full field strains leading up to quasi-static failure. Experimental results including ultimate strengths and moduli are analyzed alongside representative failure modes. The orthogonal woven material was found to have both greater strength and modulus in tension and compression, though a ply to ply woven architecture was found to outperform the remaining three-dimensional architectures. Recommendations are made for improving the manufacturing processes of certain three-dimensional woven architectures. Ó 2015 Elsevier Ltd. All rights reserved.
1. Introduction Three-dimensional (3D) woven carbon composites are currently used in structural aerospace applications including fan blades and fan casings in the next generation CFM International Leading Edge Aviation Propulsion (LEAP) engine as well as the main landing gear brace for Boeing’s 787-8 Dreamliner. In light of these industrial applications, a comprehensive understanding of three innovative 3D woven architectures was sought for expanded use in aerospace structures. This study focused on quasi-static loading of these unique fiber architectures under tensile, compressive, and in-plane shear loads. Compressive and tensile loads were evaluated in both the longitudinal (warp) and transverse (weft) directions. The 3D architectures examined utilize IM7 carbon woven from large 12 K and 24 K tows, representing the current state-of-the-art in industrial aerospace 3D woven composites actively used in industry. Large tow sizes offer a balance between part performance and cost. Using 12 K and 24 K tows allows for scalability of the manufacturing process and consistency within the required equipment as part sizes increase. An analysis of failure modes and stress–strain responses has identified possible sources for material nonlinearity within select 3D fiber architectures. Detailed failure modes and mechanisms for ⇑ Corresponding author. Tel.: +1 207 720 0774. E-mail address: [email protected] (K.C. Warren). http://dx.doi.org/10.1016/j.compositesa.2015.03.011 1359-835X/Ó 2015 Elsevier Ltd. All rights reserved.
two-dimensional (2D) woven composites are available in the literature and will not be discussed in depth in this paper. There is a deficiency in the currently available literature in reliable tensile, compressive and shear data for 3D woven composites. A small number of experimental studies are available in literature and from those, the majority present results from limited sample sets or no longer represent the latest technology and practices in structural three-dimensional preforming. The data collected in this study and discussed in this paper accurately represent materials that are currently used in complex aerospace applications. Additionally, digital image correlation (DIC) was used for full field strain measurement and analysis for all experiments in this study. DIC is a very important and necessary tool to accurately capture strains within 3D woven materials due to large unit cell sizes. The use of appropriate strain measurement techniques for mechanical evaluation of three-dimensional woven materials in tension and compression is limited in the literature. Three-dimensional woven composite materials exhibit many benefits that traditional two-dimensional composites do not. Some of these benefits include improved fracture toughness, higher interlaminar tensile strength, reduced notch sensitivity and improved impact resistance [1,2]. Since three-dimensional woven composites do not consist of laminae as traditional two-dimensional composites do, delamination, or the separation of laminae, cannot occur in the classical sense of the word. A major advantage of using 3D woven materials is near net shape
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preforming; three-dimensional woven composite preforms can be woven into shapes resembling closely the shape of the final part [3]. Reduced fabrication labor, material scrap and part cost are among the advantages realized in industry [4]. While maintaining a constant fiber volume fraction between a two-dimensional and three-dimensional material, through-thickness mechanical characteristics can be enhanced. However, fewer fibers are then found inplane which can result in a reduction of in-plane properties. While this potential reduction of in-plane properties may initially be seen as a limitation, applications of select 3D woven architectures in aerospace structures can be notably more efficient than traditional two-dimensional woven materials. Effective elimination of interfacial failure between layers seen in two-dimensional composites coupled with material nonlinearity and improved fracture toughness allows for designers to use 3D woven composites in new and innovative ways. Tan et al. [5] experimentally evaluated a 3D orthogonal woven carbon fiber composite material. The preform was manufactured from Torayca T-300 (3 K) carbon fiber and molded with Epicote 828 epoxy resin. The resulting panels were characterized in tension with loading in both the warp direction and weft direction. Tensile strength, tensile modulus and Poisson’s ratio were reported for each case. It was found that for the orthogonal weave being studied, the tensile stress–strain response in both loading directions was essentially linear until ultimate failure. Filler tow waviness was found to significantly affect the stiffness and strength of the material. A fracture mechanism and fracture path for the studied orthogonal fiber architecture was identified; interfacial debonding of the z-fiber and the matrix and stuffer yarn pull-out combined with filler yarn/matrix debonding were found to be the primary failure mechanisms for samples loaded in the stuffer direction. Similarly, stuffer yarn/matrix debonding and filler yarn rupture with slight fiber pull-out was identified as the primary failure mechanism when loaded in the filler direction. It was also noted that strain gages would need to be considered carefully prior to use to be sure an adequate representation of the material is captured. Early work by Cox et al. [6] focused primarily on the compressive response of 3D woven angle interlock composites. The findings from this study illustrated that kink band formation in the stuffer tows was the key initiating event that ultimately led to failure. Initial kink bands were found to form near the interaction between misaligned stuffer tows and warp weaver tows. Chou et al. investigated three-dimensional woven composites with reinforcement in three and five directions [7]. Of relevance to the work presented in this paper, the authors examined 12 K orthogonal woven reinforcement with Epikote 828/Epikote 1001 matrix mixture. The tensile response of this material system experienced