# In-situ open-hole tensile testing and modeling of hybrid PEEK thermoplastic laminates under burn-through kerosene flame exposure

Lanhui Lin, Benoit Vieille, Christophe Bouvet, Tanguy Davin

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Lanhui Lin, Benoit Vieille, Christophe Bouvet, Tanguy Davin. In-situ open-hole tensile testing and modeling of hybrid PEEK thermoplastic laminates under burn-through kerosene flame exposure. Composites Part C: Open Access, 2025, 18, pp.100657. 10.1016/j.jcomc.2025.100657. hal-05341723

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Submitted on 1 Nov 2025

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# In-situ open-hole tensile testing and modeling of hybrid PEEK thermoplastic laminates under burn-through kerosene flame exposure

Lanhui Lin<sup>a, b</sup>, Benoit Vieille<sup>a</sup>, Christophe Bouvet<sup>b, c</sup>, Tanguy Davin<sup>a</sup>

## ARTICLE INFO

**Keywords:**

In-situ fire test

Kerosene flame

Thermoplastic composite

Open-hole

## ABSTRACT

The objective of the present work is to investigate the thermo-mechanical behavior of open-hole hybrid carbon/glass fiber reinforced PolyEther Ether Ketone (CG/PEEK) thermoplastic laminate subjected to the kerosene flame exposure (1100 °C and 116 kW/m² heat flux) in combination with tensile loading. A specialized flame testing bench has been developed, integrating a tensile mechanical loading and a kerosene burner, to induce in-situ fire-mechanical test conditions. The novel prototype has been employed to monitor the temporal evolution of several physical quantities in the range of fire exposure times up to 900 s, including back surface and through thickness temperature, open-hole deformation and swelling ratio of thickness. The mechanisms of fire- and mechanically-induced damage are examined through the fractographic analysis using tomography and microscopy. One-sided burn-through flame exposure causes the in-plane (4.0 K/mm) and through-thickness (40.1 K/mm) temperature gradients after 300 s. Compared to the virgin state, there is a considerable reduction in the equivalent stiffness (-67%) and axial strength (-55%) following a 900 s of flame exposure, indicating the severely damaged structural integrity. The modeling of the in-situ mechanical properties over multiple phase transition temperatures of the PEEK matrix is applied to characterize and ultimately predict the thermo-mechanical response of laminate under tensile loading in fire. The approach is based on the experimental measurement of mechanical properties over a wide temperature range (isothermal heating from the glass transition temperature to the PEEK matrix pyrolysis). The model shows a high degree of effectiveness in representing the in-situ open-hole tensile behavior of TP-based laminates under fire conditions as a function of flame exposure time.

## 1. Introduction

The use of composite materials in the structures of latest generation aircraft is of increasing prevalence [1]. The flammability and rapid thermal decomposition of most thermoset (TS) matrices have been the driving force behind the emergence of high-performance thermoplastic (TP) composites in the aerospace industry over the last fifteen years [2]. Indeed, the fact that composites are joined to other structures and contain open cut-outs for access requires the drilling of laminates [3,4]. In the case of notched laminates exposed to one-sided fire exposure, the presence of holes provides open access for the heat flow, thereby potentially leading to the in-plane degradation of the laminates [5,6]. However, the issue of multi-physics couplings within Polymer Matrix Composites (PMCs) under severe heat exposure (such as the flame exposure of aircraft engines) remains open [7,8].

With regard to the stringent fire safety regulations [2], it is thus essential to understand and predict the thermo-mechanical behavior of composite materials, and ultimately of structures and assemblies [9]. This calls for the development of adapted experimental characterization means as well as modeling approaches of the fire resistance of laminates under load. To date, the predominant focus has been on the TS matrix laminates [7,10], while there is a lack of references addressing the TP-based composites under critical service conditions [11,12]. Therefore, a novel bench-scale fire test, featuring a combination of tensile loading and kerosene burner, has been specifically designed to induce in-situ fire-mechanical stress conditions [13]. Using experimental and modeling methods, the present study aims to investigate the thermo-mechanical behavior of open-hole CG/PEEK laminates subjected to combined kerosene flame exposure (116 kW/m² and 1100 °C [13]) and tensile loading.

<footnote><sup>a</sup> Groupe de Physique des Matériaux (UMR CNRS 6634), INSA Rouen Normandie, Univ Rouen Normandie, CNRS, Normandie Univ, 685 Av. Université, F-76801, Saint-Etienne-du-Rouvray, France</footnote>

<footnote><sup>b</sup> Institut Clément Ader (UMR CNRS 5312), Université de Toulouse, ISAE-SUPAERO – IMT Mines Albi – INSA, 10 Av. M. Pélegrin, Cedex 4, F-31055, Toulouse France</footnote>

<footnote>* Corresponding author.</footnote>

<footnote>E-mail address: christophe.bouvet@isae-supaero.fr (C. Bouvet).</footnote>
## 1.1. Thermo-mechanical coupling of PMCs in fire

