Voids in fiber-reinforced polymer composites: A review on their formation,characteristics,and effects on mechanical performance

Formation of voids

LCM is a process for manufacturing of thermoset composites through liquid resin infiltration into fibrous preforms. The term covers a large group of processing techniques, of which the most well-known ones are resin transfer molding (RTM) and vacuum-assisted resin transfer molding (VARTM). In LCM processes, there are several causes for void formation such as mechanical air entrapment during resin flow (main cause), ¹⁴ gas created due to chemical reactions during cure, ¹⁵ and nucleation of dissolved gases in the resin. ¹⁶ The air entrapment is mainly due to the inhomogeneous fiber architecture, resulting in non-uniform permeability of the fiber preform, which causes local variation in resin velocity. This local velocity variation is exacerbated by the capillary effect, prevailing at the micro-scale. ¹⁷

Owing to the multi-scale nature of FRCs, the voids can be formed in three different scales: macro, meso, and micro. Micro-voids are formed in between the fibers in a tow, meso-voids in between the tows, and macro-voids in a larger zone of the preform (observable with the naked eye). Micro-voids are smaller than meso-voids in transverse cross-section, while are usually larger in planar view (Figure 1(a)). Micro- and meso-voids formation is controlled by the micro-scale flow at the tow level relating to the heterogeneous medium of the preform, whereas macro-voids form with regard to the macroscopic (global) flow considering the preform as a homogeneous medium. The macroscopic and micro-scale flows are strongly coupled, interacting with each other. ¹¹ OPEN IN VIEWER

In the literature, the voids with the same scale are sometimes differently called. For instance, another name for macro-void is “dry spot”, for meso-void is “inter-bundle”, “inter-tow”, or “channel” voids, and for micro-void is “intra-bundle”, “intra-tow”, or “tow” voids. ¹¹ Moreover, meso-voids are sometimes incorrectly called macro-voids. In the present review, the terminology introduced in the beginning of this paragraph is used. Furthermore, terms “pore” and “porosity” that are used in many studies to refer to voids are changed to “void” and “voidage”, respectively, in order to ensure consistency. “Pore” is used here to specify the empty regions in the reinforcement structure, before impregnation and “porosity” refers to these before-impregnation pores.

Formation of macro-voids: A macro-void is a zone that has not been impregnated by the macroscopic resin flow yet, while the resin flow front has reached the vent. ¹⁹ This type of void, observable even with the naked eye, is induced by distorted flow patterns, due to irregular permeability of the preform, improper injection locations, ²⁰ or the presence of inserts, ribs, and cores in the mold. ²¹ The formation of the macro-voids is often modeled by a conventional mold-filling approach for a saturated flow, where a moving distinct edge (the flow front) divides the domain into two regions: saturated and unfilled. ²² This simulation approach is based on substituting the Darcy’s law ²³

v ¯ = - K ∼ μ · ∇ P

(1)

∇ · v ¯ = 0

(2)

∇ · ( K ∼ μ · ∇ P ) = 0

(3)

Liu et al. ²¹ presented a simulation feature in the Liquid Injection Molding Simulation (LIMS) software tool ²⁵ to numerically predict the location, size, and pressure of all dry spots formed during the filling process, by which optimal processing conditions for a controlled injection (with regard to dry spots) can be achieved. Lee et al. ²⁰ performed a numerical simulation, based on the mass–momentum balance for modeling the macroscopic flow, to investigate a control methodology to avoid dry spots. This was performed along with a real-time control scheme, detecting the flow front with optical sensors, to obtain experimental results for validation of the model. These simulation studies suggest methodologies to achieve optimal processing conditions to minimize the formation of macro-voids.

Formation of micro- and meso-voids: Mechanisms: Although the LCM process is considered to be successful for production of parts with no visible dry spots, the impregnation of fibers inside tows can be incomplete, resulting in un-wetted fibers, potentially causing small-scale, i.e. micro- and meso-voids. ²⁶ These voids show more complicated physics in their formation and evolution, compared to the macro-voids. Therefore, a large amount of research has been performed in this area, suggesting a dual-scale framework for the study of voids in LCM. This dual-scale voidage is characteristic for tow-based composites (short term for textile composites made of yarns, excluding unidirectional (UD) plies, random mats, and short fiber composites). It correlates with two distinct scales of gaps to be filled with the resin flow: the large gaps (meso-pores) between fiber tows and small channels (micro-pores) inside tows between individual fibers (Figure 1(a)). In fact, the fiber preform can be considered as two interacting and superimposed continuous porous media, one containing meso-pores and the other containing micro-pores. ²⁷ According to numerical investigations by Tahir et al., ²⁸ the macroscopic permeability of the dual-scale fibrous preforms is governed by meso-pores.

Between the tows, the hydrodynamic force drives the resin viscous flow, but between the fibers, due to the small pore diameter, the surface tension becomes significant, and thus the wicking flow due to capillary pressure becomes a dominant factor to drive the resin flow. Formation of meso- and micro-voids is controlled by the competition between the viscous flow and the capillary flow. If the capillary flow between fibers dominates the flow, the meso-pores remain unfilled, and meso-voids form between tows (Figure 1(a) and (c)). On the other hand, if the hydrodynamic flow is faster, the micro-pores remain unfilled, and micro-voids form inside tows (Figure 1(a) and (b)). ¹¹ This competition is quantified by capillary number ( Ca ), which is a dimensionless velocity, representing the ratio of the viscous force to the capillary force. Patel et al. ²⁹ proposed a modified capillary number ( Ca * ) that takes the liquid–fiber–air contact angle into account

Ca * = μ V γ cos θ

(4)

The macroscopic (global) resin flow can occur in two directions relative to the fiber direction. Through analytical modeling, it has been shown that when the flow is perpendicular to the fiber direction (transverse flow, Figure 1(a)), voids are distributed throughout the composite, ⁴⁰ whereas in case of flow parallel to the fiber direction (longitudinal or axial flow, Figure 1(a)), void formation is localized at the flow front region. ⁴¹ For the parallel flow, void localization at the flow front was also experimentally observed as significantly higher void content close to the ventilation side of the laminate. ¹⁴ Some exotic angles of resin flow are also studied in literature, e.g. in the literature,^42,43 which are explained further in the current section.

Various methodologies have been employed to perform in situ visualization of void formation at the flow front, during resin impregnation. They include high-resolution video-assisted microscopy, ³⁰ video recording, ⁴⁴ combination of image analysis and visualization techniques, ⁴⁵ infrared thermography, ⁴⁶ voltage sensor system, ³⁷ and X-ray micro-computed tomography (micro-CT). ⁴⁷ Rohatgi et al. ³⁰ observed that for the same capillary number, the micro-void content was higher for transverse flow than for the axial flow. In a multidirectional laminate, the voidage was much lower in layers parallel to the flow, and it increased with increase in the off-axis angle, as observed through micro-CT by Sisodia et al. ⁴⁸

Whether the macroscopic flow is longitudinal or transverse, the impregnation of the tows can take place in two directions: parallel (straight arrows in Figure 1(a)) and transverse (inclined arrows in Figure 1(a)) to the macroscopic flow direction. Using an analytical model, Binetruy et al. ⁴⁴ concluded that for the longitudinal macroscopic flow, the tow impregnation happens mainly transverse rather than parallel to the flow.

An experimental methodology, based on capillary rise monitored by fluorescence visualization, was developed by LeBel et al. ⁴⁹ to identify the capillary properties of fibrous preforms as well as the penetrativity of resins (an intrinsic property of the fluid and of the fluid–fiber interface), which were considered as the key parameters responsible of void formation by air entrapment. This technique was employed to determine a window of the optimal modified capillary number, and hence an optimal flow front velocity. They also claimed that the capillary rise analysis is more suitable for exploring the impregnation mechanisms in fibrous preforms than a separate study of dynamic viscosity, surface tension, or contact angle between fiber and resin.

Formation of micro- and meso-voids: Theory: In modeling the formation of micro- and meso-voids, unlike of macro-voids, the assumption of a distinct edge for the flow front (saturated flow) is not valid since, due to the capillary effect, there exists a partially saturated (or simply unsaturated) zone at the flow front that separates the fully impregnated zone and the unimpregnated zone. The direct consequence of this delayed impregnation is the trapped air within the preform when resin has already reached the mold exit. Therefore, a variable, called degree of saturation, i.e. the ratio of the filled volume to the total pore volume in a unit cell, is defined to take into account the delayed impregnation. The mass conservation equation at an arbitrary point reads as ⁵⁰

Ø ∂ s ∂ t + ∇ · v ¯ ( s ) = 0

(5)

v ¯ ( s ) = Ø s v ¯ p ( s ) = - k ( s ) μ K ∼ · ∇ ( P ( s ) - P c ( s ) )

(6)

Ø ∂ s ∂ t = ∇ · ( k ( s ) μ K ∼ · ∇ ( P ( s ) - P c ( s ) ) )

(7)

The difficulty in this approach is that the relative permeability and the capillary pressure should be known. This is done by either experimental measurements or micro-scale simulations. ¹¹ In the global analysis of unsaturated flow, the resin motion in the inter-tow gaps (meso-pores) is described by the Stokes flow. ²² The micro-scale flow, in small channels in between fibers in a tow, can be simulated based on the Stokes flow with distinct flow front, ²² Darcy’s law with coupling between the two flow regions, ²⁵ or Brinkman equation. ¹¹ The capillary rise can be modeled based on the modified Jurin’s law for dual-scale fabrics. ⁴⁹

The key factors for modeling of the impregnation and void formation in LCM processes are the dual-scale nature of the porous medium and its saturation.^11,33,50 This has been investigated for decades. Patel and Lee ²⁷ proposed a phenomenological model based on a modified Darcy’s law and continuity equations inside meso- and micro-pores. Depending on the permeabilities and the capillary pressures of the pores, resin saturation distribution could be calculated with the model, and the fraction of meso- and micro-voids could be determined. Kang et al. ¹⁷ proposed a mathematical model to describe mechanisms of void formation based on a microscopically non-uniform velocity field at the front of a transverse resin flow. The model showed that for a given fiber preform, the combined effect of resin velocity and capillary pressure can be described by the capillary number. The model can predict the size and content of voids within and between fiber tows.

In addition to the modified capillary number, fabric microstructure parameters such as global and tow fiber volume fractions and macroscopic and tow permeability were found to be important in prediction of void content, according to Gueroult et al. ³⁷ Moreover, the fabric shear angle and flow direction greatly influence the size of meso-voids, as revealed by the mathematical model based on the analysis of the resin flow velocities inside and outside fiber tows by Yang et al. ⁴³

The effect of geometrical anisotropy, arising from the difference in geometry of fiber bundles in the warp and weft directions, on void formation in woven fabrics was investigated by Matsuzaki et al. ⁵¹ They developed a mathematical model for inter-tow (meso) void formation, which showed that optimal flow velocity and preferable flow direction change with geometric anisotropy of the fabrics. The results of this model were in agreement with their experimental results for impregnation of woven glass fabrics with an unsaturated polyester resin in the literature. ⁵² The same authors experimentally studied the effect of resin flow angle relative to the angle of an anisotropic fabric preform on void content. ⁴² They found that with changing the flow angle from 0 to 90°, void content reaches a minimum value, corresponding to the trade-off between inter-tow and intra-tow impregnation. From the observations on the change of impregnation directions with the resin flow angle, they derived an analytical model that can predict void content at any arbitrary resin flow angle as well as the minimum-void angle for a given flow velocity.

Overall, analytical models show that besides capillary pressure and resin viscosity (capillary number), fabric microstructure, fabric shear angle, flow direction, and preform geometrical anisotropy play a role in void formation in LCM. However, in case of high injection pressure, Hu et al. ⁵³ suggested that the capillary effect is negligible. With this assumption, they developed a mathematical model, based on two simplified unit cells of multi-layer woven fabrics, to analyze the formation and size of voids. They concluded that in case of the flow in the weft direction, the void size is a function of the ratio of weft axial permeability and warp transverse permeability.

Formation of micro- and meso-voids: Numerical simulations: Void formation is also simulated numerically. In some early numerical studies, the modeling is based on methods such as Stokes/Brinkman lattice Boltzmann, e.g. Spaid and Phelan. ⁵⁴ Later, numerical studies linked the microscopic capillary effect within tows to the macroscopic resin flow by means of tow saturation. Bréard et al.^55,56 developed a numerical procedure to simulate mold filling, accounting for void formation. Determining the permeability of the fiber preform as a function of saturation, they concluded that the quality of LCM parts in terms of void content can be predicted by modeling the impregnation process that takes into account the dual-scale porous medium and saturation degree.

Parseval et al., ⁵⁷ and Pillai and Advani^58,59 proposed to use a “sink” term in the continuity equation (equation (2)) to account for the delayed impregnation of tows in numerical modeling. The name “sink” is because tows act as sinks of liquid and keep on absorbing the resin even after its front has passed them. ⁶⁰ Simacek and Advani ⁶¹ modeled this sink term by appending extra one-dimensional (1D) elements to the conventional mesh in a finite element (FE)/control volume approach. They implemented the tow saturation model in the simulation package LIMS. This model was later used as a basis in various research studies on void formation for predicting dynamic void content (representing air entrapment and void motion),^33,62 investigating the effect of capillary number, fiber volume fraction, and position along the injection length, ⁶³ simulating different processing methodologies, ⁶⁴ modeling the effect of capillary pressure and air entrapment on fiber tow saturation, etc. ¹⁹

The effect of surface tension (capillarity) and wetting in LCM was studied using a bifluid–solid contact model developed by Liu et al. ⁶⁵ within an FE framework. They concluded that dynamics of the contact line motion needs to be considered for more accurate results. Using the boundary element method, Patiño Arcila et al. ⁶⁶ simulated the impregnation process of the dual-scale fibrous preform by coupling Darcy flow in inter-tow gaps (meso-pores) with Stokes flow in intra-tow channels (micro-pores). This allowed for simultaneous fulfillment of both dynamic and kinematic interface boundary conditions at the moving flow front. The void formation was simulated in a model, where the flow direction-dependent capillary pressure was calculated without experimental factors, and the surface traction effects at the flow front were taken into account. The authors claimed that the shape of the flow front is modeled very accurately in comparison to other techniques, implementing the reconstruction of the moving boundaries. From this approach, they were able to draw some conclusions on void characteristics, namely that voids are formed at the side of the transverse tows (wefts) where the resin leaves the tow, and their aspect ratio increases as the void becomes smaller.

Numerical simulation is also a useful tool for optimization of the flow front velocity to minimize the void content.^20,31 Lundström et al. ⁶⁷ developed a three-dimensional (3D) numerical model for the impregnation of non-crimp fabrics (NCFs), using a network model that mimics biaxial fabrics. With the focus on inter-tow (meso) voids, it was shown that intrinsic perturbation in the bundle geometry is of high importance. DeValve and Pitchumani ³⁶ presented a more detailed numerical simulation of the infiltrating dual-scale resin flow through the ideal architecture of plain weave fibrous preforms accounting for the capillary effects within the fiber bundles. They predicted the air entrapment locations within the fiber preform as well as the size and shape evolution of the voids, for different Capillary and Reynolds numbers, as displayed in Figure 2. OPEN IN VIEWER

Concluding remarks on modeling of the void formation: Macro-voids, or rather dry spots, can be predicted with the state-of-the-art software, based on the solution of the Darcy equation with varying permeability over the preform because of the local preform deformations. These calculations do not ask for a detailed consideration of the preform local internal geometry and need only the local permeability as a function of the local fiber volume fraction and the local fiber direction. Such a link can be provided by analytical/numerical approximations, and the local preform compaction and distortion can be calculated using forming/draping models.

Meso- and micro-voids should be modeled taking into account the preform internal geometry, and the Stokes solution for local flow is needed, probably coupled with the Darcy approximation, in the inter-tow gaps. The key points for the prediction of void formation are modeling of a dual-scale flow, unsaturated flow, and capillary effects. The detailed modeling representing all these phenomena is appearing in the last few years, but this work is far from being finished. Approximations like “sink factor” are useful shortcuts within the existing modeling tools, but need calibration to become predictive. One can expect that in the coming decade, the progress with computational power, the numerical methods in computational fluid dynamics, and the multi-level modeling will bring the full dual-scale free surface models to the off-the-shelf simulation software.

Void compression and dissolution

After voids are created, their size and shape may change before the consolidation is finished. This can happen in two ways: they may be compressed by or dissolved into the resin. In the case of mechanically entrapped voids and no diffusion, the compression behavior of voids can be approximated by the ideal gas law ¹⁴ including surface tension and a capillary effect. ⁶⁸ Therefore, for isothermal conditions, the final volume of voids can be calculated. As pointed out by Park and Lee, ¹¹ two types of error can be introduced in such calculations: for micro-voids inside tows, the air pressure inside the voids does not increase immediately with increase in the ambient resin pressure; resin cannot permeate any further into the tow if the micro-void is already highly compacted by the resin.

In reality, molecules migrate over the void–resin interface, and the dissolution of gas into the resin takes place due to diffusion (in a steady-state situation). Lundström ⁶⁸ investigated the in situ dissolution of cylindrical micro-voids, trapped between fibers, during RTM, and made a qualitative comparison with Fickian-based theoretical analysis. He found that the voids disappear in the order of minutes due to the diffusion. Moreover, not only high pressure but also high (local) flow rate and low initial gas concentration in properly degassed resin improve the dissolution of the trapped voids. However, if initial concentration of gas or water within the resin is high, the voids will grow instead of dissolving into the resin. Matsuzaki et al. ⁴⁵ showed that the molecular diffusion can lead to void growth in VARTM process. The void shrinkage or expansion due to diffusion depends on whether the resin is undersaturated or supersaturated with the gas, respectively. During vacuum-assisted impregnation, resin pressure goes below the atmospheric pressure. Hence, the resin becomes supersaturated, and the gas molecules dissolved in the resin diffuse into the void, causing expansion of the void. Matsuzaki et al. ⁴⁵ explained that the void content changes after void formation: voids can move with the resin flow, escape from tows to inter-tow regions, shrink due to surrounding pressure change, and grow because of the diffusion at the void–resin interface. Each of these steps may be influenced by the resin injection pressure and pressure gradient.