a material softening phenomenon between 0.5% and 1% tensile strain. Brandt et al. [2] analyzed a novel 3D orthogonal woven composite in tension, bending, interlaminar shear strength, compression, impact and penetration. The study concluded that the new material (orthogonal woven) had greater energy absorption capability and damage tolerance when compared with a comparable two-dimensional woven material. Additionally, the compressive response of the 3D woven was found to be in the same range as two-dimensional woven composites. Brandt et al. conducted a study on the effect of varying the amount of through-thickness reinforcement in multiple 3D woven glass/epoxy composites with respect to tensile, compressive and interlaminar shear strength [8]. The authors found that tensile properties decrease with increasing reinforcement in the through-thickness direction. Compressive strength was found to be greatest when 4% of the fibers were located in the throughthickness direction. When subjected to peel evaluation, the 3D
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materials exhibited improved fracture mechanics over the evaluated 2D architecture. One of the conclusions from this shift in fracture mechanics is an increase in energy absorption capacity and damage tolerance in the 3D materials. Dow Tactix 138 and Shell 1895 resins combined with AS4 carbon and S-2 glass reinforcement were used by Cox et al. [9] to evaluate 3D woven composites. Both orthogonal woven and angle interlock materials were evaluated. The majority of the experimental data reported was for single samples from various materials. Determination of engineering properties for each material was secondary to evaluating the relation between the macroscopic failure modes and microstructural failure mechanisms. It was reported that for compressive and tensile results, tows failed as discrete units. That is, when fibers within a single tow fail, the entire tow section fails. Observations from failure modes suggest that some of the 3D materials were able to redistribute stresses to reach much higher stains after a neighboring tow had failed. Bending failure modes were found to be a combination of both the previously identified compressive and tensile failure modes. Tensile loading was re-examined to capture additional important failure mechanisms [10]. Described in [11,12], layer-to-layer and through-the-thickness interlock woven carbon/epoxy composites with stuffer tow inclusions were investigated. A plastic straightening of the stuffer tows was determined to be the cause of a non-linear softening in the tensile stress–strain response of these composite materials. This straightening was described to occur as misaligned stuffer tow segments began to align with the direction of loading. Stuffer tow rupture and pullout were identified as the primary tensile failure modes. Stuffer tow failure occurred in most if not all stuffers prior to the ultimate failure of the tensile sample for 3D woven interlock composites. Callus et al. performed tensile evaluation and analysis of composites with three-dimensional orthogonal, normal layered interlock and offset layered interlock glass fiber reinforcement [13]. The authors discussed the stress–strain response found in the three materials and noted that, when compared with other experimental investigations in the literature (Cox et al. [10] and Chou et al. [7]), a reduction in tensile modulus was found as tensile strains exceeded 0.4% in both the warp and the weft directions. Using SEM examination, more extensive damage was found in the non-linear region of the tensile response than in the linear region and was attributed to some degree of inelastic tow straightening. This softening response was concluded to be unique to different types of fiber architectures and composite material systems. Leong et al. investigated the response of two similar three-dimensional orthogonal woven composites [14]. Stuffer tow pull out was noted as the primary failure mechanism for the composite when loaded in the warp direction. Extensive longitudinal tow splitting occurred just prior to complete failure. This phenomenon was thought be in part caused by extensive longitudinal matrix cracking from a Poisson’s ratio mismatch between the matrix and longitudinal tows. The compressive response of two orthogonal 3-axis woven composites using 24 K stuffer reinforcement with 12 K x and y (thickness and width) dry tows was examined by Kuo et al. [15]. The first 3-axis material was manufactured with carbon/epoxy stuffer rods and the second with dry tows. For both architectures, the progression of compressive damage was investigated. Kink bands within the stuffer rods and the stuffer tows were identified to be on either the ‘miniscopic’ (tow level) or microscopic (intratow) level. Kuo et al. more recently investigated three different 3-axis orthogonal composites with varying surface patterns in compression [16]. The surface yarn pattern (surface loop density) was found to influence the damage growth within the materials. A
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higher loop density was found to be desirable because it generally resulted in less weaving-induced tow distortion. Stig and Hallstrom evaluated the mechanical performance of a fully interlaced three-dimensional woven composite material [17]. Woven with 12 K T700 carbon tows and manufactured with vacuum injection molding of vinyl ester resin, this composite material experienced a tensile softening response similar to other 3D woven materials. Occurring between 0.5% and 1.0% strain, the softening is described to be indicative of a progressive tensile failure. Damage was found to be more localized in the threedimensional woven material when compared with a twill weave. Mouritz and Cox presented a comparison between stitched, pinned and 3D woven composites [18] from published data. The study mentions that the diameter of woven z-binders, stitches or pins is usually in the 0.1–1.0 mm range, citing that thicker z-binders may result in harmful distortion of the fiber architecture. It is observed that the strength and stiffness of 3D woven and stitched composites and the amount of through-thickness reinforcement are not strongly correlated. This suggests that by taking some of the otherwise in-plane fibers and orienting them through the thickness, benefits of through-thickness reinforced can be exploited without a dramatic sacrifice of in-plane properties. Bogdanovich et al. [19] performed an experimental study focused on joining 3D orthogonal woven composites. Through co-cured and adhesively bonded joints of three-dimensional Eglass preforms, this study concluded that stitching and stapling co-cured preforms prior to infusion greatly increased the singlelap joint strength. Including approximately 2% z-fibers by weight in the preform suppressed delamination during joint failure. Bogdanovich et al. [20] and Lomov et al. [21] have identified damage events during tensile loading of three-dimensional orthogonal woven carbon composites. These events correspond to certain acoustic emission event energy and signal peak frequency. The identified damage events were, in order of occurrence, sporadic cracking of boundary yarns, small transverse matrix cracks and massive transverse matrix cracking combined with tow/matrix debonding at interface boundaries. This paper is focused on exploring the following research objectives: 1. The three-dimensional woven composites in this paper have no biased (off-axis) reinforcement. The goal of this paper is to benchmark the performance of these 3D woven composites with a competitive industry standard two-dimensional woven material with bias reinforcement. 