When Polymer Matrix Composites (PMCs) are exposed to a one-sided impinging fire, the temperature of each ply within the laminate (from the exposed surface to the back surface) gradually increases with time, resulting in a series of chemical processes involved in the pyrolysis of the polymer matrix [14], at which this process occurs is known as the temperature at the onset of the thermal decomposition ($T_d$) [15]. The knowledge of their characteristic temperatures is therefore fundamental to understand the changes in the tensile behavior and properties in fire [16]: $T_g$ the glass transition temperature, $T_m$ the melting temperature (into semi-crystalline material [10]) and $T_d$ refers to polymer matrix pyrolysis (5 % of mass loss). Fig. 1 illustrates the different thermal responses of PMCs as a function of the temperature reached in different plies [7]. In general, the high-temperature thermo-mechanical coupling in PMCs can be analyzed considering different test conditions above or below $T_d$ [15]. In the former case ($T < T_d$), as the polymer matrix heats up and then viscously softens, high temperature gradients and different thermal expansion coefficients (fibers and matrix) can promote the formation of matrix cracking and porosity [17]. In the latter case ($T > T_d$), large amounts of pyrolysis gas and smoke are released, along with the formation, growth and oxidation of char, which is a solid carbonaceous residue [18]. The carbon-rich char layer can insulate from the heat flow [8], thus presents a lower effective thermal conductivity, which reduces the through-thickness flame spread within the laminate over extended exposure time [17,19]. PMCs typically contain a high proportion of synthetic polymer [15], and generally behave as a carbonizing/charring material [10], depending on the fiber-to-resin ratio.

The thermal degradation mechanisms in PMCs have been extensively investigated for many years [15,16,18]. However, the literature on the fire behavior of composites has focused primarily on TS-based composites [2,20,21] and less on TP-based composites [6,19]. Most of the results available in the literature on the thermo-mechanical behavior of TP-based laminates are derived from thermal aggression in the cone calorimeter [10,22,23]. Among high-performance thermoplastic resins, the preliminary studies [8,17,19] were conducted on CG/PEEK laminated composites. Thermogravimetric analysis (TGA) results show that PEEK resin has a high melting temperature (350 °C) and good thermal stability up to 450 °C [19]. Under isothermal conditions (at 500 °C), the in-situ tensile properties are significantly reduced [8], with -38 % in axial stiffness and -54 % in axial strength, compared to those obtained at room temperature. Under a kerosene flame exposure, it appears that the boundary conditions have a significant influence on the creep behavior [8,17]. It is noteworthy that an increase in temperature is effective to accommodate the overstresses induced by the open hole [3], leading to lower hole-sensitivity under elevated temperature exposure [5]. The work [8] reported that the back-surface and through-thickness temperature measurements serve as an important point to investigate the thermo-mechanical coupling under severe thermal stress, through which the in-situ tensile behavior is expected to be predicted as a function of flame exposure time [16].

<figure><img src="image_2.png" /><figcaption>Fig. 1. Schematic presentation of thermal behaviors of PMCs as function of exposure temperature ranging from $T_g$ to $T_d$ (adapted from [7]).</figcaption></figure>

## 1.2. Property modeling over phase transition temperatures in polymers

Modeling of the in-situ properties of PMCs can be divided into the modeling of the thermal and mechanical responses [7]. The former accounts for the thermal degradation as a function of ply (induced by temperature gradient) [24], and the latter considers structural integrity [25], which is highly dependent on the degree of degradation (correlated with temperature evolution) [16]. Ideally, any mechanical property values (such as the stiffness or strength) would be measured and expressed as a function of the heating temperature $P(T)$ [10], as illustrated in Fig. 2(a). Under one-sided flame exposure, a composite laminate can be regarded as a composition of materials in various states (see Fig. 1), exhibiting different thermo-mechanical behaviors [26], as shown in Fig. 2(b). Based on the modeling of temperature-dependent mechanical properties and the estimation of thermal responses (temperature profiles), it is then possible to simulate the integrity of laminates subjected to a one-sided flame under tensile load [27]. Therefore, the issue at hand is the identification of a suitable curve to represent the relationship between in-situ properties and isothermal heating temperatures across multiple phase transition temperatures [28]. This can be obtained by performing the isothermal mechanical property measurement [16,25], which provides the values of the mechanical properties at different characteristic temperatures of polymer matrix.

Several functions with the required antisymmetric feature (with respect to $T_R$ in Fig. 2a) have been proposed and investigated with different types of polymers [7,10]. One of the first models to account for this behavior was developed by Mahieux et al. [25,28], where the temperature-dependent property $P(T)$ is assumed to follow a cumulative Weibull distribution. This simple functional relationship is given in the following Eq. (1):

$$P(T) = P_R + (P_U - P_R) \exp \left[ - \left( \frac{T}{T_R} \right)^m \right] \quad (1)$$

where $P_U$ and $P_R$ are the unrelaxed (low temperature) and relaxed (high temperature) property values respectively, $T_R$ is the relaxation temperature (in Kelvin), where the property has been reduced by 50 %, as indicated in Fig. 2(a), and $m$ is the Weibull exponent. Based on the hyperbolic $tanh$ function instead, Gibson et al. [21,27] proposed the property-temperature curves in the following form:

$$P(T) = P_U - \frac{P_U - P_R}{2} [1 + \tanh(k(T - T_R))] \quad (2)$$

Where $k$ is a fitting constant describing the span of the relaxation region with a corresponding $T_R$. The application of this equation has been found to be suitable to fit a functional relationship with temperature and generally shows a good prediction results to the strength data [21].