Yamaleev and Mohan ²⁶ argued that diffusion-based models can only predict the voidage in the final stage and cannot be used in evaluation of the initial void dynamics. They proposed a numerical model that includes the liquid/vapor phase transition occurring in the gas mixture entrapped inside tows. It was concluded that the condensation of the gas mixture, during the microscopic impregnation process, reduces the steady-state void size considerably in comparison with that predicted by the models with the ideal gas assumption. In addition, the void formation is much faster if the real gas effects are considered. The void size predicted by such a model can be used as initial conditions in diffusion-based models like the one in the literature. ⁶⁸

Void motion

Only some voids stay where they are formed and the rest move with the resin flow, in particular in axial flows where void formation occurs at the flow front. The mobility of voids can be expressed by the force balance between the drag caused by the pressure gradient across the void and the adhesion due to surface tension. ¹¹ In general, void mobility is characterized by two non-dimensional parameters: the capillary number and the ratio of the void size to the inter-tow or inter-fiber spacing. High resin velocity, which means high capillary number, provides high mobility for voids. ¹¹ There is a critical capillary number, above which the void becomes mobile. ²⁷ However, Kang et al. ⁶⁹ stated that a geometric hindrance force may also affect the transport of the voids, and thus geometric configuration of the mold should be considered in addition to the capillary number. About the influence of void size on its mobility, different arguments have been made.^27,69–71 In general, meso-voids (between tows) may be removed easily if they move with the resin, while micro-voids (inside tows) are more difficult to remove. Rohatgi et al. ³⁰ attributed this to the much larger gaps between the tows than those within the tows, which results in insufficient drag due to the hydrodynamic pressure to make the micro-voids mobile. Because of the difficulty in their mobilization after they are formed, it is essential to minimize formation of the micro-voids, which necessitates a proper choice of the molding conditions. Lundström ⁷⁰ concluded that for removal of micro-voids, mechanisms such as compression and diffusion are more probable to occur than motion.

Frishfelds et al. ⁷² showed that in impregnated NCFs, voids move with the resin through the inter-tow channels, and are trapped if the channels become too narrow. The voids move, on average, biased to the direction of the tows. For the same fabric geometry, Lundström et al. ⁶⁷ showed that intra-tow (micro) voids move much slower. However, it is common for them to become inter-tow (meso) voids in the case of high resin wetting of fibers.

The void mobility is the highest when it is nearest to resin flow fronts according to Gangloff Jr et al. ⁷³ They derived a relationship between the dimensionless void size, process capillary number, and void mobility, based on a simplified model proposed in the literature. ⁷⁴ By aid of synchrotron micro-CT, Vila et al. ⁴⁷ performed an in situ analysis of the motion of voids during tow impregnation. They observed that voids are trapped inside tows when the distance between fibers is below a critical value. Therefore, the micro-flow and mobility of micro-voids between fibers depend on the local microstructure, mainly the local fiber volume fraction, and the presence of convergent/divergent individual fiber trajectories. Additionally, wetting between the resin and fibers and rheological properties of the resin play a significant role. Sisodia et al. ⁴⁸ attributed the higher voidage in off-axis plies of a multidirectional laminate to the more difficult escape of voids with resin flow through the off-axis channels compared to that through the parallel channels with the flow.

Park et al. ³⁵ presented an integrated modeling approach for the void formation, in the meso-pores between fiber tows and in the micro-pores inside fiber tows, as well as compression and transport of voids, which was implemented into an FE code for mold filling simulations. They showed that a partially saturated zone (bubbly zone) forms behind the macroscopic flow front. To obtain a full impregnation, they suggest to perform a bleeding process, in which by keeping the injection pressure high, and air vents open, even after the macroscopic resin front reaches the air vents, the voids in the partially saturated zone are removed.

Concluding remarks on void formation mechanisms in LCM processes, and how to avoid them: In conclusion, to reduce the void content in an LCM product of a given material, the stages of formation, compression/dissolution, and motion of voids should be controlled (Figure 3). As explained above, void formation can be minimized by optimization of the (modified) capillary number. Although according to the model of Labat et al., ⁷⁵ a very high injection (inlet) pressure increases the formation of micro-voids, it is favorable for the other two stages. It can help to reduce the void size due to compression, to dissolve the voids in the resin due to increase in solubility of the gas, and to make the voids mobile due to increase in resin velocity, in particular causing the hydrodynamic pressure to overcome the capillary pressure for micro-voids entrapped inside tows.^11,20 OPEN IN VIEWER

Appropriate application of vacuum also can assist in reduction of voids in LCM. The effect of vacuum is explained by the reduction of air pressure, which favors the compression of the entrapped air,^18,35,67,76 or by compression of the voids when the vacuum is released after injection, ¹⁴ or by high compaction of preform as a result of vacuum. ⁷⁷ Vacuum may have negative effects as well. For instance, it can induce another source of voids, which is evaporation of (certain volatile components in) the resin. ¹⁸ In addition to optimization of the (modified) capillary number, increase in the inlet pressure, and application of vacuum, other treatments, such as degassing the resin prior to infusion^16,68 and continuing the resin flow after the mold filling is complete (bleeding),^35,64,78 are often performed to reduce the void content. Moreover, some other methods to achieve a low void content have been proposed: compressing mold walls during injection, ⁷⁹ applying a permanent post-fill pressure (packing pressure) after injection,^80–82 light waxing of the mold surface with buffing, applying vibration to the mold,^83–85 applying sufficiently high magnetic compressive pressure, etc.^86,87

However, for an accurate evaluation of the influence of processing parameters on the part quality, particularly void formation, optimization of those parameters should be performed in a set, not based on individual parameters, but taking into account any probable interdependency between the parameters. For instance, Kedari et al. ⁷⁷ showed that a combination of strong vacuum and high mold temperature requires a reduced inlet pressure for minimizing the void content, in the case of impregnation of E-glass random mats with polyester. They presented a flow compatibility model to explain that there must be a different optimal inlet/outlet pressure drop of a VARTM process at a different mold temperature.

Autoclave curing process

One of the most common manufacturing techniques used for fabrication of high-performance structural FRCs is the autoclave curing process. This is because of the high-pressure environment in the autoclave, which results in a high fiber volume fraction, and facilitates the dissolution and removal of voids in the part. There are several models proposed for formation and evolution of moisture-based voids during cure of thermoset resins. The curing process was modeled by Loos and Springer, ⁹¹ who related the cure cycle to the thermal, chemical, and physical processes taking place during cure. They explained that after a void is formed, its size might alter due to thermal expansion, diffusion effects, or changes in the void pressure because of the changes in ambient temperature and pressure. Their model accounts for the last two mechanisms. A good agreement between the experimentally measured void contents, as a function of the cure pressure, and the void contents predicted with the Loos and Springer model is found in the literature ⁹² for hot-press processing.

The dissolution and growth of moisture-based voids in the prepreg during cure was described to be diffusion-controlled by Kardos et al. ⁹⁰ They analyzed and modeled these phenomena as a function of temperature and pressure. The stability of voids was described with a pressure–temperature–humidity map, as exhibited in Figure 4, which determines conditions for void growth or dissolution throughout the cure cycle. Later, Boey and Lye ⁹³ modified the diffusion-based model by a unit matrix volume, defined as the volume of resin required to produce a certain void content with a single void of a given diameter, which is a characteristic parameter for the resin. OPEN IN VIEWER

Kardos’ model, however, did not account for surface tension of the resin. This results in an unrealistic increase in the void size with temperature. Wood and Bader ⁹⁴ developed a diffusion model that can predict the rate of growth or collapse of entrapped voids in the resin, with the advantage of accounting for surface tension. They concluded that voids can be collapsed, and their growth can be suppressed by control of pressure and temperature even if the resin is saturated with a gas. Other modifications were made to Kardos’ model by White and Kim ⁹⁵ to model the void growth during staged curing, which is a manufacturing technique applying intermediate partial cure during laying up of thick composites. They took into account the effect of surface tension in obtaining the bubble gas pressure. Moreover, the initial bubble radius was determined by a force balance at the bubble wall. They argued that due to the cooldown in the staged cure, the decrease in partial pressure of gases inside voids causes the voids to collapse if the resin viscosity is sufficiently low.

Ledru et al. ³ argued that Wood and Bader’s model has low sensitivity to the autoclave pressure and proposed a coupled visco-mechanical and diffusion void growth model, which considers surface tension as well as pressure sensitivity. Through this model, it was concluded that there are three key parameters influencing the void size evolution: the onset of pressure application, the concentration of diffusive species, and the diffusion coefficient. They showed that the void size (radius) increases with temperature in the first stage of cure, drops with the application of pressure in the second stage, and is almost constant in the third stage (high pressure and temperature).

According to de Parscau du Plessix et al., ⁹⁶ these models are not able of predicting a reasonable size for the voids (prediction of extremely large radii). Therefore, the authors developed a different model for the prediction of growth of spherical voids, which accounts for the non-Fickian (Dual-Fick) behavior of the epoxy resin as well as the slowdown of water diffusion due to polymerization. The additional slowdown of water diffusion at the interface was explained by the formation of an interphase around the void, having different diffusive properties. For three different cure cycles (Figure 5(a) and (b)), they predicted the evolution of void size, which had a reasonable order of magnitude (radius in the order of tens of micrometers, as shown in Figure 5(c)). However, they argued that the highly complex physics behind the void formation and growth during cure is not yet completely clear for researchers. Furthermore, according to Hernández et al., ⁹⁷ the above models correspond to small spherical voids growing inside polymer, but not to cylindrical voids forming in a viscous matrix in the presence of fibers. OPEN IN VIEWER

With an appropriate cure window, i.e. an optimized set of temperature, vacuum pressure, and autoclave (cure) pressure profiles versus time, large parts can be fabricated with almost no voids, because the high pressure largely suppresses the evolution of volatiles and dissolved species and eliminates any entrapped air. ⁸⁸ The removal of intra-laminar voids by means of vacuum is more difficult since they are not connected and also the through-thickness air-permeability of prepregs is low; therefore, the void evacuation is done mainly through inter-laminar zones. ⁹⁸ Stone and Clarke ⁹⁹ manipulated the autoclave pressure and the vacuum-bag internal pressure in order to obtain specimens with a variable void content for their analysis. The vacuum level in composite manufacturing is around 0.5 bar.

Shim et al. ¹⁰⁰ found that without vacuum and in the presence of autoclave pressure, a high inter-laminar voidage exists, but the composite made with vacuum and without autoclave pressure has the highest voidage. Hence, autoclave pressure plays the consolidating role, suppressing residual air growth simultaneously. Similarly, Liu et al. ¹⁰¹ concluded that an autoclave pressure is imperative to reduce the void content down to an acceptable level. Boey and Lye ⁹³ showed that without the application of vacuum, void content can be reduced by increasing the cure pressure, but even with high pressures, a small degree of voidage is present. Furthermore, they showed that in the absence of vacuum, with increase in cure temperature, the void content increases. Campbell et al. ¹⁰² explained this by the increase in the volatile vapor pressure when temperature is increased. They also mentioned that factors such as prepreg surface roughness, laying up environment, laminate thickness, ply orientation, internal ply drop-offs, tooling, and bagging play a role in void formation during autoclave processing of prepregs. According to Olivier et al., ¹⁰³ the size, shape, and distribution of voids alter with cure cycle parameters. They determined optimal curing pressure conditions to minimize void content based on the results of thermogravimetry and mechanical spectrometry tests.

The influence of residual solvent content in a prepreg, controlled by the drying time prior to cure, on the void formation was investigated by Naganuma et al. ¹⁰⁴ They observed that in the case of short times of prepreg drying, open voidage (to the thickness surface) with large and irregularly shaped voids is formed around the bundle intersections, while in the case of long drying times, closed voidage with flat and semicircular voids is formed in the spaces between plies.

Koushyar at al. ¹⁰⁵ explored the effects of variation in autoclave parameters on voidage in a composite panel and observed that maintaining the vacuum throughout the cure cycle results in a high void content. The latter was found to increase with increase in the cure pressure. They explained this by boiling volatiles in the resin when vacuum is maintained at elevated curing temperatures. Li et al. ¹⁰⁶ showed that when the cure pressure is low, a considerable number of large voids mostly at ply interfaces are formed, of which the number and size decline significantly when the pressure is increased. For low cure pressures and in the presence of vacuum, Kakakasery et al. ¹⁰⁷ observed that debulking, which is removal of air and volatiles by application of vacuum at a rather high temperature prior to cure, could significantly reduce the intra-laminar voids. Curing at high pressures drastically diminished the void content and standard deviation of the measurements.

Concluding remarks on voids in autoclave curing processes: The role of autoclave pressure is crucial in suppressing the void formation. For selecting the suitable profiles for the cure window parameters, i.e. pressure, temperature, and vacuum, they need to be optimized together, rather than separately. There are also other factors, including residual solvent content and humidity, which influence the void formation in autoclave processing. Techniques such as debulking and prepreg drying can be employed to further reduce the voidage in final parts.

OoA curing process

With the composite market growth and considering high costs of autoclave processing of large parts, OoA manufacturing techniques have shown potential over the past two decades to meet future demand. In particular, autoclave-quality parts have been manufactured through vacuum-bag-only (VBO) consolidation. ¹³ Low capital investment, lower consumable costs, and improved energy efficiency can be counted as advantages of OoA, in addition to higher production rates. ¹⁰⁸ They are due to the possibility of using a diverse range of lower cost cure setups, such as conventional ovens, heating blankets, heated tooling, ¹³ and hot presses. ⁹⁷

In contrast to the autoclave process, in which alleviation of voids is easily done by means of high pressure, the void formation and mitigation are serious issues in the OoA process. Although modifications to the prepregs for OoA manufacturing are needed (explained in the following paragraph), appropriate processing conditions, especially sophisticated thermal cycles, are essential to achieve a high consolidation and low void contents. ¹³ OoA production of large parts with void contents less than 2% is reported in the literature; some examples are reviewed in the literature. ¹³ The current section is devoted to OoA curing of prepregs, while other non-autoclave techniques dealing with LCM such as RTM and VARTM are already discussed in the Liquid composite molding section.

VBO consolidation: VBO is an OoA technique that only applies vacuum along with high-temperature initial- and post-cure to consolidate laminates. Voids in OoA, in particular in VBO composites, can be classified into “flow-induced” voids, resulting from insufficient impregnation of fibers before resin gelation, and “gas-induced” voids, due to the presence of entrapped air, moisture, or resin volatiles. In an unconsolidated stack of plies, voidage can exist in three forms: micro-voids within dry tow cores, meso-voids within the resin-rich regions around tows, and inter-laminar voids between the plies. During VBO, when vacuum is applied, air is evacuated through the dry tow cores, and the content of meso- and inter-laminar (gas-induced) voids reduces. Then, when temperature is increased, resin flows into the dry tow areas, causing the content of flow-induced voids to decrease (Figure 6). It can be concluded that to get rid of the gas-induced voids during VBO consolidation, surprisingly, the introduction of an open-cell porosity into the prepreg, which facilitates the air evacuation, is required. ¹³ This can be achieved by partial impregnation of the preform during the prepreg production, which makes the VBO prepregs “breathable” through the intra-tow voids, as illustrated in Figure 7. Thus, the degree of impregnation in a prepreg becomes a prominent factor determining the void content in the final part made of this prepreg.^13,109–112 OPEN IN VIEWER

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Figure 7.

Schematic and micrograph of a “breathable” prepreg consisting air evacuation channels suitable for VBO – EVaC stands for “Engineered Vacuum Channels”. ¹³

Nevertheless, partial impregnation can be present in the final laminate in the form of flow-induced micro-voids inside tows if the material and thermal conditions are such that the curing reaction finishes before complete impregnation of tows takes place. According to Centea and Hubert, ¹¹³ this can occur due to a high initial degree of cure (induced for example by high out-time ¹¹² ), fiber-dense tows, slow heating rates, low dwell temperatures, as well as low impregnation time. ¹¹⁴ Therefore, to produce a void-free laminate, in addition to the partial impregnation of the as-received prepreg (special prepreg for VBO processing), which is a requirement for sufficient air evacuation, a proper rheological evolution of resin during VBO is essential. Resin viscosity must evolve during VBO such that at each stage, favorable viscosity for void elimination at that stage is acquired. This means a high viscosity during vacuum application to retain the intra-tow voids as paths for evacuation of gas-induced voids, and a low viscosity during heating to allow resin flow into the intra-tow voids, removing flow-induced voids. This is schematically shown in Figure 6. This proper evolution of the resin viscosity can be achieved through optimization of material parameters and process conditions. ¹³

The ability of vacuum to reduce gas-induced voids and that of heating to decrease flow-induced voids were reported by Centea and Hubert. ¹¹⁵ They performed micro-CT on laminates processed to different stages of a simple cure cycle. After an hour-long room temperature vacuum hold, the number and size of the meso- and inter-laminar voids decreased dramatically, whereas the micro-voids tended to be reduced through progressive impregnation only once the material was heated. From this work, the relationship between process conditions and voidage in VBO was perceived and it was further investigated in a later study, ¹¹³ where high ramp rates and isotherm temperatures were found to rapidly complete the tow impregnation, which may prevent the removal of meso- and inter-laminar voids. On the other hand, if ramp rates and isotherm temperatures become very low, tow impregnation might be stopped due to gelation leaving micro-voids in the tows. In an earlier study, Davies et al. ¹¹⁶ used Quickstep cure ^a as an OoA technique (but not VBO) and showed that higher process ramp rates in Quickstep processes can reduce resin viscosity, thus facilitating void removal. In particular, the consolidation of the laminate was improved with a lower void content due to maintaining a lowered resin viscosity for a long period. In addition to the application of vacuum for evacuation of the gas-induced (trapped) voids and providing sufficient heat, pressure, and time for removal of the flow-induced voids, Fernlund et al. ¹¹⁷ suggested an extra intermediate stage necessary for control of void formation in OoA: Heating the laminate up to cure while keeping the resin pressure high enough to maintain the volatiles dissolved until resin gelation.