2. Identify material nonlinearities and make design recommendations based on these findings. 3. Characterization of failure modes and fracture paths experienced by the selected three-dimensional woven architectures. 4. Quantify the amount of crimp seen in three-dimensional ply to ply architectures and draw conclusions relating crimp to material properties. 5. Collect reliable data on various three-dimensional woven architectures for use in numerical modeling and analysis. 2. Fiber architectures Four different fiber architectures were examined during this study: a 3D orthogonal preform, two different 3D ply to ply angle interlock preforms, and a traditional two-dimensional weave. Geometric representations of general three-dimensional woven architectures are seen in Fig. 1. All materials were woven at Albany Engineered Composites in Rochester, NH, and molded with a resin transfer molding (RTM) process. IM7 carbon was used in all materials. Tow sizes were either 12 K or 24 K and are explicitly
defined for each material in subsequent sections. Toughened epoxy resins used were either ST-15 or PR-520. ST-15 has a tensile strength of 89.6 MPa and a tensile modulus of 3.31 GPa [22]. Likewise, the PR-520 resin system has a tensile strength of 82.1 MPa and a tensile modulus of 4.0 GPa [23]. IM7 carbon has a tensile strength of 5.6 GPa and a modulus of 276 GPa [24]. The two resin systems have similar physical and mechanical properties and the choice for their use reflects the current trend in industrial aerospace applications. Crimp of a tow is a textile description of the waviness of the tow based on the straightened length as a percentage of the unstretched length. As the waviness of a tow increases, the crimp is said to increase. To quantify crimp in the two ply to ply architectures, measurements were taken along cross sections from digital microscopy images. Crimp is related to tow waviness which is defined in this paper as the amplitude of a tow in a given direction divided by the wavelength of that tow. The tow waviness ratio is the tow waviness in the weft tows divided by the tow waviness in the warp tows. Tow waviness has a significant contribution to the mechanical performance of composites with three-dimensional woven ply to ply fiber reinforcement. 2.1. 3D orthogonal A three-dimensional orthogonal fiber architecture was examined. This preform used 24 K filler tows and 12 K warp weaver and warp stuffer tows. The average total fiber volume fraction for this composite was found to be 51.9%. The preform was molded to a nominal thickness of 4.19 mm with ST-15 resin. Four warp stuffer tows and five weft filler tows are woven as a preform connected by warp weaver tows that travel over two top surface picks and under the next two bottom surface picks. Based on weaving predictions, the warp weaver tows are estimated to be 8.7% of the total architecture by weight. A digital microscopy image illustrates the cross-section of this 3D orthogonal material in Fig. 2. From representative cross sections for this material, a unit cell size of repeating geometry was found to be 10.5 mm in the warp direction and 8.50 mm in the weft direction. The stuffer tows in this fiber architecture in Fig. 2b are notably straighter than the filler tows in Fig. 2a. The inherent waviness of the filler tows is in part a result of the warp weaver tow distorting the outer filler tows during the weaving process. 2.2. 3D ply to ply 12 K/24 K The second three-dimensional fiber architecture subjected to laboratory investigation used 12 K warp tows and 24 K weft tows arranged in a ply to ply angle interlock preform. The preform was molded to a nominal thickness of 4.19 mm with ST-15 resin, resulting in a fiber volume fraction of 50.4%. Fig. 3 depicts cross sections of the fiber architecture showing weft tows (a) and warp tows (b). The amount of crimp in the weft tows is notably greater than the warp tows. Measurements of the unit cell size were made from these representative cross sections and were found to be 10.5 mm in the warp direction and 8.36 mm in the weft direction. The warp tow wavelength was measured to be 10.5 mm with an amplitude of 0.34 mm. The weft tow wavelength was measured to be 8.36 mm with an amplitude of 0.55 mm. With these measurements, the resulting warp tow waviness is 0.0324 and the weft tow waviness is 0.0657. The tow waviness ratio for this architecture is 2.03. 2.3. 3D ply to ply 24 K/24 K The last three-dimensional fiber architecture evaluated was a ply to ply angle interlock with 24 K tows in both the warp and
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(a)
(b)
Fig. 1. Geometric representations of general three-dimensional architectures including (a) orthogonal and (b) ply to ply angle interlock. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
the weft directions. Molded with PR-520 toughened epoxy resin, the nominal thickness of this material was also 4.19 mm. The volume fraction of this material was 59.0%. Similar to the 12 K/24 K ply to ply architecture, the warp tows shown in Fig. 4b are found to have less crimp and are therefore
straighter than the weft tows seen in Fig. 4a. From representative cross-sections, a unit cell size was found to be 10.5 mm in the warp direction and 8.80 mm in the weft direction. The warp tow wavelength was 10.5 mm with an amplitude of 0.383 mm. The weft tow wavelength was 8.80 mm with an amplitude of 0.515 mm. The
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Fig. 2. 3D orthogonal digital microscopy image of fiber architecture. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
warp tow waviness and weft tow waviness are 0.0365 and 0.0585, respectively, resulting in a tow waviness ratio of 1.60 for this 3D architecture.