Low-voidage carbon/polyether ether ketone (PEEK) laminates was manufactured through oven vacuum-bag processing by Zhang et al. ¹¹⁸ Using a resin infusion void filling model, ¹¹⁹ they could estimate the time required to achieve zero voidage. They underlined the importance of the gas diffusion at the void–resin interface in predicting the final void content. Moreover, the material variability within prepregs is suggested by Helmus et al. ¹²⁰ to have a significant influence on void formation in OoA prepregs. They proposed a stochastic method to model the material variability, which can be used to simulate, more accurately, the void formation during OoA consolidation.

The change in the air evacuation capacity with time, during VBO cure, can be measured by the change in through-thickness air permeability, which was performed by Tavares et al. ¹²¹ According to their results, through-thickness air permeability of prepregs, being influenced by progressive prepreg impregnation and evolution of resin viscosity, changes by several orders of magnitude during VBO cure. This means that also the potential for removing gas-induced voids changes during VBO cure. Characterizing the prepreg surface roughness and the evolution of the surface contact during OoA cure, Helmus et al. ¹²² suggested that the entrapped inter-laminar air can either be evacuated from the surface roughness valleys, move inside the prepreg and then exit through the tow evacuation channels, or migrate in the out-of-plane direction. Therefore, fiber distribution and in-plane and out-of-plane air permeability play a role in evacuation of the entrapped inter-laminar air.

So far, the formation and removal of voids, mainly caused by air entrapment either inside tows or between plies during VBO cure, have been discussed. VBO suffers from another source of voids, which is moisture in the resin. Grunenfelder and Nutt ⁸⁸ showed that the high pressure in autoclave is sufficient to suppress void formation even for prepregs with high level of moisture absorbed, whereas in VBO processing, void volume fraction increases exponentially as a function of moisture content (Figure 8). They validated their results against the modified diffusion-based model of Boey and Lye. ⁹³ Kay et al. ¹²³ confirmed that moisture and entrapped air are the dominant sources of voids within VBO prepregs, while cure-arising volatiles rarely become a source of voids. OPEN IN VIEWER

Using mass spectrometry, Agius et al. ¹⁰⁸ proved that the main volatile within a prepreg is water, and traces of other volatiles such as alcohol, acetone, and ethanol are very small. If the water vapor pressure exceeds the hydrostatic resin pressure during cure, the moisture-rich voids can grow, which happened only during the second stage of the recommended cure cycle, in their case. Later, they ¹²⁴ found that a conditioning procedure (120 min at 40℃ under −97 kPa absolute pressure in a vacuum oven) prior to laying up and cure could effectively reduce the moisture and solvent volatiles within the prepreg. They showed that the conditioned composite had consistently lower void content throughout the VBO cure, without any visible void growth. In general, the absence of high pressure in VBO processes makes it vulnerable to the process deviations that reduce the void-suppression capacity. ¹³

Besides resin moisture, another undesirable process deviation may arise from decreased resin pressure (due to reduced ambient pressure) ¹²⁵ or increased void pressure because of the poor vacuum,^123,125 and insufficient air evacuation as discussed above. ¹²⁵ Moreover, geometric complexity ¹²⁶ and large size of the part ¹²³ can aggravate void control in VBO processing. Techniques such as applying resonance by a pneumatic vibrator to the curing system ¹²⁷ are proposed for reduction of void content in OoA processes.

Hot press curing: Hot press manufacturing, as another OoA technique, is studied with regard to void formation, as well. Employing a hot press with an identical constant pressure, but different temperature cycles, Hernández et al.^97,128 concluded that the final void content depends on the actual evolution of the resin dynamic viscosity throughout the cure cycle. Through micro-CT, they observed that initial voidage in an unconsolidated prepreg was limited and mainly concentrated within the tows, hence intra-laminar. In the cured laminate, however, most of the voidage, being elongated in the fiber direction, was inter-laminar resulting from the air entrapment during laying up. This represents the case that partially impregnated tows exist in the as-received prepreg, but have been fully impregnated before the gas-induced voids are removed through the unimpregnated tow cores. They figured out that a low voidage in OoA products could be achieved through a proper temperature cycle, which results in a wide processing window, in which the resin viscosity stays low. Moreover, the resin flow in these works^97,128 was anisotropic and mainly occurred along the fiber direction, in agreement with the higher permeability factor in this direction. The dominant resin flow in the fiber direction caused a channel-type structure, and enabled the transport and coalescence of voids along the fibers. ⁹⁷ As a consequence of a fiber-rich skeleton in the composite (explained in the Location and spatial distribution section), the pressure distribution during consolidation becomes inhomogeneous, which may influence the volume fraction, size, and location of the voids. ¹²⁸ Observing a rise in the void content with restricting the flow region, Anderson and Altan ¹²⁹ found the resin outflow as the dominant mechanism for evacuation of voids during hot-press cure at low pressures.

The combined effect of plate pressure and prepreg humidity exposure on void formation was investigated by Anderson and Altan. ¹³⁰ They observed that at low processing pressures, the composite with high moisture content have the highest void content. However, the effect of moisture was diminished with increase in pressure such that at high pressures, void content tend to reach an asymptotic value of ∼1.6%, independent of moisture content. They also found that the moisture absorption/desorption of the prepreg can be described by a 1D Fickian diffusion model. A model for the prediction of void content based on the processing pressure and moisture content was developed, being able to predict the non-zero asymptotic value. In a similar study by Gu et al., ¹³¹ an asymptotic value of ∼0% was obtained. Moreover, increase in gel temperature was found to increase the void content, at a given pressure and moisture content. This was attributed to higher void growth at higher temperatures.

Concluding remarks on voids in OoA curing processes: Void formation in VBO processes is a serious issue due to low (vacuum) pressure during the process. Qualitatively, the formation and evacuation of the voids are well understood, but methods for quantitative prediction and design of the process is yet to appear due to complexity of the phenomena involved. For flow-induced voids, these phenomena relate to change of the resin viscosity with advancement of the gelation, for gas-induced voids – complex diffusion of air, resin volatiles, and moisture.

Automated prepreg laying

As stated in the beginning of the Prepreg technology section, modern prepregs possess a very low volatile and moisture content. Therefore, the final voidage is predominantly controlled by the air entrapment during laying and consolidation pressure applied to collapse the entrapped voids. ¹³² The automated prepreg laying involves a unique stacking process and sometimes in situ consolidation, and hence it is interesting to investigate this technique from the voidage point of view. Automated tape laying (ATL) and automated fiber placement (AFP) are the two main technologies of automated prepreg laying, used today to manufacture advanced composites. The main difference between ATL and AFP is in the width of the prepreg they use. The tapes used in ATL are much wider than those used in AFP. Thus, AFP can deal with more complexity in the part geometry, while ATL is suitable for production of flat or low-curvature parts. ¹² Both thermoset and thermoplastic automated lay-ups may be consolidated in an autoclave or OoA (for example, using in situ consolidation during the laying process), depending on the materials and manufacturer. For thermoplastic prepregs, the application of AFP is more common than ATL. ¹² Moreover, automated tow/tape placement (ATP), a technique in which consolidation already happens during the tow/tape placement, is used to lay up thermoplastic prepregs. This technique can significantly reduce manufacturing costs benefiting from in situ consolidation. ¹³³ It is worth noting that the void formation in the automated lay-ups can be influenced by the presence of defects in the tow laying, caused by not-that-precisely-laid tows, namely gaps (matrix pockets between tows) and overlaps (zones of increased local fiber volume fraction on the overlapping tow boundaries).

Regarding the voidage in thermoset automated lay-ups, Lukaszewicz and Potter ⁹⁸ argued that variability of the uncured prepreg in terms of initial voidage, resin content, and surface roughness determines the final voidage in the cured composite. Therefore, they measured these three parameters in two different uncured prepregs. Inside the prepregs before cure, a high voidage was found, which decreased toward the center of the prepreg cross-section. The void content did not correlate with the void size. Similarly to the void distribution, the average resin content decreased toward the middle of the uncured prepreg. This similarity may be linked to the prepreg manufacturing method. Regions with a high resin content are probable to have higher voidage. One of the prepreg types had a surprisingly high surface roughness, which is more likely to trap inter-laminar voidage, in particular in large parts. Overall, it could be deduced that the variability in the unconsolidated prepreg is too high to allow simple development of robust analytical models for predicting final voidage in these materials.

Later, Lukaszewicz et al. ¹³² showed that high-quality carbon/epoxy laminates, with respect to voidage, can be produced from autoclave prepregs without application of debulking and additional pressure, but with sufficient compaction of each ply during the automated laying process at elevated pressure and temperature and using VBO (OoA) curing. They developed an automated laying simulator to provide proper operating parameters. By compaction of each ply during laying at high temperatures, surface roughness of prepregs is reduced, decreasing the inter-laminar sites for air entrapment. At very high laying temperatures (∼70℃), much of the intra-laminar voidage in the uncured prepreg is also removed, achieving a void content of ∼2%, as can be qualitatively observed in Figure 9. OPEN IN VIEWER

For automated laying of thermoplastic prepregs, usually ATP is used. ATP can include the following steps: heating, development of intimate contact between overlying tows and removal of inter-laminar voids, interfacial healing, consolidation and squeeze flow due to the compaction roller force (reducing the intra-laminar voids), void growth resulting from high temperature, and polymer degradation. ¹³⁴ The quality of ATP composites is assessed mainly with the interfacial healing and voidage in the final part. Pitchumani et al. ¹³⁴ modeled the void growth as the expansion of air in a quiescent polymer melt at certain pressure and temperature, and void consolidation as a squeeze flow of a compressible fiber–resin–voids mixture under the compaction rollers. They concluded that as the consolidation force exerted by the rollers increases, the interfacial bonding increases and the final voidage reduces. When the heat input is increased, the void growth is promoted and voidage is increased although interfacial bonding is improved. However, if a forced cooling is applied after consolidation, the void growth is suppressed, leading to a reduced final voidage.

Tierney and Gillespie ¹³³ determined the through-thickness void distribution in ATP parts based on variations in processing conditions by means of a macroscopic flow model and a microscopic void dynamics model, coupled with the heat transfer solution. Large gradients of internal void content were found through the thickness, attributed to the preheating of the composite surface (tacking approach) and repeated compaction (multiple passes) during the laying process. The model showed that these multiple passes drastically reduce the final void content. The deconsolidation, i.e. the void growth in high-temperature low-viscosity matrix, is somewhat controlled within the laminate, but it increases the void content in the laminate surface. As a solution, the authors proposed quenching of the material surface directly after the consolidation step. Lamontia et al. ¹³⁵ explained that the temperature, force, and contact time of the compaction rollers at the laying head play an important role in controlling voidage through resin flow. However, they argued that high-quality in situ thermoplastic ATP anyway needs flat tapes with low initial voidage since tape roughness hampers the layer-to-layer intimate contact, leaving behind inter-laminar voids. The removal of initial voids is also hardly possible since the through-thickness escape route is eliminated by the compaction force, leaving intra-laminar voids in the laminate. Therefore, the application of flat tapes with low voidage was found to be a more practical remedy than the use of low placement velocities or large compaction forces. Through a non-local void filling model, Simacek et al. ¹¹⁹ showed that the predicted void content, in addition to the initial void content, depends on the initial void distribution along the prepreg width. Their model describes the void dynamics in thermoplastic prepregs during the tape placement, and accounts for the volatile pressure in voids, compaction load, fiber bed response, and resin pressure due to squeeze flow.

In contrast to the autoclave processing of semi-crystalline thermoplastic composites, the in situ OoA consolidation processes of these composites have a narrow processing window where elevated temperature and pressure are applied. Therefore, macroscopic resin flow, reptation of polymer chains across ply interfaces, and numerous void reduction mechanisms (e.g. coalescence, migration, compression, bubbling), which take place usually in autoclave, happen to a smaller extent in the in situ consolidation processes. Consequently, prepreg variables such as void content, level of crystallinity, dimensional tolerance, and fiber volume fraction become the crucial factors determining the quality of the semi-crystalline thermoplastic composites produced by ATP and in situ consolidation. In addition, the heat source, which can be hot-gas, infra-red, or laser, plays a key role. ¹³⁶ Comer et al. ¹³⁶ observed through micro-CT that the voids, in an unconsolidated carbon/PEEK prepreg, are highly elongated in the fiber direction (Figure 10(a) and (b)), as found in the literature. ¹¹⁸ They also found that voidage increased through laser-assisted ATP (an OoA consolidation process, which couples an elevated temperature with high dynamic shear forces exerted by the compaction roller), whereas it reduced during autoclave processing (Figure 10(c) and (d)). This could be due to the addition of inter-laminar voids, trapped during laying, which are controlled by prepreg roughness and processing parameters. Moreover, rebound of intra-laminar voids because of inadequate extracted heat could be the reason for the increase in the void content. OPEN IN VIEWER

UD-ply composites ^b

Void morphology in composites built up from UD prepregs has been studied the most. Stone and Clarke ⁹⁹ reported a change in the void morphology with an increase in the void content, from small and spherical (volatile-induced) voids between plies to inter-laminar (air-entrapment) voids, which are “flattened and elongated”. Hsu and Uhl ¹⁴⁷ carried out one of the first systematic studies on void morphology in UD-ply composites. They conducted a “manual” tomography by cutting consecutive thin slices perpendicular and parallel to the fiber direction, and performing microscopy and image analysis on them. It was observed that voids, which were mainly inter-laminar, do not have a circular cross-section, but mainly flat irregular cross-sections. The average height-to-width ratio of the voids was 0.35 (void size parameters are illustrated in Figure 14). Most of them were elongated in the adjacent fiber direction, and voids with large cross-sections were quite long (“needle-like voids”). The authors reported that there is a large length-to-width or length-to-height ratio (respectively ∼15 and ∼50) for inter-laminar voids between plies with the same as well as different orientation, which seems to be common for different volume fractions and manufacturing processes. OPEN IN VIEWER

Gürdal et al. ¹⁴⁴ observed that intra-laminar voids in each UD ply are also elongated mainly in the fiber direction. Their microscopy and image analysis results showed the voids have “cylindrical morphology and elliptic cross-section”, with an average length/radius ratio of ∼40. Rubin and Jerina ²²⁰ stated that inter-laminar voids in UD laminates have “airfoil cross-section” and run along the fibers. Using the manual tomography method, Olivier et al. ¹⁰³ observed that no void developments had happened in perpendicular direction to fibers; larger voids were present mainly between the plies (inter-laminar), and the voids were in contact with the fibers, which caused local change in the distribution of surrounding fibers. The elongated morphology of voids in UD-ply composites is reported in many other studies and expressed with different terminology such as “elliptically cylindrical”, ⁸⁹ “cylindrical inter-laminar”, ²²¹ “cigar-shaped”,^222,223 “large and long”, ²²⁴ “flat and elliptical in cross-section”, ¹⁴⁶ “flattened”, ¹⁰¹ “rod-like shape”,^97,128 and “cylindrically shaped”. ²⁰³

The change in the void morphology with the void content in UD-ply composites, reported by Stone and Clarke, ⁹⁹ was endorsed by Stamopoulos et al. ¹⁹⁸ through micro-CT analysis of voids, where the authors showed that the shape of the voids changed from “spherical” or “ellipsoidal” to a “needle-shape”, when the void content increased. This was related to the curing pressure by Liu et al. ¹⁰¹ and Guo et al. ²²⁵ They concluded that for high curing pressures (low void content), only small spherical voids in resin-rich regions may exist, while for low pressures (high void content), much larger, flattened, and elongated voids with elliptical or different irregular cross-section exist mostly at the ply interfaces. Li et al. ¹⁰⁶ affirmed that “globular” (using their terminology) voids mainly result from resin volatiles and “columnar” shape voids are the result of gas entrapment during fiber impregnation or laying up.

For quasi-UD flax FRCs, Li et al. ²²⁶ observed that for low hot-press pressures, large and semi-spherical voids were mostly situated between the flax yarns and at the interface between fiber and matrix, whereas for high pressures, more microscopic voids were located inside the yarns. It was reported by Liebig et al.^218,227 that intra-laminar voids between carbon fibers have a high length-to-width ratio (“wormhole” void), while inter-laminar voids with an irregular shape (“cumulus-clouds” voids) could be observed in resin-rich regions between plies. The typical morphologies of voids in UD-ply composites are schematically exhibited in Figure 14(a).

In a precise analysis of void morphology in UD laminates produced in a hot press with the identical constant pressure, but four different temperature profiles, Hernández et al., ¹²⁸ using micro-CT, observed that voids in the final part were mainly inter-laminar with a rod-like shape, resulting from air entrapment between the plies during laying up. They fitted each void into an equivalent elliptic cylinder with the same volume, centroid, and moments of inertia as those of the void (Figure 15(b)) and found that the maximum misalignment between the major axis of the cylinders and fiber direction was below ∼1.5°. Flatness ratio (f), i.e. the ratio between the transversal axes of the elliptical cross-section, was ∼1.5, regardless of the cure cycle (Figure 15(d)). They pointed out a dominant effect of the applied pressure on the cross-sectional shape of the voids. The elongation factor, i.e. the ratio between the major axis and the average transversal axis, increases with the volume of individual voids, however, with a deceasing increase rate. In other words, the larger the void, the longer it becomes. This affirms the existence of two different void sources: small voids, with more rounded shape, originate mainly from the prepreg itself, whereas large and elongated voids arise from air entrapment between the prepregs during laying up. In a later study, Hernández et al. ⁹⁷ observed that in a quasi-isotropic (QI) laminate, voids are predominantly elongated in the fiber direction of their (adjacent) ply (Figure 16). Fitting them to an equivalent ellipsoid, it was observed that cure cycle did not change the morphology (which was determined by elongation and flatness factors) of the voids, but the laminate stacking sequence did. OPEN IN VIEWER

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Figure 16.

Void density (number of voids per mm ³ ) versus the orientation of the major axis of the equivalent ellipsoid in the quasi-isotropic carbon/epoxy composites, manufactured through the three different cure cycles shown in Figure 15(c). The graph indicates that the voids are predominantly elongated in the fiber direction of their (adjacent) ply. ⁹⁷

Little et al. ¹³⁸ also performed void characterization by means of micro-CT, confirming the observations made by Hernández et al., ¹²⁸ but measuring sphericity of voids, which is the inverse of their elongation. They observed that sphericity almost linearly decreased with an increase in the projected area diameter of the void, which indicates that a void becomes more elongated as its volume increases. In the absence of possibilities to evaluate the accurate void morphology, a similar conclusion was made, but with regard to the total volume of voids, the axial void length increases with the void content increase.^89,147,224

Non-UD-ply composites

In the studies on void morphology in woven-prepreg composites, more variability is noticed. Madra et al. ¹⁴¹ detected three different void types in a glass/polypropylene (PP) woven composite: small spheres in the resin, usually between two yarns; long voids inside the yarns; irregular large voids in the resin at tow corners, usually in conglomerates. The main morphologies observed and reported for voids in woven fabrics are (1) narrow and elongated in fiber direction inside tows, (2) irregular or spherical in resin-rich regions at tow intersections, (3) flat pancake-shape between the overlaying tows of two adjacent plies, and (4) small spheres in resin-rich regions. These voids can be seen in Figure 13(a) and are schematically depicted in Figure 14(b). A more detailed description of the corresponding studies is presented in Table 3, in which the results on void morphology in other types of composites are summarized as well.

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Table 3.

Void characteristics reported in the literature for different composites.

Manufacturing type	Ply structure	Material	Morphology	Location	Size	References
Prepreg	UD	Void shape (morphology) section	Mainly cylindrical with elliptic cross-section, more details in the Void shape (morphology) section	Mainly at ply interfaces, more details in the Void shape (morphology) and Location and spatial distribution sections	Void size section (Table 4)	Void shape (morphology) section
	Woven	Carbon/epoxy Carbon/PI	Small and spherical voids	Primarily in resin-rich pockets	–	89
		Carbon/epoxy (8-harness)	A truncated cone with a narrow elliptical cross-section	Between plies	Void length mainly: 0.15–0.70 mm	222
		Carbon/epoxy (8-harness)	Flat elliptical voids with patterns depending on the reinforcement architecture	Inter-ply	Average void length: 0.48 mm	220
		Carbon/epoxy Carbon/PI	Spherical, in carbon/epoxy composites some larger voids with elliptical cross-sections as a result of coalescence of a few smaller voids	Localized in the resin-rich pockets at the cross-over points	Average diameter in carbon/PI composite: 30–80 µm	146
		Carbon/epoxy (5-harness)	Pancake-shaped voids, very wide due to low degree of intermingling of tows in the prepregs	Between fabric plies	Width: 100 µm to 9 mm thickness: 25 to 300 µm	215
		Carbon/epoxy Carbon/bismaleimide (8-harness)	Carbon fabric/epoxy: triangular voids → Carbon fabric/bismaleimide: elongated voids →	Preferentially at the crossing of the fiber tows At the interface of the fiber tows	– –	155
		Carbon/epoxy	Large voids → Spherical voids →	Between the plies, mostly in contact with fibers Inside plies	Equivalent diameter: mostly between 3.6−50.0 µm aspect ratio (l/w): mostly <10	217
		Glass/PI (8-harness)	Open voidage (to the thickness surface) with large and irregularly shaped cross-section voids → Closed voidage with flat and semicircular shape voids →	Located around the warp and weft intersections (intra-laminar) In the space between prepregs in the composite (inter-laminar)	– –	104
		Carbon/epoxy (5-harness)	Quasi-elliptically shaped voids → Very wide and flat voids →	Inside tows (intra-laminar) Between plies, in the inter-tow gaps	– –	115
		Carbon/epoxy	Spherical and small voids → Large and elongated voids →	In resin-rich regions At the ply interface, in contact with fibers	Void length mainly: <500 µm aspect ratio (l/w): mostly <10	228
		Carbon/epoxy	Spherical → Elongated or elliptical → Triangular shape →	Preferentially located at the resin-rich areas In the resin along the fiber/matrix interface At the crossing of the woven fiber tows, within the resin-rich areas	– – –	229
		Carbon/epoxy	Narrow and elongated in fiber direction → Small spheres → Flat (pancake-shape) →	Intra-laminar voids in tows In resin-rich regions Inter-laminar voids	– – –	151
		Glass/PP	Spherical → Cylindrical → Irregular →	in the resin, usually between two yarns in the middle of the yarn in the resin, usually in conglomerates	Diameter: 5–20 µm Diameter: 5–30 µm Diameter: 25–75 µm	141
LCM	UD	Glass/epoxy	Cigar-like	In-between fibers	Average diameter ≈ 45 µm	223
	Woven	Carbon/epoxy (5-harness)	Intra-laminar voids: small and nearly spherical or larger and asymmetric → Inter-laminar voids: less wide than those in prepreg fabrics →	At warp/weft intersections Between plies	Spherical voids: diameter between 100 and 200 µm, asymmetric voids: width: 250 µm and 4 mm, height: 150 µm to 400 µm	215
		Carbon/epoxy (plain weave) (5-harness)	Plain weave laminate: elliptical void → 5-harness weave: from elliptical to the more common asymmetric shaped void →	At tow intersection, large through the thickness linkage At tow intersection, more separated through the thickness	Width: 50 µm to 1.2 mm thickness: 20 to 500 µm Width: 50 µm to 3 mm thickness: 20 to 350 µm	216
	NCF	Glass/vinyl ester	Large, spheroidal voids → Long, cylindrical voids with smaller diameter orientated in the flow direction →	Between transverse fiber bundles at positions where weft bundles cross transverse bundles Within fiber bundles mainly at laminate edge	– –	230
		Carbon/epoxy	Irregular shape	Between the adjacent tows	–	231
		Glass/epoxy	Spherical voids → Cylindrical voids → Large irregular cross-section voids →	Between rovings Within the rovings Between plies	Max diameter of about 400 µm Max diameter of 1.6 mm	232,233
		Glass/epoxy	Cigar-like shape → Spherical →	between fibers resin-rich regions	average void diameter ≈ 41 µm	219
		Carbon, glass/epoxy	Needle-shaped → Almost spherical → Long and curved →	Within fiber bundles Between bundles Within or close to the holding yarns	10–1000 µm3	48
	Other	Carbon/phenolic resin (3D braided)	Vacuum assisted RTM: irregular strip, triangular, circular and ellipse → Vacuum and compression co-assisted RTM: polygon shape →	In resin-rich areas Within fiber bundles	Equivalent diameter: 88–262 µm Equivalent diameter: 6–61 µm	234
		Glass/epoxy (random mat)	Spherical shape voids with circular or elliptical cross-section → Cylindrical voids with irregular cross-section →	In matrix/preform interface (predominant) or in matrix-rich areas Inside fibrous preform (predominant)	Equivalent diameter: <50 µm Equivalent diameter: 50–100 µm	150
Other	Random chopped	Glass/epoxy	Approximately spherical shape	–	Varying diameter	235

UD-ply composites

For UD-prepreg composites, characterized by means of the “manual” tomography approach (explained in Void shape (morphology) section), Hsu and Uhl ¹⁴⁷ conducted a 3D analysis of the voids and reported ranges, averages, and distributions for width, height, and length of the voids, located mainly between UD plies, which are presented in Table 4. Wide ranges for these characteristics were found even in one sample. Higher average width and length were observed for samples with higher void contents as observed by Rubin and Jerina. ²²² A size analysis mainly for intra-laminar voids was performed by Gürdal et al. ¹⁴⁴ on QI laminates. The average values for the void cross-section equivalent diameter were within 15–38 µm, with an average of 24 µm, and for the length, within 148–1118 µm, with an average of 497 µm.

OPEN IN VIEWER

Table 4.

Distribution of (mainly inter-laminar) void size parameters in UD-ply carbon/epoxy composites.

Lay-up (reference)	Characterization technique	Void content (%)	Width range (µm)	Average width (µm)	Height range (µm)	Average height (µm)	Average aspect ratio	Length range (µm)	Average length (µm)
UD – [0]16 ( 147 )	Microscopy (manual tomography)	0.20	9.1–120	21.0	3.6–27.3	8.6	0.37	–	–
		1.14	10.3–137	46.4	4.1–31.3	14.4	0.36	–	–
		2.04	3.2–973	69.2	1.5–105	12.6	0.37	–	–
		6.51	1.9–896	66.4	1.1–61.9	11.2	0.32	–	–
Dispersed QI – [+45/−45/0/90]2S ( 147 )	Microscopy (manual tomography)	0.34	12.5–168	33.1	4.6–51.2	10.8	0.45	100–1700	380
		1. 25	12.8–292	56.4	4.8–47.2	15.7	0.42	100–8500	660
		2.82	11.6–631	59.5	4.6–69.8	17.0	0.30	100–10900	1010
		4.05	11.5–995	100	4.6–72.8	19.0	0.32	–	–
Lay-up (reference)	Characterization technique	Void content (%)	Semi-middle axis range (µm)	Semi-middle axis average (µm)	Semi-minor axis range (µm)	Semi-minor axis average (µm)	Aspect ratio range	Semi-major axis range (µm)	Semi-major axis average (µm)
UD – [0]10 (97,236)	Micro-CT	1.1	8.4–50.3	–	8.0–22.0	–	0.29–1.00	38.9–738.9	–
Dispersed QI – [45/0/−45/90]3S (97,236)	Micro-CT	1.78	∼9–66	–	8.3–22.0	17.8	–	25.6–371.4	222.0
		0.12	∼9–67	–	8.1–32.8	18.2	–	28.5–629.3	244.1
		0.60	∼9–67	–	7.7–26.2	17.3	0.24–0.91	29.8–719.2	241.4
Clustered QI – [453/03/−453/903]S (97,236)	Micro-CT	1.30	∼9–59	–	8.4–21.8	12.6	–	32.3–705.1	393.1
		0.24	∼8–60	–	7.6–21.8	12.9	–	31.6–854.0	370.2
		0.26	∼9–44	–	8.3–36.1	13.3	0.21–0.911	26.9–777.5	353.7

Chambers et al. ¹⁵³ classified voids in an OoA UD carbon/epoxy composite by their cross-section area into four size groups: 8–4000 µm², 4000–10,000 µm², 10,000–30,000 µm², >30,000 µm². Having a significant number of voids in the first size group, the number of voids in each size group decreased with the size increase. Agius et al. ¹⁰⁸ reported similar observations for a different definition of size groups: <1000 µm², 1000–2000 µm², 2000–5000 µm², 5000–15,000 µm². Ledru et al., ³ by means of a visco-mechanical based void growth model, calculated that the radius of intra-laminar spherical voids in resin-rich regions would be around 3 µm and experimentally observed that ∼80% of the intra-laminar voids have a radius less than 3 µm. The frequency of voids decreased with the increase in their equivalent radius. However, de Parscau du Plessix et al. ⁹⁶ claimed that this model does not provide acceptable results for the void size. Applying some modifications (explained in the Autoclave curing process section) to the void growth model, they predicted the radius of spherical voids to be within 30–80 µm, depending on the cure cycle.

Little et al. ¹³⁸ presented statistical data on the size of mainly intra-laminar voids with the following interpretation. About 45% of the total number of voids have a projected diameter less than 50 µm, while only accounting for about 6% of the total void area fraction. And 1.5% of the total number of voids have a projected diameter larger than or equal to 200 µm, accounting for about 24% of the total void area fraction. Twenty percent of the total number of voids have a diameter below 20 µm, contributing less than 0.5% to the total void area. Hernández et al. ⁹⁷ measured the major (length) and minor (height) axes of the equivalent ellipsoid for each void detected via micro-CT. The semi-major axis was below 900 µm, and the semi-minor axis was mainly between 8 and 30 µm. These data together with some results regarding the void size from the literature ²³⁶ are also reported in Table 4. The data presented in this table correspond mainly to inter-laminar voids. Through micro-CT analysis, Tserpes et al. ²⁰⁰ confirmed that UD composites with higher void contents have a higher number of large voids, and the larger the void, the longer it is. In a filament wound composite, Scott et al. ²⁰² measured through synchrotron micro-CT, mean diameters of 12.3 µm and 9.9 µm in the two transverse directions and a mean length of 84.7 µm in the fiber direction for intra-laminar voids.

Non-UD-ply composites

For other composite systems, the available data on void size are summarized in Table 3. In woven-prepreg composites, large voids are observed in regions with both low and high void contents, whereas in RTM woven laminates they are only seen in regions with high void contents. ²¹⁵ In woven-prepreg composites, the proportion of larger to smaller voids increases as the void content increases. ²¹⁷ The average size of small and spherical voids located primarily in resin-rich pockets of woven composites varied with changing the void content, but did not monotonically increase as in UD laminates. ⁸⁹ In NCF composites produced through liquid resin infusion, Maragoni et al. ²¹⁹ found an asymmetrical distribution for the void cross-section diameter with an average of 41 µm. They also observed that half of the voids have an area (in top view) below 400 µm². Bodaghi et al. ¹⁴⁹ observed that in a woven carbon/epoxy composite with one type of reinforcing fabric, the void size has a bell-shape distribution for prepreg composites (produced via autoclave or VBO), but a right-skewed distribution for LCM composites (produced by high injection pressure RTM). This was attributed to the random nature of the void size in prepreg composites and its dependence on flow velocity and distance to the gate in LCM composites. Through synchrotron micro-CT analysis of an RTM quasi-UD NCF carbon/epoxy composite by Sisodia et al., ⁴⁸ the volume of voids was found to be within 10–1000 µm³ and suggested a bell-shape distribution. Melenka et al. ¹⁹⁰ fitted the void volume distribution for a Kevlar/epoxy braided composite, measured through micro-CT, by a Weibull distribution, as presented in Figure 17. The mean void volume, obtained through the Weibull distribution was 7.4 × 10⁶µm³ with a variance of 7.3 × 10⁴ µm³. The median void volume was found to be 3.3 × 10⁶µm³. OPEN IN VIEWER

ILSS in UD-ply composites

Carbon/epoxy UD-ply composites: In one of the earlier works, Judd and Wright ⁷ conducted a review on the influence of voids on mechanical properties, in particular ILSS, of FRCs. They drew the conclusion that regardless of the constituent materials, ILSS reduces almost linearly by ∼6% for every 1% increase in void content, up to about 4% of void content, beyond which the reduction rate diminishes. Yoshida et al. ²³⁹ found an average coefficient of correlation of −0.84 that shows a highly linear relationship between ILSS and void content, with similar proportionality to that of the literature. ⁷ Using the analysis of variance technique, they calculated a critical void content of 0.5%, below which no significant change in ILSS is detected. According to Ghiorse, ¹⁴⁰ each 1% increase in void content in the range of 0–5% leads to ∼10% decrease in ILSS, which is significantly higher than the value found in Judd and Wright. ⁷ He argued that a “significant decline” in properties happens in the presence of voids regardless of the constituent materials, the void characteristics, and the loading mode. Tang et al. ⁹² found an increase in void sensitivity of ILSS at higher void contents, which was also reported by Liu et al. ¹⁰¹ Although Liu et al. ¹⁰¹ attributed this to the decrease in ply interface area, this trend cannot be identified in their presented ILSS versus void content plot. Their observation of 6% reduction in ILSS with each 1% increase in void content is in line with the conclusion from Judd and Wright. ⁷

Uhl et al. ⁸⁹ noticed a higher sensitivity of ILSS to the void content in UD laminates compared to that of QI laminates. A simple empirical exponential model, attributed to the literature, ²⁴⁵ was used by them to fit the monotonic reduction in ILSS with void content. According to this model, the strength (σ) at a void content of V_v (can be correlated to the US attenuation, as explained in the Ultrasonic testing section) can be calculated from the following equation, where σ max is the strength at zero voidage and B is an empirical constant that depends on void characteristics

σ = σ max e - BV v

(9)

In a step-forward research, de Almeida and Neto ¹⁵⁴ used a fracture criterion, based on linear elastic fracture mechanics, analogous to the equation proposed by Mar and Lin²⁴⁶ for tensile failure of FRCs with holes, to correlate the dependency of (inter-laminar shear) strength on void content, or exchangeably US absorption coefficient of FRC laminates (see the Ultrasonic testing section). This criterion implies that the fracture stress σ f of an FRC with the absorption coefficient of α is

σ f = { σ f 0 if α ≤ α cr H ( α ) - m if α ≥ α cr

(10)

Recent research does not focus on void content anymore. Olivier et al. ¹⁰³ attributed the high sensitivity of ILSS to voids to the fact that voids are mainly located between the plies, where failure in the ILSS test is triggered. Moreover, they showed that besides the void content, the size and location of voids play a role in the effect on ILSS. Wisnom et al. ²²⁴ performed a more systematic study on failure affected by voids, and observed that failure initiates from the (artificially produced) discrete voids for carbon/epoxy specimens. ILSS decreased consistently with the length of the voids increasing; more specifically, the increase of length from 0.28 mm to 3.00 mm led to a decrease of 23% in ILSS, with regard to the reference laminate without voids. It is not clear whether the void content was kept constant, when the void length was varied. They found out that linear elastic fracture mechanics could not properly predict the reduction in strength of specimens starting to fail form voids, since it does not deal with failure initiation. However, an FE method, including both a stress criterion for yielding and an energy release rate criterion for failure, worked very well in prediction of failure initiation and propagation, regardless of the shape of the void. ²²⁴ In addition to Olivier et al. ¹⁰³ and Wisnom et al., ²²⁴ the effect of void size was also studied by Kousourakis et al. ²⁴⁷ Placing removable silicone tubes to create sensor cavities used for structural health monitoring, representing elliptic–cylindrical voids, at the mid-plane of the specimens, they found that ILSS decreases linearly with increasing the diameter or the number of the cavities. The cavities were aligned perpendicular to the direction of inter-laminar shear cracking.