in-plane material properties and improves out-of-plane properties (e.g., inter-laminar shear strength, and impact toughness [18]). 3. Experimental method
2.4. 2D woven 3.1. Sample preparation In addition to three-dimensional preforms, a two-dimensional woven fiber architecture was also investigated. Eight layers of [0/ 90] were molded using the same RTM process described for the 3D fiber architectures. A [45, 0, 45, 90] symmetric layup was utilized, with layer orientations corresponding to the warp fiber orientations, to represent a quasi-isotropic material (in-plane). Each layer was woven from 12 K warp, 12 K weft tows in a plain weave. The nominal thickness was 3.30 mm. The fiber volume fraction was found to be 49.5%. This conventional two-dimensional woven configuration has no through-the-thickness fiber reinforcement and is therefore susceptible to interlaminar failure modes including delamination. 2.5. Fiber volume fraction When making comparisons between two-dimensional and three-dimensional composites of the same fiber volume fraction, one must consider the relative amount of in-plane fiber reinforcement. Two-dimensional composites maintain roughly all fibers inplane; traditional composites have no through-the-thickness fiber reinforcement. A significant percentage of the fibers within the 3D woven architectures are oriented through the thickness. Taking some of these fibers out of plane in 3D architectures reduces
To ensure the highest sample quality and process repeatability, all materials were cut using a waterjet with garnet abrasive. After samples were cut to the appropriate nominal dimensions, they were conditioned at 23 °C, 50% relative humidity for at least 24 h prior to measuring sample dimensions. 3.2. Digital image correlation With the relatively large unit cell size characteristic of the evaluated 3D woven materials, traditional strain gages have been found to exhibit large variation in measured strain. Studies focused on 3D woven materials have found that strain gauges need to be carefully considered to be sure that the data captured is representative of the material [5]. For this reason, all experimental work utilized the ARAMIS system for digital image correlation for accurate measurement of full field strains. Digital image correlation utilizes a specialized paint spatter pattern on the surface of samples to track points on that surface through time. Of particular interest is the size of the point tracked in space, called facets, and the separation distance between each facet, called the step. A facet is a collection of pixels in a neighborhood that represent a discrete point in a set of points called the
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Fig. 3. 3D ply to ply 12 K/24 K digital microscopy image of fiber architecture. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
field. By first establishing a strain-free state, a deformation gradient is initialized and updated as the camera system accurately tracks the movement of facets through a calibrated control volume in space. As the step between facets becomes smaller, the amount of information extracted from a sample in space increases. Previous work has identified a facet size of 15 pixels and a step of 13 pixels to produce accurate results for similar control volumes [25,26]; all strain fields were analyzed with these settings. 4. Procedure Laboratory procedures are described in this section for tensile, compressive and in-plane shear evaluation. Samples from each material were evaluated with loading in the warp direction and the weft direction for tension and compression. Additionally, samples were subjected to in-plane shear loading. A minimum of five samples were considered for each data set. Standard laboratory atmosphere, 23 °C, 50% relative humidity, was maintained within the laboratory during all mechanical evaluation. 4.1. Tension The procedure for evaluating the tensile response of the four materials was in accordance with ASTM D3039, Tensile Properties of Polymer Matrix Composite Materials. Experiments were run with displacement control at a rate of 1.27 mm/min on a 500 kN Instron servo-hydraulic frame. A 250 kN load cell was used with 250 kN hydraulic wedge grips. Samples were centered between the wedge grips prior to gripping at 20 MPa. The ARAMIS system for DIC was calibrated with a
100 mm 80 mm 80 mm control volume. The nominal width of tensile samples measured 25.4 mm. The corresponding gage area is dependent upon the actual material thickness as listed previously. A macro script was created within the ARAMIS system to ensure that the area in the gage used for calculation of strain was held constant from one sample to another. This calculation area measured approximately 19 mm wide and 100 mm long. Using this method, numerous unit cells were captured for an average strain calculation. The tensile modulus of elasticity was calculated by assuming a linear elastic region between 1000 and 3000 microstrain. 4.2. Compression The combined loading compression (CLC) fixture was used to evaluate materials in compression in accordance with ASTM D6641. Experiments were conducted in displacement control. The standard stipulates that any given experiment should last between 1 and 10 min to be considered quasi-static. At the recommended initial load head displacement rate of 1.27 mm/min, experiments did not last for one minute, prompting an adjustment of the displacement rate allowing experiments to last between 2 and 3 min, corresponding to a rate of 0.254 mm/min. The nominal width of compressive samples measured 12.7 mm. The gage area is dependent upon each of the material thicknesses as listed previously. Evaluation of the four materials was performed in both the longitudinal (warp loading) and the transverse (weft loading) directions. The ARAMIS system was calibrated to a 35 mm 28 mm 28 mm control volume for full field strain analysis. A 100 kN Instron frame was used with a 100 kN load cell.