Davies et al. ¹¹⁶ concluded that in the presence of voids, other parameters such as fiber/matrix adhesion can also influence ILSS. They presented peculiar results showing that despite a higher void content, Quickstep (OoA) cured panel exhibited an improvement in ILSS over the autoclave cured panel, but this was observed only for normalized (to a common equivalent fiber volume fraction) ILSS. The authors suggested that this could be due to enhancement in interfacial fiber/matrix adhesion. Olivier et al. ²⁴³ proposed equations defining ILSS as a function of the void content and Mode II inter-laminar fracture toughness. Koushyar et al. ¹⁰⁵ witnessed a severe decline in ILSS by voids for contents above 1%. They fitted the decline with an exponential equation. In the post-mortem micrographs, the voids seemed to have a major role in crack propagation through the plies, and inelastic deformation was more pronounced for composites with higher void contents. Assuming that cylindrical voids are arranged in a regular square pattern, Hernández et al. ¹²⁸ proposed a simple analytical net-section model to correlate the ILSS reduction with the increasing void content. One specimen, out of four, did not follow the model trend. As the nano-indentation and the push-in tests showed similar matrix and interface properties for specimens with different void contents, they concluded that the lower ILSS of that specific specimen, in spite of its lower void content, could be due to the void morphology.

In contrast to the previous studies, Nikishkov et al. ¹⁹⁶ claimed that ILSS has a relatively low dependency on void content. Taking advantage of micro-CT as a precise characterization tool, they showed that the crucial parameter is the presence of large voids in critical locations. In their case of short-beam-shear specimen, the mid-section of the specimen was the critical location due to the inter-laminar shear stress concentration. The inter-laminar failure appeared as two cracks extending in half of the specimen length, where the critical void lay: one initiating at the critical void, and the second one immediately after the start of the first crack due to the reduction in the effective specimen thickness. Through a numerical methodology, employing a LaRC04 fracture toughness-based failure criterion, they could predict the load level and location of failure for laminates with significant voidage. Stamopoulos et al. ¹⁹⁸ reported 11% reduction of ILSS for ∼3% of voidage. Similarly to Olivier et al., ¹⁰³ they stated that the more severe effect of voids on ILSS, compared to other properties, is because many voids are placed between the plies. Furthermore, based on the pre-analysis of the specimens via micro-CT, they argued that for composites with similar void contents, the degradation of properties is more severe in the one with few large voids than in the one with many small voids. Thus, the insufficiency of the void content and the necessity of consideration of other void characteristics in the evaluation of composite properties, particularly ILSS, are endorsed by the use of micro-CT.

Glass/epoxy UD-ply composites: In one of the first studies on the effect of voids, Bascom and Romans ²⁴⁸ observed that reducing the void content from 5% to less than 0.1%, through control of the filament winding process, leads to a 50% increase in ILSS. In a similar and simultaneous work, Kohn et al. ¹⁴² indicated that a linear inverse relationship describes adequately the correlation between ILSS and void content. Thomason ²⁴⁹ showed that ILSS, besides linearly correlating with void content, depends enormously on the nature of the fiber surface coating.

Wisnom et al. ²²⁴ conducted a broad research on ILSS of glass/epoxy composites influenced by discrete as well as distributed voids. The discrete voids were artificially created by small PTFE inserts at the laminate mid-plane, and the distributed voids were produced by eliminating the positive pressure during cure. Unlike in carbon/epoxy specimens, failure hardly initiated from the voids in glass/epoxy specimens, even for those with discrete voids, in which the voids were large compared to the cross-section. Nevertheless, the failure initiated mostly above and below the void. In this case, the increase in stress due to the reduction in net cross-section seemed to play a critical role. Therefore, net section stress analysis would be more appropriate than the stress-based failure criteria, which predict the strength based on the stress concentration caused by voids. Through this analysis, Wisnom et al. concluded that the stress concentrations associated with the voids are less critical than might be expected. For the specimens with highly distributed voidage also, the failure was controlled mainly by the reduction in net cross-section. They suggest that the degradation of ILSS with increase in voidage is caused by a combination of the decrease in cross-sectional area and the failure initiation from large voids, as stress concentrators.

Other UD-ply composites: The influence of voids on ILSS in thermoplastic composites exhibits the same phenomena as in thermoset composites. Bowles and Frimpong ²²¹ investigated carbon/polyimide (PI) laminates. They found that for low void contents, the ILSS depends on the location of the voids, such that voids closer to the inner, high-shear-stress areas of the specimen become more critical. The standard deviation of ILSS increased with increasing void content, attributed to the increase in randomness of void location, so that the chance that a void has a critical location to cause premature failure scatters. The relationship between ILSS and composite density was fitted with both linear and power relationship equations. The composite density was expressed in terms of the fiber volume fraction and void content. Although they observed that the power equation fits better to the measurements, the predictions were not highly accurate. According to de Almeida and Neto, ¹⁵⁴ the reason can be that these equations neglect stress concentrations and material toughness, and assume non-realistic geometries for voids.

On a composite manufactured by pultrusion of a yarn, containing discontinuous fibers of carbon and polyamide, Wiedmer and Manolesos ²⁵⁰ concluded that ILSS is the most sensitive property to process parameters. Although a high pulling speed results in a high voidage and low ILSS, proper preheating treatments provided low-voidage composites, even at high pulling speeds.

The relationship between ILSS and voids in natural-fiber composites is also a point of interest. Li et al. ²²⁶ studied the formation of voids and their effect on the mechanical properties of quasi-UD flax/epoxy composites. They noted that with each 1% decrease in voidage, ILSS increases ∼11%, and concluded that flax FRCs are more void-sensitive than carbon and glass FRCs (5–8% ILSS increase with 1% decrease in voidage).

ILSS in woven-ply composites

Carbon/epoxy woven composites: Uhl et al. ⁸⁹ and later Jeong ¹⁴⁶ spotted the same trend of ILSS reduction for woven plies as for UD laminates, but with a lower sensitivity. The void sensitivity of woven composites was found to be between that of UD and QI laminates of UD plies. Jeong ¹⁴⁶ suggested that the higher void sensitivity of UD laminates than that of fabric composites could be due to “the natural resistance of fabrics to crack propagation in the matrix”. One can understand it as reference to intermingling of woven plies, which affects the crack propagation. The reason might also be the difference in void morphology and location as he noted that UD laminates have mostly elongated voids at the ply interfaces, whereas the voids in woven laminates are more spherical and located in resin-rich pockets. Goodwin et al. ²¹⁶ noted a difference in the dependency of ILSS on voids between plain weave and 5-harness fabric laminates, produced through RTM. For each 1% increase in void content up to 10%, the plain weave laminate showed 4% reduction in ILSS, while the 5-harness had 7% reduction. The higher sensitivity of the 5-harness laminate was attributed to the higher stress concentrations resulted from a higher level of asymmetric voids with corners of lower radii, higher stress intensification caused by the longer voids, and lower stress for crack initiation due to a higher population of voids throughout the laminate. Intra-laminar cracks were observed to initiate from voids, and to propagate to the ply interface, transforming into inter-laminar cracks. A reduction of 25% in ILSS of a woven carbon/epoxy composite was reported by Di Landro et al. ¹⁸⁸ for ∼6% increase in void content.

Costa et al. ¹⁵⁵ applied the fracture criterion (equation (10)) proposed in the literature ¹⁵⁴ to find a correlation between the measured values of ILSS and the void content for a woven carbon/epoxy composite and could establish a reasonable fit. They noted that during short-beam-shear tests, intra-laminar transverse cracks emanate from the matrix voids. In a later study, ²⁵¹ they also concluded that the slope parameter m is mainly controlled by the ply reinforcement structure. UD-ply laminates showed to be more void-sensitive than woven-ply laminates.

Zhu et al. ²¹⁷ observed a higher scatter in ILSS with increasing void content, which was attributed to the shapes and sizes of voids as the initiation and propagation of cracks were affected by the voids. The ILSS of the laminates hosting voids with a higher aspect ratio was more sensitive to voidage. Zhang et al. ²⁵² established a macro-level FE analysis to predict the laminate strength, including ILSS, in the presence of voids, based on the improved Hashin failure criterion, using the tensile, compressive, and shear strength values for plies with different void contents as input. Through this model, they could predict the decrease in ILSS of a laminate with increasing void content.

Other woven plies: Woven carbon/PI^89,146 and woven carbon/bismaleimide composites ¹⁵⁵ exhibited a similar trend to woven carbon/epoxy composites in reduction of ILSS with voidage. On an example of a woven glass/polyester composite, Mouritz ¹⁵⁷ showed that although there is a clear relationship between ILSS and US absorption coefficient, even at extremely high void contents (up to 30%), the Mar and Lin theory,²⁴⁷ modified by de Almeida and Neto (equation (10)), ¹⁵⁴ cannot correlate this relationship in FRCs with high level of voidage (above ∼15%). In case of a thermoplastic matrix, Bureau and Denault ²⁵³ showed that ILSS of glass/PP composites consolidated in different molding conditions is linearly correlated with their void content.

Concluding remarks on the effect of voids on ILSS

Table 5 makes a summary of literature treating the void effect on ILSS. As a general guidance, it can be accepted that 1% increase in void content in carbon/epoxy and glass/epoxy UD-ply and woven composites leads to 5–10% decrease in ILSS, with woven laminates showing a lower sensitivity than UD-ply. More detailed studies reveal dependencies on void configuration, laminate architecture, matrix toughness, etc., but generalization on these details does not seem feasible for the current state-of-the-art. The main mechanisms of the ILSS decrease are believed to be the reduced cross-sectional (“presented”) area because of the presence of voids at the inter-laminar interfaces and the stress concentrations around the voids. However, a full theoretical/modeling treatment of the ILSS in the presence of voids is still to be developed.

Only about 5% of all the studies dealing with the effect of voids on ILSS take advantage of modeling approaches. This is attributed to the lack of simulation tools that could predict composite failure by simultaneously modeling fibers and voids – features of distinctly different sizes and properties – which strongly interact.

Tensile elastic properties

Engineering constants of UD composites, influenced by inter-laminar voids, were studied numerically by Huang and Talreja. ² The normalized reduction of the in-plane moduli, i.e. longitudinal (E_x) and transverse (E_y) tensile and shear ( G xy ) modulus, with voids was linear and not substantial. For different aspect ratios of voids, a 5% increase in void content caused a reduction in the range of 5–10% for E_x, in the range of 6–12% for E_y, and in the range of 6–7% for G xy . On the other hand, the normalized reduction of the out-of-plane moduli, i.e. out-of-plane tensile (E_z) and out-of-plane shear ( G xz ) modulus was significantly higher such that with 5% increase in void content, E_z reduction was in the range of 12–40%. The effect of void morphology was investigated as well. The width-to-height aspect ratio (the shape of the void cross-section) was found to have a more significant effect on reduction of the moduli, than the length-to-width aspect ratio does, except for the out-of-plane shear stiffness G xz , which was sensitive to both aspect ratios. Larger width-to-height aspect ratio (more flat voids) led to less modulus reduction for in-plane moduli, but larger reduction for out-of-plane moduli. Higher length-to-width aspect ratio (longer voids) caused more degradation of G xz . A good agreement was noted between the results of the FE analysis and an analytical modeling based on the Eshelby’s equivalent inclusion theory (Mori–Tanaka solution). The results of this study were presented in the literature^4,5 as well. In the following, more studies on elastic properties, influenced by voids, are reviewed.

Tensile modulus: In an early study, Harper et al. ²⁵⁴ and later Olivier et al. ¹⁰³ found that in UD-ply composites, fiber-dominated elastic properties, i.e. longitudinal tensile modulus and Poisson’s ratio, related to contraction in the fiber direction, are practically insensitive to voidage, but the matrix dominated moduli, i.e. transverse tensile and shear modulus, degrade significantly with voids. Liu et al., ¹⁰¹ Zhan et al., ¹⁰⁶ and Stamopoulos et al. ¹⁹⁸ for UD carbon/epoxy composites, and Li et al. ²²⁶ for quasi-UD flax/epoxy composites confirmed the low sensitivity of longitudinal tensile modulus to voids. Selmi ²⁵⁵ and Huang and Talreja ² concluded the low sensitivity through analytical and numerical modeling. For quasi-UD NCF laminates (produced through RTM) also, the detrimental effect of voids on transverse tensile modulus was reported by Varna et al., ²³⁰ where 5% voidage caused ∼15% decrease. Experimental results of Gürdal et al. ¹⁴⁴ showed that the out-of-plane tensile modulus is largely sensitive to voids such that 1% increase in void content leads to ∼10% decrease in out-of-plane modulus.

For woven composites, Farouk et al. ²⁵⁶ proposed an analytical model, based on Weng’s inclusion approach, and Ishikawa and Chou’s crimp method, for predicting the effect of void content as well as aspect ratio on the longitudinal modulus. They showed that the longitudinal modulus reduces negligibly with increasing void content. For a fixed void content, the longitudinal modulus increases with moving from oblate to prolate voids, and reaches a plateau at an aspect ratio around 1.5. Through a simple numerical methodology considering randomly distributed voids in the matrix, Van Den Broucke et al. ²⁵⁷ found that in a carbon/epoxy woven composite, produced with resin infusion process, 1% increase in void content results in a reduction of ∼5% for in-plane (because of the plate thickness increased with the voidage) and ∼7% for out-of-plane tensile modulus. Taking into account the real geometry of a woven composite, obtained via micro-CT, Choudhry et al. ²⁵⁸ presented an FE micromechanical approach, which accounts for the yarn volume fraction and the composite voids. The voids were taken into account by reducing the matrix properties, through simply multiplying them by a void compensation factor (explained in the Prediction of void effects in the elastic regime section). With this modification, the resulting tensile modulus in the warp direction and the in-plane shear modulus were decreased so that they matched the experimentally measured results.

Naganuma et al. ¹⁰⁴ observed that the effect of voids on tensile modulus is dependent on the morphology of the voids in woven composites, as is the case for tensile strength (explained in the Tensile failure in multidirectional laminates section). When the residual solvent of prepreg is below 20 wt% (enough drying time), closed voidage (flat and semicircular voids between the prepregs) is prevalent (explained in the Autoclave curing process section), with almost no effect on tensile modulus. However, above 20 wt% of the residual solvent, open voidage (large and irregularly shaped voids around the bundle intersections, open to the thickness surface) become dominant, reducing the tensile modulus. The effect of voids can be dissimilar for different stacking sequences. Zhu et al. ²²⁸ showed that the 0°-tensile modulus of a lay-up of carbon/epoxy woven plies with a high number of ±45° plies decreases 14% by ∼8% of voidage, whereas the modulus of another lay-up with a high number of 0° plies is virtually insensitive to voids.

The void sensitivity of 3D braided carbon/epoxy composites is numerically explored by Xu and Qian. ²⁵⁹ In their model, some elements in both yarn and matrix regions were randomly selected to represent yarn and matrix voids, respectively. They concluded that for a given void content, the elastic properties are not sensitive to the random distribution of void. The reduction level of transverse tensile moduli with voids was larger than that of longitudinal tensile modulus. The effect on longitudinal tensile modulus (in the 0° direction w.r.t. the braiding angle) increases with increasing the braiding angle. In a similar study, the degrading effect of yarn and matrix voids on elastic properties (except on longitudinal Poisson’s ratios) of a braided carbon/epoxy was endorsed by Dong and Huo ²⁶⁰ through a two-scale numerical approach. They concluded that the effect of yarn voids is more remarkable than that of the matrix voids.

For discontinuous FRCs (glass/epoxy), a modeling study was carried out by Srinivasulu et al., ²⁶¹ in which randomly oriented fibers and spheroidal voids are numerically modeled in an FE simulation, and analytically modeled, using the Mori–Tanaka method along with orientation averaging. Both methodologies predicted a reduction in the effective (homogenized) tensile modulus and Poisson’s ratio with the void content increase. The analytical results were more close to the numerical results at lower void contents. The results showed that the aspect ratio of the void and void size distributions have less influence on the effective elastic properties of discontinuous FRCs, compared to the effect of void content. Furthermore, FE analysis proved that the degrading effect of voids is more significant for curved fiber composites than that for straight fiber composites. For random-chopped FRCs, Yang et al. ²³⁵ derived a micromechanical constitutive equation accounting for voids, and incorporated it into an FE code. The numerical results showed an insignificant effect of spherical voids on the tensile modulus even for high levels of voidage. A summary of the studies on void influence on tensile modulus is presented in Table 6.

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Table 6.

Effect of voids on the tensile modulus of different composites as reported in the literature.