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Fig. 4. 3D ply to ply 24 K/24 K digital microscopy image of fiber architecture. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
Plates were used in the grip wedges to create flat surfaces for transmitting loads to the CLC fixture. Calculation of compressive strains for each sample using the ARAMIS system involved creating a macro script for repeatable selection of the calculation area. The calculation area was selected to be approximately 10 mm 10 mm, centered in the gage, to include as much of the gage area as possible. The compressive modulus of elasticity was calculated by assuming the linear elastic region to be between 1000 and 3000 microstrain. The sample thickness was sufficient to prevent buckling of the material during compressive loading. 4.3. In-plane shear Qin et al. [27] evaluated a three-dimensional woven material with an orthogonal architecture. Digital image correlation was used with the Iosipescu shear fixture. The investigation found that, given the small unit cell size of the particular orthogonal architecture, measuring approximately 2.4 mm, the use of the Iosipescu shear fixture was appropriate and allowed for multiple repeating unit cells along the gage length. The unit cell of the 24 K warp, 24 K weft 3D woven ply to ply architecture in this paper is roughly 9.8 mm in the warp direction and 8.1 mm in the weft. The other three-dimensional materials have unit cells of similar sizes. To capture the response of three-dimensional preforms woven with 12 K and 24 K tows, a gage area that captures more repeating geometries within a material is preferable. The gage length used with the Rail Shear Fixture from ASTM D7078 nominally measures 30.5 mm, whereas the gage length of samples utilizing the Iosipescu shear fixture is 12 mm. In-plane shear response was therefore evaluated in accordance
with ASTM D7078, Shear Properties of Composite Materials by VNotched Rail Shear Method. The gage areas for each material are dependent upon their respective thicknesses. Samples were placed in the Rail Shear Fixture and bolts in the fixture were torqued 54 Nm while the sample was held square to the fixture and centered. The fixture was placed in a 100 kN Instron frame with a 25 kN load cell. The ARAMIS system was used for shear strain measurement with a 65 mm 52 mm 52 mm calibrated control volume. A macro script was created to generate consistent calculation areas for shear angle from the ARAMIS strain field. The calculation area used for all samples evaluated for inplane shear response was 6 mm wide (perpendicular to the gage length) and 29 mm long, centered between the notches.
5. Failure modes 5.1. Tension Three-dimensional woven ply to ply composites demonstrate tensile failure mechanisms that are sensitive to the amount of crimp in the warp direction relative to the weft direction. When evaluated in the warp direction, warp tow rupture and withdrawal from surrounding weft tows was evident. Failure planes were found to be repeatable between numerous samples and located between the weft tow/matrix interfaces. The fracture path began on one surface of the sample and propagated directly through the thickness, and normal to the direction of loading. Matrix cracking was also observed on the surface of the samples, adjacent to the warp tows and parallel to loading. Representative failure modes for both ply to ply architectures, 12 K/24 K and 24 K/24 K, are seen in
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Fig. 5. 3D ply to ply failure modes for (a) 12 K/24 K architecture and (b) 24 K/24 K architecture. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
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Fig. 6. 3D ply to ply inclined failure mode for (a) 12 K/24 K and (b) 24 K/24 K fiber architectures. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
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Fig. 5a and b, respectively. This observed failure mode was consistent between both ply to ply architectures when loaded in the warp direction. Evaluation with loading in the weft direction revealed different failure modes. For the ply to ply architecture with 12 K tows in the warp and 24 K tows in the weft direction, the failure path is inclined with respect to the weft tows. A representative failed sample is seen in Fig. 6a. The strain to failure of the composite was found to be 1.46% with a 5.54% coefficient of variation (COV) when loaded in the warp direction and 2.13% (14.8% COV) when loaded in the weft direction. This discrepancy in strain to failure can be partially explained by the amount of crimp in the weft tows vs. the warp tows. Weft tows within the studied ply to ply architectures have more crimp than warp tows. The crimp ratios for the 12 K/24 K and 24 K/24 K three-dimensional architectures described previously are 2.03 and 1.60, respectively. This characteristic is thought to allow the material to experience higher strain to failure when loaded in the weft direction as the weft tows begin to straighten. This failure mode was seen in all samples of this architecture subjected to the same experimental conditions. The trend in the strain to failure for the 24 K/24 K ply to ply architecture is similar to that of the 12 K/24 K ply to ply architecture in that the strain to failure when loaded in the warp direction was 1.85% (2.86% COV) as compared with 2.67% (6.24% COV) when loaded in the weft direction. The observed failure path, however, was found to also be inclined with respect to the weft tows but in the opposite direction, as seen in Fig. 6b. This discrepancy in a
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matrix-dominated onset of ultimate failure between two distinctly different three-dimensional woven ply to ply architectures may partially be caused by the different tow sizes found in the warp and the weft directions as well as the mismatch in matrix between the architectures. Fig. 7 illustrates the typical stress–strain responses observed for both ply to ply fiber architectures. A bilinear response or material softening can be seen in each of the curves. The trend is more pronounced for each of the two materials when loading is oriented in the weft direction. Weft samples experience not only lower strengths, but also reduced linear elastic properties and greater strain to failure. The trend in weft and warp loaded bilinearity for both materials can be in part attributed to each of the materials having more crimp in the weft direction. The crimp in a tow tends to reduce when the tow is subjected to strain. This reduction imparts strain on the surrounding matrix. The greater the amount of crimp found in a tow, the stronger this tendency is. The onset of this phenomenon in warp tows (similar to ‘plastic tow straightening’ seen by others [7,10,13]) results in localized matrix damage, corresponding to the transition region between the linear-elastic region and the second region seen in Fig. 7. Unlike the work by Cox et al. [10], the evaluated ply to ply architectures did not include any stuffer tows in the warp direction and therefore the associated tow straightening is experienced in the weaver tows. Fig. 8 shows the tensile strains developed in the loading direction for the 24 K/24 K ply to ply architecture for loading in both the warp (8a and b) and weft directions (8c and d). Each of these