Ply structure	Material	Lay-up	Production technique	Void content range (%)	Notable observation(s)	Reference
UD	Carbon/epoxy	UD	Prepreg	1–5.5	Insensitivity of longitudinal tensile modulus and sensitivity of transverse tensile and shear modulus to the voids	254
		UD	Prepreg–autoclave	0–10	Insensitivity of longitudinal tensile modulus and sensitivity of transverse tensile and shear modulus to the voids	103
		UD	Prepreg–autoclave	0–6	Linear and not substantial reduction of the in-plane moduli with voids; significant reduction of the out-of-plane moduli with voids; more significant effect of the width–height aspect ratio than the effect of the length–width aspect ratio, except for the out-of-plane shear stiffnessa	2
		CP	Prepreg–autoclave	0–3.5	Low sensitivity of longitudinal tensile modulus to voids	101
		UD	Prepreg–autoclave	0–6	Presenting some of the results of 2  a	4
		UD	Prepreg–autoclave	0.7–5.4	Low sensitivity of longitudinal tensile modulus to voids	106
		UD	N/A	0–15	Low sensitivity of longitudinal tensile modulus to voidsa,b	254
		UD	Prepreg–autoclave	0.8–3.4	Low sensitivity of longitudinal tensile modulus to voids c	198
	Carbon/polyester	UD	N/A	0–8	Presence of voids: less significant decrease in transverse Young’s modulus than that in transverse shear modulusa	263
	Flax/epoxy	UD	Prepreg–hot press	0.8–3.4	Low sensitivity of longitudinal tensile modulus to voids	226
Woven	Carbon/PI	N/A	Compression molding	0–20	Increase in longitudinal modulus with moving from oblate to prolate voids, a plateau at an aspect ratio around 1.5 b	256
	Carbon/epoxy	N/A	Resin infusion	0–5	Reduction of ∼5% for in-plane and ∼7% for out-of-plane tensile modulus with 1% increase in voidagea	257
		[(±45)4/(0,90)/(±45)2]S [(±45)/04/(±45) (0,90)]S	prepreg–autoclave	0.4–9	Dependency of the effect of voids on the stacking sequence	228
	Glass/PI	[(0/90)]4	Prepreg–autoclave	0–10	Almost no effect of closed voidage on tensile modulus, but reduction of tensile modulus by open voidage (to the thickness surface)	104
	Glass/phenolic resin	[0/0flip]2	Prepreg– Quickstep	5.3	Decrease in the predicted tensile modulus in warp direction and the in-plane shear modulus with considering voids, resulting in a match with the experimentally measured resultsa,c	258
Other	Glass/vinyl ester (UD NCF)	Quasi-UD	RTM	0–4.5	Detrimental effect of voids on transverse tensile modulus	230
	Carbon/epoxy	QI NCF	RTM	0.8–5	Negligible influence of voids on tensile stiffness b	240
	Carbon/epoxy	(braided)	RTM	0–3	Higher reduction of transverse tensile moduli with voids than that of longitudinal tensile modulus (in the 0° direction w.r.t. the braiding angle)	259
	Glass/epoxy	(random discontinuous fibers)	N/A	0–12	Reduction in effective (homogenized) tensile modulus and Poisson’s ratio with increasing void content; less influence of aspect ratio of the void, compared to the effect of void content, on the elastic properties; higher void sensitivity of curved fiber composites than that of straight fiber compositesa,b	261
	Glass/epoxy	(random-chopped fibers)	Multi-spray laying up	10.1–10.7	Insignificant effect of spherical voids on tensile modulus even for high levels of voidagea	235
	Carbon/epoxy	(braided)	RTM	0–12	Degrading effect of yarn and matrix voids on elastic properties, except for longitudinal Poisson’s ratios; more remarkable effect of yarn voids than that of the matrix voidsa	260

Shear modulus: Longitudinal shear modulus ( G LT ) of cylindrical (tubular or rod) UD carbon/epoxy specimens was measured in a torsion test by Hancox. ²⁶² For low void contents (up to ∼1.5%), the reduction of shear modulus with voids was found to agree with theoretical models and was ∼10% for 1% increase in void content. Above this void content, the shear modulus dropped severely, so that for 5% of voidage, the reduction reached 70%. The author attributed the drop in modulus to the type of voids and their distribution.

Rubin and Jerina ²²⁰ developed an analytical model to calculate reduced elastic properties, accounting for inter-laminar voids. Different ply interfaces were assigned to have a different length of the porous section, which was an input variable for the model. A linear correlation between US attenuation and the total inter-laminar void length to the ply thickness ratio was found. The compression and in-plane shear moduli, measured in the rail shear and open-hole compression tests, respectively, matched the analytical results. The rate of moduli reduction depended on the laminate type, laminate thickness, and stacking sequence. For a woven carbon epoxy laminate, the highest reduction rate of in-plane shear modulus was found for the lay-up with the minimum number of 45° plies. For UD laminates, the reduction rate for the shear and transverse compression moduli is much higher than that for the longitudinal compression modulus, and the reduction rate increases with the fraction of 0° plies.

Numerical results of Van Den Broucke et al. ²⁵⁷ showed that in a woven composite, 1% increase in the volume of randomly distributed voids in the matrix results in a reduction of ∼4% for in-plane and ∼10% for out-of-plane shear modulus. In a numerical analysis on a 2D elastic FE model of a UD composite with circular voids in the matrix, Nikopour ²⁶³ obtained a higher decline of transverse shear modulus ( G TT ) and transverse Poisson’s ratio ( ν TT ) than that of transverse Young’s modulus ( E TT ). Selmi ²⁵⁵ showed that the difference between the transverse tensile modulus of a 2%-voidage UD composite and that of a void-free UD composite increases with increasing fiber volume fraction. This was the case also for the longitudinal shear modulus, while there was practically no difference in their longitudinal tensile modulus. This was achieved through a two-step analytical modeling based on the Mori–Tanaka homogenization scheme validated against an FE simulation. The numerical results from Xu and Qian ²⁵⁹ for a braided composite showed that in-plane and out-of-plane shear moduli decrease slightly and monotonously with the void content increase up to 3%. This degrading effect was less pronounced for higher braiding angles. A summary of the studies on void effect on shear modulus is presented in Table 7.

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Table 7.

Effect of voids on the shear modulus of different composites as reported in the literature.

Ply structure	Material	Lay-up	Production technique	Void content range (%)	Notable observation(s)	Reference
UD	Carbon/epoxy	UD tube/rod	Pultrusion, filament winding	0.5–4.6	Monotonic reduction in shear modulus ( G LT ) with increasing void content up to ∼1.5% before a drastic modulus drop a	262
		N/A	prepreg	N/A	Linear correlation between US attenuation and the total inter-laminar void length/ply thickness ratio; good agreement between the analytical results and rail shear test data; dependency of the rate of moduli reduction on laminate type, laminate thickness, and stacking sequence; the highest reduction rate of shear modulus for the lay-up with the maximum number of 0° plies a	220
		UD	Prepreg–autoclave	0–6	High sensitivity of out-of-plane shear modulus to inter-laminar voids than that of in-plane shear modulus; more detrimental effect of larger width–height and length–width aspect ratios of inter-laminar voids on out-of-plane shear modulus b	2
		UD	N/A	0–15	Numerical and analytical results: increase in the difference in longitudinal shear modulus and transverse tensile modulus of a 2%-voidage composite compared to a void-free composite with increasing fiber volume fraction b	255
	Carbon/polyester	UD	N/A	0–8	Higher decline of transverse shear modulus and transverse Poisson’s ratio than that of transverse Young’s modulus b	263
Woven	Carbon/epoxy	N/A	Prepreg	N/A	The highest reduction rate of shear modulus for the lay-up with the minimum number of 45° plies a	220
		N/A	Infiltration	0–5	Numerical results: reduction of ∼4% for in-plane and ∼10% for out-of-plane tensile modulus with 1% increase in voidage b	257
Braided	Carbon/epoxy	N/A	RTM	0–3	Higher reduction of transverse tensile moduli with voids than that of longitudinal tensile modulus b	259

Tensile failure in UD laminates

The effect of voids on composite tensile strength in the fiber axial direction received less attention in the literature. Olivier et al. ¹⁰³ noticed a very small reduction of longitudinal tensile strength, being ∼10% decrease with increase in void content from 0% to 10%. They explained that this reduction by the voids could be caused by local fiber deformations and micro-debonding between matrix and fibers.

In contrast, voids have a significant influence on the transverse tensile properties. In an analytical study, Greszczuk ²⁶⁴ derived a failure criterion for transverse tensile loading of UD composites, based on the Hencky-von Mises distortion energy theory, assuming a square array of cylindrical voids. The tensile strength was analytically evaluated as a function of the properties and volume fraction of the constituents, including voids, and the interaction between stress concentrations due to fibers and voids was taken into account. Olivier et al. ¹⁰³ reported ∼30% reduction in transverse tensile strength of a UD carbon/epoxy composite with increase in void content from 0% to 10%. In a later study, ²⁴³ they fitted the reduction to a power–law relationship. Stamopoulos et al. ¹⁹⁸ noted a considerable (∼15%) decrease in the transverse tensile strength with an increase of 3% in void content. The non-linearity of the reduction was attributed to the shape, size, and location of voids, evaluated via micro-CT. Unlike what is reported for the scatter of ILSS in UD ²²¹ and woven laminates, ²¹⁷ the scatter (standard deviation) for the transverse tensile strength was reduced at higher void contents.

Through a numerical analysis, Chowdhury et al. ²⁶⁵ investigated the local failure of a UD glass/poly(methyl methacrylate) (PMMA) composite, under transverse loading. Their model, explained in the Strength prediction with voids section, accounted for a rate-, pressure-, and temperature-sensitive behavior as well as for crazing in the matrix, and an energy-based debonding criterion for fiber/matrix interface. At low strain rate and room temperature, fracture started from the fiber/matrix debonding. The model with void showed higher macroscopic fracture strain. It might be due to some crazing that initiated and developed near the void before the main crack affected the void. Upon increasing the temperature, fracture initiated at the void, induced by crazing in the polymer. This was attributed to the thermal softening of the matrix and the decline in the crazing critical stress. Similarly, at high strain rates, the craze-induced damage initiated from the void edge. It should be noted that in spite of valuable results of this study, they are highly dependent on the input material properties. Moreover, the equality of fiber volume fraction and void content (modeling 2 circle quarters of the same size in the unit cell, one as the fiber and one as the void) as well as the large distance between fiber and void are too unrealistic in this numerical study.

The effect of very small inter-fiber micro-voids on the transverse strength properties of UD glass/epoxy composites was evaluated by Ashouri Vajari et al., ²⁶⁶ using micromechanical models. This was accomplished through FE modeling, prescribing an isotropic modified Drucker–Prager model for the matrix and cohesive zone model for the fiber/matrix interface. Micro-voids were generated either trapped between fibers (irregular shape) or dispersed in the matrix (circular). The simulation results showed that voidage reduces the strength, but does not alter the dominant failure mechanism, which was interface debonding. The initial damage, and thus the strength, was governed by the trapped voids, but the circular voids could change the final crack path if they were large enough (Figure 19). It was shown that a small amount of micro-voids (1%) drastically reduces transverse tensile strength and the ultimate strain. Further increase of the void content up to 5%, however, does not lead to such significant degradation in strength and failure-strain. This is because only one or two trapped voids are sufficient to cause damage initiation, and hence reduction in strength. A good agreement was found between the failure locus of the composite lamina obtained from the micromechanical model and that predicted by the Puck’s model. The authors surprisingly attributed the lower transverse tensile strength of the composite, compared to the matrix tensile strength and the interface strength, to only the presence of micro-voids. It is known that even without voids, the transverse strength of a composite is lower than the strength of the matrix and interface, which is a result of stress concentrations generated by fibers. Figure 19 also highlights the necessity for modeling the stochastic nature of voidage. With only one realization of the placement of voids for each void radius, it seems that the ultimate strain increases with the increasing void radius; such an effect should be supported by considering a statistically representative sampling of the placement of voids. OPEN IN VIEWER

In a similar study, ²⁶⁷ it was found that the spatial distribution of inter-fiber micro-voids, with a fixed void content of 2%, can cause a 17% deviation in the transverse tensile strength. Furthermore, 1% voidage can decrease the strength by 25% under combined loading conditions of transverse tension and longitudinal shear loading.

Based on micro-CT analysis, Tserpes et al. ²⁰⁰ proposed a numerical methodology, which is performed in three consecutive scales. The results, falling within the range of experimental scatter, showed that the transverse tensile strength is highly void-sensitive. However, they claimed that the transverse tensile stiffness is negligibly sensitive to voids, even for void contents of ∼3.5%, which contradicts to the literature.^2,103,254

Tensile failure in multidirectional laminates

In a quasi-UD NCF glass/vinyl ester laminate (5% of fibers in the transverse direction) produced through RTM, Varna et al. ²³⁰ found an insignificant effect of voids on transverse tensile strength. In contrast, the transversal strain-to-failure was much more sensitive to voids, increasing significantly with increasing void content. In specimens with a low void content, only a few large transverse cracks were found prior to failure, while high-voidage specimens had several large transverse cracks with irregularly shaped and numerous smaller cracks. They proposed a simple analytical model, based on the “shear lag” approach, to show that in laminates with a high void content, the stress level inside the small amount of weft bundles oriented in the loading direction is lower, compared to the laminates without voids (due to the irregular cracks). This could explain the higher strain-to-failure of laminates with voids. The negligible influence of voids on the tensile strength is also reported for QI NCF composites produced with (VA)RTM in the literature.^233,240

In a cross-ply carbon/epoxy laminate, ∼15% reduction in tensile strength for an increase of 2.5% in voidage was reported by Liu et al. ¹⁰¹ In the literature, ²²⁵ the tensile strength of a cross-ply laminate was fitted with the failure criterion (equation (10)) proposed by de Almeida and Neto ¹⁵⁴ for shear and flexural strength (explained in the Inter-laminar shear strength section). Fedulov et al. ²⁶⁸ used a simple numerical methodology to predict the influence of voids on the tensile failure load of cross-ply and QI laminates. The input strength parameters for an individual ply in the absence and presence of voids were estimated based on the (previously measured) experimental data. The void content was found to reduce the tensile failure load, which was a consequence of the shear strength reduction.

In woven glass/PP laminates with different void contents, produced at different molding conditions, Bureau and Denault ²⁵³ noted an increase in ultimate tensile strength when the void content decreased, except for the “extreme temperature condition” processing that led to a lower strength in spite of a fairly low void content. This might be due to the occurrence of thermal degradation in the material, and not linked to voids. In a satin woven glass/PI composite, Naganuma et al. ¹⁰⁴ found that the effect on the strength depends on the voidage type, which was described in the Autoclave curing process section: The open voidage (to the thickness surface) markedly reduces the tensile strength of the woven composite, whereas the closed voidage has a negligible effect. Similarly, Zhu et al. ²²⁸ observed that long voids, present in high-voidage specimens, can initiate cracks, while no cracks were found to emanate from the voids in the low-voidage specimens. For two different lay-ups, one with a high number of ±45° plies and the other with more 0° plies, the degradation of tensile strength with 9% of voidage was ∼2%, indicating a low sensitivity of the tensile strength of woven composites to voidage. The slight reduction of tensile strength of a woven fabric by voids was also numerically predicted by Zhang et al., ²⁵² using the strength values for plies with different void contents as input. The studies about the effect of voids on tensile failure are summarized in Table 8.

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Table 8.

Effect of voids on the tensile failure of different composites as reported in the literature.

Laminate structure	Ply structure	Material	Lay-up	Production technique	Void content range (%)	Notable observation(s)	Reference
UD	UD (transverse)	Carbon/epoxy	UD	Prepreg–autoclave	0–10	∼20% reduction in transverse tensile strength with 10% increase in void content	103
				Prepreg–autoclave	0.3–10.3	Fitting the reduction in transverse tensile strength to a power–law relationship	243
				Prepreg–autoclave	0.8–3.4	A considerable non-linear reduction in transverse tensile strength with increase in void content; decrease in the strength scatter at higher void contents a	198
				Prepreg–autoclave	1.6–3.4	The reduction in transverse tensile strength predicted with a combined-level numerical methodology falling within the range of experimental scattera,b	200
		Glass/epoxy	UD	N/A	N/A	Deriving an analytical equation for determining the transverse strength, taking into account the voids and their resulting stress concentrations c	264
				N/A	0–5	Dependence of the strength on the trapped voids, and dependence of the crack path on the circular voids; good agreement between the failure locus obtained via micromechanical modeling and that predicted by the Puck’s model b	266
				N/A	0–5	High importance of the spatial distribution of voids in degrading the transverse tensile strength b	267
		Glass/PMMA	UD	N/A	N/A	Dependency of location of failure initiation (either from fiber/matrix interface or from void) on strain rate and temperature; higher macroscopic fracture strain for the model with void b	265
	UD (longitudinal)	Carbon/epoxy	UD	Prepreg–autoclave	0–10	A very small reduction of longitudinal tensile strength with voids (∼10% decrease with 10% increase in void content)	103
Multi-directional	UD	Carbon/epoxy	CP	Prepreg–autoclave	0–3.5	∼15% reduction in tensile strength for 2.5% increase in voidage	101
			CP	Prepreg–autoclave	0–3.5	Using the fracture criterion proposed (equation (10)) in the literature 154 to correlate tensile strength with void content	225
			CP, QI	N/A	0–4	Reduction in tensile failure load with increasing the void, as a consequence of the shear strength reduction b	268
	Woven	Glass/PP	N/A	(compression) Molding	1–10	Increase in ultimate tensile strength with decrease in void content, except for “extreme temperature condition” processing	253
		Glass/PI	[(0/90)]4	Prepreg–autoclave	0–10	Large effect of open voidage (large and irregularly shaped voids around the bundle intersections – intra-laminar) on the tensile strength; negligible effect of closed voidage (flat and semicircular voids between the prepregs – inter-laminar)	104
		Carbon/epoxy	[(±45)4/(0,90)/(±45)2]S [(±45)/04/(±45) (0,90)]S	prepreg–autoclave	0.4–9	Crack initiation from long voids in high-voidage specimens; low void sensitivity of tensile strength of woven composites	228
		Carbon/epoxy	[(±45)4/(0,90)/(±45)2]S	Prepreg–autoclave	0.3–1.5	FE analysis to predict the reduction in laminate tensile strength with increasing void content, using the improved Hashin failure criterion b	252
	NCF	Glass/vinyl ester (UD NCF)	Quasi-UD	RTM	0–4.5	Insignificant effect of voids on transverse tensile strength due to the presence of 5% weft fibers; increase in the strain to failure with increasing void content due to multiplication of large and small cracks	230
		Carbon/epoxy	QI	RTM	0.8–5	Negligible influence of voids on tensile strength	240
		Glass/epoxy	[0/+45/90/−45]S	VARTM	0–4.5	Negligible influence of voids on tensile strength	233

Concluding remarks on the effect of voids on tensile

Tables 6 to 8 make an overview of literature on the void effect on tensile properties. As far as stiffness is concerned, the voids affect stiffness of transverse and bias plies with few percent reduction per 1% increase of the void content. This leads to a small overall reduction of stiffness in cross-ply and QI laminates. Out-of-plane shear modulus can be substantially reduced because of void concentration at the inter-laminar interface.

Voids should affect strength of longitudinal plies because of the potential change in stress transfer and stress redistribution next to a broken fiber; there are experimental indications confirming this, but the phenomenon is not thoroughly studied.

Voids definitely affect strength of transverse and bias plies. The reduction in strength is as high as 10–20% per 1% of the void content increase. This reduction is much higher than what can be expected from the reduction of the effective cross-sectional area and is attributed to stress concentrations around the voids and easier crack initiation; these mechanisms are also revealed in the micro-level models. Transverse strength is related to the cracking of the transverse plies, which is the subject of the Transverse cracking section.