Fig. 7. Tensile stress–strain response for 3D ply to ply architectures loaded in the warp and the weft directions. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
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direction from the warp to the through-thickness direction. A typical tensile failure can be seen in Fig. 9a, with a magnification highlighting stuffer tow and warp weaver rupture in Fig. 9b. When subjected to loading in the weft direction, a strain to failure of 1.43% (7.14% COV) was experienced. Filler tow rupture was found to be the primary failure mechanism, with matrix cracking along the warp weavers and parallel to the stuffer tows. This suggests further agreement with Tan et al. [5] where a weft loaded orthogonal architecture was also found to fail adjacent to stuffer tows. 5.2. Compression
Fig. 8. Tensile strain in the loading (vertical) direction for 3D ply to ply 24 K/24 K sample loaded in the warp direction (a and b) and weft direction (c and d). Strain fields (a) and (c) correspond to a state just prior to the bilinear transition region whereas strain fields (b) and (d) correspond to a strain state just after this transition region. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
figures represent strain fields from just prior to the bilinear transition region (8a and c) and just after the transition (8b and d). Areas of high surface strains were found in the matrix material with notably higher strain concentrations in the weft loaded direction. This localized strain and corresponding damage experienced by the matrix is a contributor to the bilinear stress–strain behavior seen in both ply to ply architectures. Observed failure modes in the 3D orthogonal material coincide with a previous study by Tan et al. [5]. Warp loaded tension samples were found to exhibit a linear stress–strain response to failure. With a 1.61% (4.50% COV) strain to failure when loaded in the warp stuffer direction, fracture planes were identified along the filler tows accompanied by matrix/filler tow debonding. Additionally, stuffer tow withdrawal was seen with rupture found in the stuffer tows. Leong et al. noted a similar stuffer tow withdrawal that was accompanied by longitudinal (stuffer tow) cracking prior to failure [14]. Longitudinal stuffer tow cracking was not found in any of the evaluated samples in this experimental investigation. Matrix cracking was present in the surrounding composite parallel and adjacent to the filler tows. The warp weaver tows were found to have ruptured near the global failure when the tow changes
Different failure modes were observed in the warp and the weft directions when subjected to compressive loading conditions for the 3D orthogonal fiber architecture. A typical sample failure for warp-direction loading is seen in Fig. 10a. Warp stuffer tows are seen to be relatively straight running in the warp (left–right) direction. This straightness is a result of the weaving process for this orthogonal preform. When on the loom, the warp stuffer tows are kept in tension throughout the weaving process. The high degree of stuffer tow alignment has led to a mild brooming-type failure mixed with a through-thickness angled failure caused by macroscopic compressive stuffer tow failures and kink band formation. Unlike warp stuffer tows, weft filler tows are not kept in tension during the weaving process for this preform. Fig. 10b depicts a weft-loaded failure of the same material. It is seen that the weft filler tows, now running from the left to the right in the figure, do not maintain the same level of fiber straightness through the gage area of the sample, leading to an angled through-thickness failure from fiber microbuckling. The stress–strain response from this orthogonal woven material was essentially linear until failure for both the warp and weft loaded cases. Callus et al. [13] noted a softening in the stress–strain response for both warp and weft loaded compression of an orthogonal woven composite that was not seen in these samples. Compressive failure modes were found to be largely a concurrent debonding between the matrix and the surrounding tows coupled with tow microbuckling and isolated kink band formation for both the 3D ply to ply architectures. Failure modes and fracture paths found in samples loaded in the weft direction were similar to those found in warp direction loading. Typical warp-loaded sample failures are found in Fig. 11a and b for the 12 K/24 K ply to ply and 24 K/24 K ply to ply architectures, respectively. For comparison, a typical 2D woven quasi-isotropic failure is seen in Fig. 11c. Kink band formation as a primary compressive failure mode was found to coincide with previous work by Kuo et al. [15] and Cox et al. [6]. 5.3. In-plane shear In-plane shear samples were oriented in the warp/weft plane with the warp tows perpendicular to the gage length as defined between the notches in the samples. Failure modes observed during shear evaluation were consistent between both 3D ply to ply architectures. Shear strains were found to be greatest in the matrix material following the interface between one weft tow and the adjacent weft tow and were isolated to the gage area between the two notches. Shear loads are initially carried by the warp and weft tows in the ply to ply architectures. As shear strains build in the matrix it begins to degrade. This degradation is thought to transfer the initial shear loads from the tows to the weakening matrix. Concurrently, the weft tows begin to angle with increased shear loading. The combination of these two observed failure mechanisms leads to the softening seen in the shear stress–strain response in Fig. 12.
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Fig. 9. Typical (a) 3D orthogonal tensile failure in warp loading with (b) magnification of 3D orthogonal stuffer tow and warp weaver tow rupture. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
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Fig. 10. Typical (a) longitudinal and (b) transverse 3D orthogonal compressive failures. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
The 3D orthogonal woven architecture experienced a similar response and failure mechanism to the two 3D woven ply to ply architectures. Shear angles were once again found to be isolated to the area between the notches. Seen in Fig. 13 is a shear angle field distribution of the orthogonal architecture in a representative sample prior to failure. Highlighted as areas of greater shear strain are the matrix pockets found between filler tows along the gauge. The discontinuities in these matrix rich regions between filler tows are warp weaver tows on the surface of the material. Similar to the ply to ply architecture, significant matrix cracking is observed in the regions just described. 6. Results and discussion 6.1. Tension Tensile experimental results are presented in Table 1. Poisson ratios are also presented as calculated in the linear elastic region used for modulus calculation. The 3D orthogonal woven material
outperformed all other evaluated materials in tension for both strength and tensile modulus, with the exception of the two-dimensional woven material in transverse loading. The 3D orthogonal preform is made from five warp stuffer tows woven to four weft filler tows with orthogonal through-thickness reinforcement. The reduction of strength and stiffness seen in weft vs. warp loading conditions is attributed to the mismatch in inplane fibers in the warp and the weft directions. The weft-loaded modulus was found to be 84.9% of the warp-loaded modulus, reflecting the 4:5 weft to warp tow ratio. Weft loaded strength was found to be 61.7% of the warp loaded strength. Deviations from the 4:5 weft to warp tow ratio are likely a result of tow waviness and the addition of warp weaver tows. The three-dimensional ply to ply architectures, when compared with one another, reveal that the 24 K warp, 24 K weft preform has greater strength and stiffness for both warp and weft direction loading conditions. The transverse (weft-loaded) tensile strength of the 24 K/24 K ply to ply material is 54% of the warp loaded strength. For the 12 K warp, 24 K weft ply to ply architecture, the
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Fig. 11. Typical longitudinal compressive failure for (a) ply to ply 12 K/24 K, (b) ply to ply 24 K/24 K and (c) quasi-isotropic 2D woven fiber architectures. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
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Fig. 12. Shear stress–strain response for 3D architectures. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
tensile strength when loaded in the weft direction is 46% of the tensile strength when loaded in the warp direction. This suggests that despite volume fraction differences between architectures, a mismatch of 12 K warp, 24 K weft tows combined with greater crimp in the weft tows yields a proportionally lower strength when comparing the weft strength to the warp strength. Although the resin systems used for each of these two 3D architectures were different, this is thought to have only a minor contribution to the mechanical properties reported. Trends in material nonlinearity were common between both material systems. 6.2. Compression Compressive results are presented in Table 2. Of the 3D fiber architectures examined, the orthogonal woven was found to have the highest strengths when loaded in both the warp and the weft directions. This is attributed to the straightness of the warp stuffer and weft filler tows. Both ply to ply architectures lack the inclusion of ‘straight’ fibers within the respective preforms, resulting in noticeably lower compressive strengths. A trend seen in the compressive behavior for each of the 3D architectures is lower strength and stiffness when loaded in the weft direction as compared with the warp direction. This trend can once again be explained by the increase in crimp in the weft direction as compared with crimp in the warp direction for both ply to ply architectures. This induced waviness transfers stresses from tows to the matrix, resulting in the observed lower strengths and stiffnesses.