The reduced strength of transverse and bias plies is translated into a moderate decrease of the laminate tensile strength, few-percent reduction per 1% increase in void content. This is also true for woven laminates.

Compressive elastic properties

The compressive modulus, influenced by voids, has not been widely studied. In QI laminates, Hernández et al. ⁹⁷ found a small effect of voids on the compressive modulus, whereas Cinquin et al. ²⁶⁹ observed a linear reduction of the compressive modulus with increasing voidage for high void contents. In multidirectional NCF laminates produced with VARTM, according to Kosmann et al., ²³² the compressive modulus decreased with increasing void content in specimens with distributed voids. This reduction was more than what is expected from the reduced cross-sectional area following the rule of mixture. It was attributed to local fiber kinking in longitudinal bundles in the vicinity of voids and to early micro-cracking due to stress concentrations.

Compressive failure

Longitudinal compressive failure, or shortly compressive failure, of UD FRCs is controlled by many factors including constituent properties, interfacial effects and internal structure. Voids as a part of the internal structure can influence the compressive failure since they cause misalignment of the fibers in their vicinity as well as variations in the local shear behavior of the matrix. ²⁷⁰ It is worth noting that in a number of recent investigations of compressive strength, influenced by voids, modern methodologies such as in situ microscopy, micro-CT, photoelastic analysis, and digital image correlation (DIC) are employed that are explained in the current section.

Uhl et al. ⁸⁹ found the same monotonic trend in reduction of compressive strength with voidage as that of ILSS, for QI laminates with UD carbon/epoxy plies as well as for laminates with woven carbon/epoxy or carbon/PI plies. The degradation of the compressive strength with voidage was much less than that of ILSS, being attributed to the location and morphology of the voids. For a certain void content, the reduction of compressive strength was around two times higher in woven carbon/epoxy laminate than in QI carbon/epoxy laminate and in woven carbon/PI laminate. They fitted the results to the simple empirical exponential model of the literature, ²⁴⁵ explained in the ILSS in UD-ply composites section. Tang et al. ⁹² observed a ∼30% reduction in compressive strength by increasing the void content from 5% to 10%. The reduction rate increased with void content in their study, in contrast to the linear reduction in the literature. ⁸⁹

Bazhenov et al. ²⁷¹ found a square root relationship between the compressive strength and the void content in UD glass/epoxy composites. Their results showed that as the fiber diameter increases, the dominant failure mechanism changes from splitting (delamination) to micro-buckling, and the compressive strength becomes less void-sensitive. According to Lee and Soutis, ²¹⁴ thicker compression specimens have higher void contents, which can cause damage initiation through triggering fiber micro-buckling, leading to strength reduction. For a QI carbon/epoxy laminate, Cinquin et al. ²⁶⁹ observed that intra-laminar voids are detrimental for compressive strength, causing a reduction of ∼16% with increase of 6% in void content.

Hapke et al. ²⁷⁰ conducted an in situ analysis of the sequence of events prior to and during compressive failure in the presence of voids, as one of the pioneer studies performing more than post-mortem analysis on voids' effects. For a void-free UD carbon/epoxy specimen, the failure development started by matrix plastic shearing and continued by kink band formation (no micro-buckling) and fiber fracture. In the void-containing specimen, besides the overall void content, the void location relative to the developing kink band seemed to play a significant role. In the case of a kink band propagating along a path above the void tip, the band orientation altered due to the change in the lateral support. In the case where a void was located with its center in front of the notch, the kink band was suppressed by propagating through the void. In both location-cases, fiber fracture did not happen along the entire band length, while it always occurred in void-free specimens after the kink band initiation.

With the help of micro-CT, Hernández et al. ⁹⁷ could distinguish the effect of voids and of the stacking sequence on the compressive strength of QI carbon/epoxy laminates. Their compressive strength decreased with the void content increase and/or ply clustering. The effect of the latter is due to earlier onset of delaminations in laminates with thicker plies. The compressive strength after impact was also investigated and found to be insensitive to the voidage, which indicates that the damage produced during impact prevailed the effect of the voids.

In a number of recent studies, Liebig et al.^218,227,242 explored the influence of voids on the compressive failure in a model composite containing a single void laying between two UD glass fibers embedded in epoxy matrix. Through a photoelastic analysis, ²¹⁸ they observed that a high equivalent stress exists alongside the void, whereas close to the end caps of the void, the load is minimum (Figure 20). The length of the void had a vital influence on the compressive failure mechanism. The shorter the void, the smaller stress concentration appears alongside the void. The failure type for the case of a short/circular void was fiber/matrix debonding, while for the case of an elongated void, it was fiber buckling because of the only one-sided support for the fiber. Both types of failures happened in the zone influenced by the void where a high equivalent stress existed. The stress distribution together with the high stiffness mismatch between the fiber and matrix resulted in an overloading of the interface along the fiber close to the end cap of the void. This led to the interface debonding or fiber buckling. OPEN IN VIEWER

Later, they ²²⁷ performed a broader analysis on a model composite produced with discrete glass fibers as well as an extended model composite with glass rovings, both containing a single void. In the extended model composite, similarly to the model composite, the failure mechanism is: (1) fiber/matrix debonding, Euler buckling and failure of one fiber; (2) stress rearrangements, formation of a damage zone (3) failure of additional fibers, kink-band formation by the nearest fibers around the void; (4) propagation of the kink-band through the composite (Figure 21). Premature compressive failure of composites occurs in the presence of voids as an interaction between stress rearrangement alongside the void and fiber misalignment, which result in fiber/matrix debonding and local overloading of the fibers. This is followed by an ongoing loss in stability of the fibers, and finally fiber kinking. Through a numerical analysis with plasticity for the matrix and cohesive behavior for the interface, it was shown that the influence of fiber misalignment on the first fiber/matrix debonding is more significant than that of the aspect ratio of the void. Furthermore, in fractography of typical UD carbon/epoxy composite containing 8% of voidage, similar failure steps to those of the extended composite was observed. OPEN IN VIEWER

In a further study, ²⁴² the authors showed that DIC indicates also minimum and maximum strains in the summit and in the middle of the void, respectively. The high strain alongside the void was linked to the fiber deflection into the void, which accompanies the non-linear deformation of the matrix. Using an analytical approach, it was also concluded that besides fiber misalignment, the central angle of the fiber sector embedded in the matrix influences the compressive strength in the presence of voids. Therefore, to estimate the compressive strength, the critical number of affected fibers by voids needs to be determined through a probabilistic approach, in addition to the void content. Although in contrast to the approach presented by Hapke et al., ²⁷⁰ the methodology investigated by Liebig et al.^218,227,242 shows information also about failure initiation, it is still not representative of the compressive failure in real composites. This can be attributed to the lower fiber volume fraction, larger fiber diameter (∼70 µm), and larger size of voids (up to 3500 µm length and 800 µm height) compared with typical composites.

In NCF glass/epoxy composites produced through VARTM, Kosmann et al. ²³² observed that a significant reduction of compressive Young’s modulus and compressive strength takes place in the presence of voids, whereas they do not influence the strain to failure. Acoustic emission results showed that under compressive loads (distributed), voids could cause a shift in the matrix damage initiation to much lower strains in comparison to the case without voids. For woven carbon/epoxy composites, Zhang et al. ²⁵² numerically predicted the reduction of compressive strength as a function of the void content, using strength values for plies with different void contents as input.

Transverse compressive strength and its dependence on voids are less studied. Through numerical modeling, Ashouri Vajari ²⁶⁷ showed that 1% void content can decrease the strength by 15% in dominant transverse compression combined with longitudinal shear loading. In a rare study on through-thickness compressive strength of a cross-ply laminate with polyethylene fibers and a PU thermoplastic matrix produced in a hot press, O’Masta et al. ²⁰³ concluded that the load-carrying capacity in through-thickness compression mode is reduced greatly at the regions of high concentrations of voids. By accounting for load shielding within the high voidage regions (a simple statistical model), the large reduction in through-thickness compressive strength could be predicted.

Concluding remarks on the effect of voids on compressive properties

Table 9 makes a summary of literature treating the void effect on compressive properties, loading in the fiber direction. The main mechanism for the void influence on the compressive resistance is facilitating the fiber kinking/buckling and other instability modes because of the presence of free surfaces and free volumes introduced by the voids. The scale of voids is within the scale characteristic for the fiber kinking phenomenon. This leads to a reduction of the compressive strength by few percent per 1% of the void content increase. There exist a number of micro-level observations of this effect, as well as simple models of fibers kinking near a void. A full predictive model of the phenomenon, yet to appear, must account for void morphology and size distribution, and for void tendency to concentrate at the interplay interfaces.

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Table 9.

Effect of voids on the longitudinal compressive failure of different composites as reported in the literature.

Ply structure	Material	Lay-up	Production technique	Void content range (%)	Notable observation(s)	Reference
UD	Carbon/epoxy	QI	Prepreg	0.4–4.0	Monotonic (∼11%) reduction in compressive strength with (4%) voidage; fitting the results to the simple empirical exponential model, explained in the ILSS in UD-ply composites section; less degradation of compressive strength compared to ILSS	89
		UD	Prepreg–hot press	5–12	∼30% decrease in compressive strength with increasing the void content from 5% to 8–11%	92
		UD, QI	Prepreg–autoclave	N/A	Higher void contents in thicker specimens: reduction in compressive strength	214
		UD	Prepreg–autoclave	N/A	In situ analysis: depending on their location, voids can change the orientation of a propagating kink band or suppress the band propagation; fiber fracture did not occur along the entire band length in the presence of voids	270
		QI	Prepreg	0–11	∼16% reduction of compressive strength with 6% increase in intra-laminar void content	269
		QI	Prepreg–hot press	0.1–1.8	Degrading effect of voids on compressive strength, but not on compressive strength after impact a	97
	Glass/epoxy	UD rod	Pultrusion- and winding-type	1–4	A square root relationship between compressive strength and void content; change in failure mechanism from delamination to micro-buckling for high fiber diameters: less void sensitivity of compressive strength for high fiber diameters	271
		UD discrete fibers	Unique resin casting	N/A	Photoelastic analysis: high equivalent stress alongside the void; high effect of void length on compressive failure mechanism	218
		UD discrete fibers	Unique resin casting	N/A	Kink-band formation by the nearest fibers around the void; premature compressive failure of composites in the presence of voids; numerical analysis with voids, using plasticity for the matrix and cohesive behavior for the interface: significance of fiber misalignment +	227
		UD discrete fibers	Unique resin casting	N/A	Maximum and minimum strain in the summit and in the middle of the void, respectively; importance of the critical number of affected fibers by voids, determined through a probabilistic study, in estimating the compressive strength b	242
Woven	Carbon/PI	N/A	Prepreg	0–9.6	Monotonic (∼10%) reduction in compressive strength with (∼4%) voidage; fitting the results to the simple empirical exponential model, explained in the ILSS in UD-ply composites section; less degradation of compressive strength compared to ILSS	89
	Carbon/epoxy	N/A	prepreg	0–5.1	Monotonic (∼19%) reduction in compressive strength with (∼4%) voidage; fitting the results to the simple empirical exponential model, explained in the ILSS in UD-ply composites section; less degradation of compressive strength compared to ILSS	89
		[(±45)4/(0,90)/(±45)2]S	Prepreg–autoclave	0.3–1.5	FE analysis to predict the reduction in laminate compressive strength with increasing void content, using the improved Hashin failure criterion c	252
NCF	Glass/epoxy	[0,−45,90,45]S	VARTM	0–4.5	No significant effect of voids on compressive strength	232

Flexural elastic properties

Ghiorse ¹⁴⁰ observed a steady (in total 25%) decline in flexural modulus of a cross-ply carbon/epoxy composite with a void content increase from 0% to 5%. For a UD carbon/epoxy laminate, Olivier et al. ¹⁰³ concluded that the reduction in flexural modulus is highly dependent on the shape and size of the voids, in addition to their volume fraction. According to Liu et al., ¹⁰¹ in a cross-ply laminate, the flexural modulus decreases significantly with voids (∼10% decrease with 1% increase in voidage), which is not the case for the tensile modulus. The degradation of flexural modulus with voidage was also reported by Stamopoulos et al. ¹⁹⁸ for UD laminates, by Vergani et al. ²⁷² for multidirectional NCF laminates produced through RTM, and by Hou et al. ²⁷³ for thermoplastic matrix composites with woven plies, produced by compression molding. The latter attributed the modulus degradation to the decrease in the load transfer efficiency caused by voids. For cross-ply and random-chopped-strand composites of carbon/PP, produced with molding, Hayashi and Takahashi ²⁷⁴ observed, respectively, ∼3% and ∼11% decrease in flexural modulus with 1% increase in void content.

Hagstrand et al. ²⁷⁵ found a low level of degradation in a UD glass/PP laminate with low fiber volume fraction (35%), produced by compression molding. Fourteen percent of voidage led to 20% reduction in flexural modulus. They also pointed out that voids may produce positive effects on some structural properties of composites, in addition to lowering density and reducing manufacturing costs for the composite component. They, for example, can improve such structural properties as beam stiffness and flexural failure load if their positive effect on geometry (increasing the beam thickness) prevails their negative effect on material properties. This may happen only in special cases such as extremely high voidage and/or very low fiber volume fractions. Likewise, Yang and El-Hajjar ²³⁸ confirmed the enhancement of flexural rigidity by voids. They analytically calculated the bending stiffness matrix of a laminate using classical laminate theory, assuming layers of voids in between plies. The increase in bending stiffness was validated against experimental results. The studies about the effect of voids on flexural elastic properties are summarized in Table 10.

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Table 10.

Effect of voids on the flexural modulus of different composites as reported in the literature.

Ply structure	Material	Lay-up	Production technique	Void content range (%)	Notable observation(s)	Reference
UD	Carbon/epoxy	CP	Prepreg–autoclave	0–5	A steady (in total 25%) decline in flexural modulus with the void content increase from 0% to 5%	140
		UD	Prepreg–autoclave	0–10	High dependency of the reduction in flexural modulus on the shape and size of the voids, besides their volume fraction	103
		CP	Prepreg–autoclave	0–3.5	Significant reduction in flexural modulus with voids (∼10% decrease with 1% increase in voidage)	101
		QI	Prepreg–hot press	1–25	Enhancement of flexural rigidity (not flexural modulus) by voids	238
		UD	Prepreg–autoclave	0.8–3.4	Degradation of flexural modulus with voidage a	198
	Glass/PP	UD	Compression molding	1–14	Only 20% reduction in flexural modulus with 14% of voidage	275
	Carbon/PP	CP	Compression molding	0.9–3.2	∼3% decrease in flexural modulus with 1% increase in void content	274
Woven	Carbon/PEI	UD	Compression molding	0–6	Degradation of flexural modulus with voidage	273
NCF	Carbon/epoxy	QI	RTM	1.5–2.9	Degradation of flexural modulus with voidage	272
Chopped UD strands	Carbon/PP	Randomly oriented strands	Compression molding	0.9–2.1	∼11% decrease in flexural modulus with 1% increase in void content	274

Flexural failure

Ghiorse ¹⁴⁰ noted that the degradation of flexural strength in the presence of voids is as severe as that of ILSS. There is a steady decline of 10% in flexural strength with each 1% increase in void content, from 0% to 5%. Bowles and Frimpong ²²¹ found a non-linear relationship between the flexural strength and ILSS in a UD carbon/PI composite. They suggest that the change in the flexural failure mechanism is the reason for the non-linearity. As the void content increases, ILSS decreases, and the flexural failure mechanism changes from tensile failure in the tension side of the specimen to compressive failure in the compression side. On an example of a UD glass/PP laminate with low fiber volume fraction, Hagstrand et al. ²⁷⁵ argued that voids can also have a positive effect on structural properties like flexural failure load due to the increase in beam thickness. This may occur only for extremely high voidage and/or very low fiber volume fractions (explained in the Flexural elastic properties section).

The flexural failure evolution of a UD carbon/epoxy composite was characterized in the literature ¹⁵³ as: (1) matrix cracking (not originating from voids), compression damage underneath the roller; (2) linking of blunt intra-laminar and inter-laminar cracks by their propagation, which involves coalescence of large voids; (3) forming a through-thickness crack, final fracture on the tension side. With increasing void content, the mean flexural strength was found to decrease, while the standard deviation increased, indicating a dependency on the probability of voids of a critical size and shape being located in highly stressed regions. For a cross-ply carbon-epoxy laminate, Liu et al. ¹⁰¹ noted that the effect of voids on flexural strength is even more severe than that on ILSS. A reduction of ∼20% was observed with 2.5% voidage in a cross-ply carbon/epoxy laminate. For a UD laminate, Stamopoulos et al. ¹⁹⁸ reported ∼17% reduction of flexural strength for 3% increase in voidage. The flexural strength of Quickstep (OoA) cured panels was lower than that of autoclaved panels due to higher void contents, according to Davies et al. ¹¹⁶ The influence of voids on flexural strength of molded carbon/PP composites was investigated by Hayashi and Takahashi. ²⁷⁴ They found that with 1% increase in void content, the reduction of flexural strength is drastic (∼35%) for cross-ply laminates, while it is less significant (∼15%) for random-chopped-strand composites, attributed to the higher ratio of carbon fibers oriented in the longitudinal direction in the cross-ply composite.

The effect of voids on the residual flexural strength of UD carbon/epoxy composites after a high-temperature impact was investigated by Kakakasery et al. ¹⁰⁷ They concluded that the residual strength is slightly higher for a high-voidage laminate (e.g. the non-debulked one). This is because voidage may enhance the resistance to delamination, which is the predominant failure mode in post-impact flexural loading. Studying the residual flexural strength after low energy impact, Arthurs et al. ¹⁹⁹ observed that delaminations, caused by intra-laminar cracking, could be suppressed by a group of large voids, which blunt the crack propagation path. In specimens with a high voidage, a multi-stage process was identified, in which delaminations deflected into intra-laminar and inter-laminar regions and were blunted by voids. These phenomena are further discussed in the Inter-laminar fracture toughness section.

The fracture criterion (equation (10) in the Inter-laminar shear strength section) was used to fit the flexural strength of woven ¹⁵⁴ and UD ²²⁵ carbon/epoxy laminates. For a woven carbon/polyetherimide (PEI) laminate, Hou et al. ²⁷³ reported a 40% reduction of flexural strength with 6% of voidage, which, was linked to the reduction of the load transfer efficiency by voids. Zhang et al. ²⁵² numerically predicted the change in the laminate flexural strength with void content, using input data for plies with different void contents. The studies exploring the effect of voids on flexural failure are summarized in Table 11.

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Table 11.

Effect of voids on the flexural failure of different composites as reported in the literature.

Ply structure	Material	Lay-up	Production technique	Void content range (%)	Notable observation(s)	Reference
UD	Carbon/epoxy	CP	Prepreg–autoclave	0–5	∼10% reduction in flexural strength with each 1% increase in void content (as severe as ILSS degradation)	140
		UD	Prepreg–VBO	1–6	Coalescence of large voids during linking of blunt intra-laminar and inter-laminar cracks; decrease in the mean flexural strength, but increase in standard deviation, with increasing void content: the dependency on the probability of voids of a critical size and shape being located in highly stressed regions	153
		CP	Prepreg–autoclave	0–3.5	∼20% reduction in flexural strength with 2.5% increase in void content (more severe than ILSS degradation)	101
		UD	Prepreg–Quickstep, autoclave, oven cure	0.6–12.3	Lower flexural strength of OoA panels than autoclaved panels due to higher void content	116
		CP	Prepreg–autoclave	0–3.5	Using the fracture criterion (equation (10)) proposed in the literature 154 to correlate flexural strength with void content	225
		CP	Prepreg–autoclave	0.2–6.3	Slightly higher residual strength (after a high-temperature impact) for the high-voidage laminate: enhancement in resistance to delamination by voids	107
		CP	Prepreg–autoclave	0–5.1	Post-impact bending: arrest of delaminations by a group of large voids through blunting the propagation path a	199
		UD	Prepreg–autoclave	0.8–3.4	∼17% reduction in flexural strength with 3% increase in void content a	198
	Carbon/PI	UD	Prepreg–autoclave	1–10	Reduction of flexural strength with increasing void content; change in flexural failure mechanism from tension to compression with increasing void content	221
	Glass/PP	UD	Compression molding	1–14	Possibility of positive effect of voids on structural properties despite their negative effect on materials properties	275
	Carbon/PP	CP	Compression molding	0.9–3.2	Drastic (∼35%) reduction of flexural strength with 1% increase in void content	274
Woven	Carbon/epoxy	UD	Prepreg–autoclave	1.3–5.9	Fitting flexural strength versus void content with their modified fracture criterion (equation (10))	154
		[(±45)4/(0,90)/(±45)2]S	Prepreg–autoclave	0.3–1.5	FE analysis to predict the reduction in laminate flexural strength with increasing void content, using the improved Hashin failure criterion b	252
	Carbon/PEI	UD	Compression molding	0–6	∼40% reduction in flexural strength with 6% increase in void content	273
Chopped UD strands	Carbon/PP	Randomly oriented strands	Compression molding	0.9–2.1	Less significant (∼15%) reduction of flexural strength (compared to that in the cross-ply laminate) with 1% increase in void content	274

Concluding remarks on the effect of voids on flexural properties

Tables 10 and 11 review the literature on the void effect on flexural properties. The voids' influence on the flexural load-carrying ability of the composites is a combination of their influence on ILSS, and tensile and compressive strength. Complex combination of these mechanisms makes it difficult to generalize the void influence. However, the sources agree on decrease of the flexural stiffness by a few percent and of the flexural strength by up to 10% per 1% of the void content increase.

Concluding remarks on the effect of voids on transverse cracking

The studies, in which the effect of voids on transverse cracking is investigated, are reviewed in Table 12. The overall effect of the void presence is a shift of the crack initiation load/strain, which can be decreased by a factor of 2–3. Some studies found the crack saturation density insensitive to voids, while some other investigations revealed a significant increase of the saturation density in the laminates with voids. The earlier crack initiation is caused by the stress/strain concentration around the voids. This facilitates the crack initiation, which nevertheless continues to propagate in the same material, as in the absence of voids. However, details of the phenomena are still under active investigation. This asks for reliable and precise tools for characterization of matrix cracking dynamics. Moreover, models of the cracking in the presence of the voids distributed in the ply are available, whilst models accounting for voids at interfaces are still to appear. The scale issue is an additional reason for the overall limited number of computational works in this area. Voids are defects that have microscopic features and interact with other micro-heterogeneities like fibers. Prediction of their effects on statistically controlled properties, such as transverse cracking, requires a multi-scale modeling approach and such approaches are still rare.

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Table 12.

Effect of voids on transverse cracking in different composites as reported in the literature.

Ply structure	Material	Lay-up	Production technique	Void content range (%)	Notable observation(s)	Reference
UD	Glass/vinyl ester (UD NCF)	Quasi-UD	RTM	0–4.5	In laminates with higher void content: larger maximum crack density, lower strength, higher strain to failure	230
	Polymer–matrix composite	CP	Prepreg–autoclave	N/A	Formation of few initial transverse cracks from voids; reduction in effect of voids on further cracking with proceeding of loading	276
	Carbon/epoxy	CP	Prepreg–autoclave	N/A	Random formation of initial transverse cracks from random voids; more uniform distribution with proceeding of loading Statistical analysis: w.r.t. the degradation by defects, the laminates are ranked as: with vacuum-with pressure, with vacuum-no pressure, no vacuum-with pressure, and no vacuum-no pressure	6
					Occurrence of transverse cracks in the inner plies of the voidy material, not present in the reference material; earlier cracking initiation in the outer plies of the voidy material, compared to the reference material; irregular crack propagation through the specimen’s width in the voidy material a	277
					Prediction of the earlier initiation of transverse cracking via a combined-scale methodology, using computational micromechanics and extended finite element method b	278
		[902/05]S	Prepreg–autoclave	0–4.6	Strain concentrations caused by voids, leading to local plastic deformation of the matrix at lower global strains, causing earlier initial failure a Post-mortem micro-CT: transverse cracks running through voids c	201
Woven	Carbon/epoxy carbon/ bismaleimide	UD	Prepreg–autoclave	0.5–5.6	Initiation of transverse cracks from voids	155
NCF	Carbon/epoxy	QI	RTM	0.8–5	Void sensitivity of the crack density only at low strains; less effect of voids on saturated crack density; higher effect of voids on transverse cracking in 90° plies than in 45° plies	240
		CP, [(±45)3]S	RTM	1.7–4.5	Initiation of transverse cracks at the void locations; high void content specimen: low strain to first crack, high crack density; a good correlation between the applied strain to crack initiation and the void strain concentration factor; inadequacy of void content b	231

Concluding remarks on the effect of voids on inter-laminar fracture toughness

The literature where the effect of voids on inter-laminar fracture toughness has been studied is summarized in Tables 13 and 14. In the presence of voids, there are competing phenomena, which on one hand, facilitate the crack initiation and propagation, providing readily available free surfaces for the crack, and on the other hand, hinder the crack, increasing its surface, blunting the crack tip, and promoting fiber bridging. Due to the complex interplay between different mechanisms, no consensus could be made on the effect of voids on the inter-laminar fracture toughness.

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Table 13.

Effect of voids on the inter-laminar fracture toughness of different composites as reported in the literature.

Ply structure	Material	Lay-up	Production technique	Void content range (%)	Fracture mode	Notable observation(s)	Reference
UD	Carbon/epoxy	[0°12//(±5°/0°4)S]	Prepreg–autoclave	0–0.9	I, II, mixed	Fracture toughness at propagation in the presence of voids: strong increase in Mode I fracture toughness due to change in failure mechanism, slight decrease in Mode II fracture toughness, increase in mixed mode fracture toughness due to multiple cracking at the crack tip; either deleterious or no effects on the fracture toughness at crack initiation, in all three modes	279
		CP	Prepreg	discrete voids	I	Increase in Mode I fracture toughness when elliptic–cylindrical sensor cavities are aligned normal to the direction of crack growth due to crack tip blunting and the increased size of the plastic zone at the crack tip	247
		UD	Prepreg–autoclave	0.1–15.3	I	Decrease of 22% in Mode I inter-laminar fracture toughness with 5% increase in void content	243
Woven	Carbon/PEI	UD	Compression molding	0–6	I, II	In specimens with higher voidage: higher Mode I fracture toughness at propagation due to multiple cracking as a result of fiber/bundle bridging caused by the inter-laminar voids, lower Mode II fracture toughness at propagation, and lower Mode I and Mode II fracture toughness at initiation	273
	Glass/polyester	N/A	Hand lay-up +  cold-cured	3–30	I	Rapidly reduction in Mode I fracture toughness at both crack initiation and crack propagation with increasing voidage up to ∼15% due to facilitating the initiation and growth of the delamination instead of causing the delamination to jump between plies	157
NCF	Carbon/epoxy	CP, [(±45)3]S	RTM	1.7–4.5	I	A slightly higher Mode I inter-laminar fracture toughness (both for crack initiation and propagation) for laminates with higher voidage due to diffused damage, in particular diffused fiber bridging	231

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Table 14.

Effect of voids on inter-laminar fracture toughness (increase or decrease) in different modes as reported in the literature.

Ply structure	Material	Lay-up	Production technique	Void content range (%)	Mode I at initiation	Mode II at initiation	Mode I at propagation	Mode II at propagation	Reference
UD	Carbon/epoxy	[0°12//(±5°/0°4)S]	Prepreg–autoclave	0–0.9	↑	↑	↑	↓	279
		CP	Prepreg	Discrete voids	↑	–	–	–	246
		UD	Prepreg–autoclave	0.1–15.3	↓	–	–	–	243
Woven	Carbon/PEI	UD	Compression molding	0–6	↓	↓	↑	↓	273
	Glass/polyester	N/A	Hand lay-up+cold-cured	3–30	↓	–	↓	–	157
NCF	Carbon/epoxy	CP, [(±45)3]S	RTM	1.7–4.5	↑	–	↑	–	231

↑: increase; ↓: decrease; PEI: polyetherimide; UD: unidirectional; CP: cross-ply; N/A: not available/applicable; NCF: non-crimp fabric; RTM: resin transfer molding.

Concluding remarks on the effect of voids on fatigue behavior

An overview of the research on the effect of voids on fatigue of FRCs is presented in Table 16. Different types of axial, flexural, and inter-laminar fatigue performance are investigated. Typically, fatigue loading acts as an amplifier for the void influence on the composite strength. This is especially visible for the case of loading in the fiber direction: If influence of the voidage on the static strength is within few percent, the fatigue life can deteriorate drastically, with number of cycles to failure decreasing up to 10 times for 2% of the void content increase. However, this amplification mostly happens at low- and medium-cycle fatigue. Literature links this phenomenon with the void influence on transverse cracking (see the Transverse cracking section), which is important for the crack initiation (low-cycle fatigue) and less important for the crack saturation (high-cycle fatigue). Application of micro-CT reveals that investigation of the criticality of void characteristics is necessary for a precise analysis of the fatigue performance.

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Table 16.

Effect of voids on the fatigue behavior of different composites as reported in the literature.

Cyclic loading	Ply structure	Material	Lay-up	Production technique	Void content range (%)	Notable observation(s)	Reference
Axial	UD	Carbon/epoxy	UD	Compressional molding	0.2–1.8	Tension–compression: local heating by voids during cyclic loading resulting in matrix softening, and thus local buckling of the adjacent fibers; voids can trigger crack initiation as they are potential heat accumulation sites	280
			[0/+45/−45]3S	Prepreg–VBO	2.8–3.7	Tension–compression: finding a correlation between fatigue life and the volume of the largest single void in the critical damage layers a	195
		Glass/epoxy	CP, [0/452/0/−452]S	VARTM	0.3–6.7	Tension–tension: dramatic influence of voids on life to crack initiation and crack growth rate in both lay-ups; influence of voids on crack density in the 90°-ply only in the laminate with homogeneous distribution of voids	223
	Woven	Glass/epoxy	UD	Prepreg	0.1–2.7	Tension–tension: higher drop in fatigue strength than that in static (tensile) strength: attributed to the stress concentration around the voids b	145
	NCF	Glass/epoxy	N/A	VARTM	0–8.8	Tension–tension: a high crack density already in the early stages of fatigue life due to the voids resulting in a shorter fatigue life Compression–-compression: less significance of voids in degradation of fatigue life; premature kinking initiation due to voids Tension–compression: extension of fatigue life by voids, possibly due to the drastic decrease in inter-laminar cracking	241
			QI	Filament winding+RTM	1–2	Biaxial tension–compression: higher modulus degradation for specimens with voids; shorter fatigue life of the specimens with large accumulated voids due to early formation of delaminations, resulting in catastrophic failure, and of the specimens with distributed small voids due to a hot-spot, produced by critical delaminations, causing the final failure	172
			[0/+45/90/−45]S	VARTM	0–4.5	Tension–compression: no sensitivity of fatigue life and stiffness degradation to void content	233
			[45/−45/0]S	Infusion molding	0–4.5	Tension–tension: a lower number of cycles for the initiation of off-axis cracks in the presence of voids; increase in micro-scale damage evolution rate by voids b	219
		Carbon/epoxy	[0/−45/90/+45]2S	RTM	0.8–20	Tension–tension: degradation of fatigue life for ∼3 orders of magnitudes by 20% of void content; significance of voids in micro-crack initiation in 90° plies, but small effect in 45° plies; higher crack densities for higher void contents only at lower number of cycles c	240
Flexural	UD	Carbon/epoxy	UD	Prepreg–VBO	1–6	The same order of degradation of fatigue performance as that of flexural strength; existence of a critical level, not only for void content but also for other void characteristics, i.e. shape, size, and location: significance of large and inter-laminar voids	153
	Woven	Carbon/epoxy	UD	Prepreg–autoclave	1.3–5.9	More severe degradation of fatigue life due to voids than that of static strength; existence of a critical void content in degradation of fatigue life	154
		Glass/pp	N/A	(compression) Molding	1–10	Explaining the effect of voids on fatigue behavior by its dependency on flexural strength and ILSS through Basquin’s equation	253
Inter-laminar	UD	Carbon/epoxy	UD	Prepreg–autoclave	N/A	Inter-laminar tensile fatigue: reducing the scatter in the curved-beam fatigue test data through transferring the micro-CT data, including critical voids, to the FE model; obtaining a reliable inter-laminar tensile S–N curvea,b	197
	NCF	Carbon/epoxy	CP, [(±45)3]S	RTM	1.7–4.5	Inter-laminar (double cantilever beam) fatigue: similar Paris-like curves of specimens with and without voids; independency of crack growth rate of voids b	231

Concluding remarks on hygrothermal effect on mechanical properties

A summary of the research on the effect of voids on hygrothermal mechanical behavior of FRCs is presented in Table 17. In general, the rate and equilibrium level of moisture absorption of FRCs depends on voids. The moisture diffusion becomes non-Fickian in the presence of a high level of voidage. Moisture and voidage can have synergistic effects on mechanical properties. Hygrothermal conditioning affects the same properties as the voids do. Therefore, the degrading effect of voids on mechanical properties can be much higher in the presence of moisture and heat, particularly important for composites that will be used in humid and/or warm environments. However, the effect of hygrothermal conditioning on the strength properties influenced by voids depends on the reinforcement structure, fiber and matrix material, and loading mode.

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Table 17.

Influence of voids on the hygrothermal effects in different composites as reported in the literature.

Ply structure	Material	Lay-up	Production technique	Void content range (%)	Notable observation(s)	Reference
UD	Carbon/epoxy	UD	Prepreg	1–5.5	Dependence of rate and the equilibrium level of moisture absorption (saturation moisture content) on void content; some non-Fickian diffusion anomalies for high-voidage specimens; possible synergistic effects of voids and moisture on mechanical properties	254
		UD, CP	Prepreg–autoclave, prepreg–hot press	0–5.6	Non-Fickian diffusion behavior in the presence of high level of voidage a	281
		UD	Prepreg–autoclave	0.2–4.2	Hygrothermal conditioning affects the same properties as the void content does; moisture absorption influence the laminate strength through change in the matrix toughness caused by matrix plasticization; dependency of the effect of hygrothermal conditions on the strength properties influenced by voids on the reinforcement structure, fiber and matrix material, and loading mode; more sensitivity of ILSS in the presence of voids to hygrothermal conditioning for woven carbon/epoxy composites than for woven carbon/bismaleimide and UD carbon/epoxy composites; more sensitivity of compressive strength in the presence of voids to hygrothermal conditioning for woven carbon/bismaleimide composites	251
		±55°	N/A	0–5	Self-consistent model results: decrease in the hygroscopic expansion coefficient with increase in void content, unlike the transverse effective diffusion coefficient; strong interaction between voidage and humidity a	282
	Glass/epoxy	UD rod	Pultrusion-type technique	0.5–6	Dependence of rate and the equilibrium level of moisture absorption (saturation moisture content) on void content	248
Woven	Kevlar/epoxy	QI	Prepreg–autoclave	1–15	Decrease in flexural strength with voids only at high temperatures; decrease in flexural strength of the high voidage specimen containing moisture with freeze-thaw cycling	283
	Carbon/epoxy	UD	Prepreg–autoclave	1.1–10.5	Less sensitivity of ILSS to hygrothermal conditioning for higher void contents; less sensitivity of ILSS to void content at higher temperatures	284
		UD	Prepreg–autoclave	0.2–4.2	Explained above	251
		[(±45)4/(0,90)/(±45)2]S	Prepreg–autoclave	0.3–1.5	Decrease in strength, particularly ILSS and compressive strength with increase in void content and/or immersion time; specimens sorted w.r.t. ILSS and compressive strength: aged, dried, and unaged; some effects of hygrothermal conditioning are recoverable through drying; specimens sorted w.r.t. the bending strength: unaged, aged, and dried	229
	Carbon/ bismaleimide	UD	Prepreg–autoclave	0.2–4.2	Explained above	251