Compressive failures seen in ply to ply architectures with debonding between matrix and surrounding tows oriented in the loading direction were isolated when compared with two-dimensional woven failures. Two-dimensional failure modes were observed to contain extensive material brooming protruding beyond the bounds of the original thickness. Through-thickness reinforcement found in the 3D woven architectures contains the failure to closely within the original thickness. By doing this, certain failure modes are minimized or eliminated, including delamination. The ply to ply architecture with 24 K warp, 24 K weft tows has a higher volume fraction than the other materials, which likely contributes to the increase in strength and stiffness from the other ply to ply fiber architecture. Additionally, the 12 K/24 K ply to ply architecture has 27% more tow waviness than the 24 K/24 K ply to ply architecture. Compressive strength and stiffness from weft loading normalized to the compressive strength and stiffness from warp loading is one possible comparison to evaluate the performance of a material when subjected to a range of loading conditions. Normalized strengths and stiffnesses can also be used as a measure of the three-dimensional weaving process with regards to weaving-induced crimp and other tow nonlinearities. A visual comparison between all evaluated materials is seen in Fig. 14. The ply to ply 24 K/24 K architecture has greater relative stiffness ratios from a weft-loaded to warp-loaded orientation in compression as compared with the 12 K/24 K ply to ply architecture. The strength ratio,
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Fig. 13. Shear angle along the gauge area in the matrix for the 3D orthogonal fiber architecture. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
Table 1 Experimental tensile results. Fiber architecture
Tensile strength MPa (COV) 0°
3D 3D 3D 2D
orthogonal ply to ply 12 K/24 K ply to ply 24 K/24 K woven
1150 703 855 765
Tensile modulus GPa (COV)
90°
(4.83%) (7.45%) (7.82%) (3.41%)
710 323 461 717
0°
(7.81%) (3.44%) (3.84%) (3.79%)
69.6 53.4 62.5 49.9
m12 (COV)
90°
(2.61%) (5.56%) (2.88%) (6.06%)
59.1 33.0 43.0 46.6
(2.08%) (3.78%) (2.61%) (4.97%)
0.066 0.257 0.244 0.330
(13.8%) (4.93%) (11.3%) (5.15%)
Table 2 Experimental compressive results. Fiber architecture
Compressive strength MPa (COV) 0°
3D 3D 3D 2D
orthogonal ply to ply 12 K/24 K ply to ply 24 K/24 K woven
494 270 301 483
(3.23%) (6.15%) (7.07%) (8.02%)
however, was found to be slightly smaller in the 24 K/24 K ply to ply material as compared with the 12 K/24 K material. The compressive weft to warp stiffness ratio for the 3D orthogonal fiber architecture was found to be greater than all other relative weft to warp stiffness or strength ratios for all evaluated materials. 6.3. In-plane shear The examined architectures have no biased plies; there is no fiber reinforcement found in the three-dimensional woven
Compressive modulus GPa (COV)
90° 328 203 217 435
(2.81%) (3.78%) (6.76%) (3.33%)
0° 55.6 47.9 59.3 45.5
(10.3%) (7.58%) (6.67%) (4.98%)
90° 53.3 31.2 41.2 41.4
(7.68%) (2.20%) (4.16%) (2.08%)
materials outside of the orthogonal warp, weft and through-thickness directions. An observed nonlinearity and softening in the shear stress–strain response was found in all 3D architectures. This nonlinearity is not specific to 3D woven architectures. Plain woven [0/90] carbon composites of similar fiber volume fraction have exhibited similar trends in shear stress–strain nonlinearity [28]. The primary difference seen in 3D woven composites, including the three architectures examined in this paper, is that the strain to failure is much greater for 3D woven composites than for 2D. Experimental in-plane shear results are presented in
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Fig. 14. Weft-loaded compressive strength and stiffness normalized to compressive warp-loaded strength and stiffness. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
Table 3 Experimental shear results. Fiber architecture 3D orthogonal 3D ply to ply 12 K/24 K 3D ply to ply 24 K/24 K
Shear strength MPa (COV)
Shear modulus GPa (COV)
84.8 (3.07%) 89.6 (3.29%) 106 (3.68%)
3.94 (4.68%) 4.10 (5.07%) 5.00 (4.16%)
Table 3. The increased shear performance seen in the 24 K warp, 24 K weft ply to ply architecture over the 12 K/24 K can be partially attributed to the higher fiber volume fraction found in the 24 K/ 24 K material. 7. Conclusions Comprehensive mechanical evaluation of four composite materials has been discussed. Composites molded with three different 3D fiber architectures were evaluated and compared with a baseline 2D woven composite. Tensile, compressive and in-plane shear properties were examined. One of the main benefits of a composite molded from a 3D woven preform is that the fiber architecture can be tailored to achieve specific types of performance based on the application. Comparing the response of an orthogonal woven composite to a ply to ply woven composite demonstrates this benefit. Recognizing that other three dimensional woven
architectures are possible, this paper highlights only three specific architectures. As a result of the experimental investigation presented, the most important conclusions are: 1. Material nonlinearity is seen in the tensile response of 3D woven composites with ply to ply architectures when loaded in the weft direction. This nonlinearity has been explained and is attributed to the degree of crimp in the warp direction relative to the weft direction. Nonlinearity is also observed in orthogonal architectures under shear and select tensile loading. As a result of these significant nonlinearities, strain-based design should be considered with 3D woven composites. 2. Failure mechanisms for 3D woven composites in tension, compression and in-plane shear have been identified and characterized. One common initial failure mechanism witnessed in all evaluated samples was debonding of the matrix from tow inclusions for both resin systems. 3. The ratios between compressive warp-loaded strength and stiffness normalized to compressive weft-loaded strength and stiffness have been presented. This architecture-dependent relationship was used to compare the effects of crimp in compressive loading for each 3D woven material. Similar comparisons can be made in samples loaded in tension. The weaving process can be tailored to adjust the relative crimp between the warp tows and the weft tows. If a similar material response
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between warp loading and weft loading is desired, it is recommended that tension be controlled in weft tows during the weaving process to result in a tow waviness ratio closer to 1.0. 4. A reliable set of quasi-static data from three novel 3D woven architectures is now available. This data can be utilized for modeling of 3D woven materials as well as benchmarking of new architectures to these existing materials that are actively used in the aerospace industry.
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