It is worth noting that for TS-based laminates, $T_R$ refers to the mechanical glass transition temperature $T_g'$ [27], which is actually different from the $T_g$ value measured by thermal techniques, such as DSC. The measured temperature of the latter is typically about 20 °C higher than that of the former [10]. Moreover, in contrast to TS-based laminates, the TP resin undergoes not one, but two phase transitions before the decomposition of resin [28]: with a second crystalline transition that occurs prior to the temperature of matrix melting $T_m$ (noted as $T_m'$). In the TP-based composites case, therefore, an additional relaxation process is added to the curve-fitting model [21], so that Eq. (2) was generalized to include an extended term:

$$P(T) = P_1 - \frac{P_1 - P_2}{2} [1 + \tanh(k_1(T - T_2))] \\ - \frac{P_2 - P_3}{2} [1 + \tanh(k_2(T - T_3))] \quad (3)$$

Where the subscripts 1, 2 and 3 refer to the low, intermediate, and high temperature states, corresponding to room temperature (RT), glass transition temperature ($T_g'$) and crystalline transition temperature
<figure><img src="image_3.png" /><figcaption>Fig. 2. Schematic illustration of the modeling method: (a) The effect of isothermal heating temperature on the mechanical properties (adapted from [10]) – (b) The stress-strain curve shape of a parametric model (Eq. (5)) with increasing heating temperatures (adapted from [21]).</figcaption></figure>

(T_m'), respectively [27]. Furthermore, Mahieux et al. [28] developed an extending model to adapt the three transitions for the TP resin, up to the resin decomposition onset (noted as T_d'). The multiple transitions of TP-based laminates can be described as follow:

$$P(T) = (P_1 - P_2)\exp\left[-\left(\frac{T}{T_1}\right)^{m_1}\right] + (P_2 - P_3)\exp\left[-\left(\frac{T}{T_2}\right)^{m_2}\right] + P_3\exp\left[-\left(\frac{T}{T_3}\right)^{m_3}\right] \quad (4)$$

Where the subscripts i = 1, 2, 3 refer to different phase transitions (T_g', T_m' and T_d' respectively), P_i is the instantaneous property in each region before the phase transition, and T_i is the phase transition temperature. The values of P_i and T_i can be obtained from the isothermal heating tests, the principal focus is thus the statistical parameters m_i. The proposed model was found to well describe the behavior of four types of commercial TP-based composites (PEEK, PMMA, PPS and AS4/PPS) over the entire application temperature range [28]. Subsequent to the acquisition of the stiffness and strength values, the stress-strain curves can be modeled [21], as depicted in Fig. 2(b), employing the following empirical nonlinear relationship:

$$\sigma = \sigma_{max} \left[ 1 - \exp\left(\frac{-E\epsilon}{\sigma_{max}}\right) \right] \quad (5)$$

Where E is the Young's modulus in low strain region, σ_max is the asymptotic yield stress, which is used to describe the increasing nonlinearity under elevated temperature, and σ_f is the ultimate stress at failure in Fig. 2(b). The effectiveness of the model was validated by the high agreement between model predictions and experimental results [21,27]. In the present study, it is intended to implement the laminate analysis modeling to CG/PEEK laminates, with the objective of predicting the open-hole tensile response under fire conditions as a function of flame exposure time. Nevertheless, the thermal modeling will not be addressed in this study. Instead, the experimental through-thickness temperature profiles will be employed as input data.

## 1.3. Objectives

The presence of holes in notched laminates provides open access for heat flux, inducing detrimental thermal gradients when exposed to fire. This scenario would constitute more severe conditions with regard to the design of assemblies. The present work thus aims at characterizing and modeling the thermo-mechanical behavior of the open-hole CG/PEEK laminates subjected to combined burn-through kerosene flame exposure (1100 °C and 116 kW/m² [13]) and tensile loading. The objective is

threefold: (i) to determine the thermal responses (back surface and through-thickness temperature profiles) and the in-situ open-hole tensile properties as a function of flame exposure time – (ii) to examine the effect of heating temperature on the changes in the tensile properties (stiffness and strength) under isothermal testing conditions – (iii) to represent the in-situ open-hole tensile behavior of CG/PEEK laminates, using the proposed modeling method over multiple phase transition states (up to the decomposition of resin). It is then possible to provide a practical approach to assess the load-bearing behavior of composites in fire scenario.

# 2. Materials and experimental set-up

## 2.1. Materials and specimens

The composite materials considered in this study are CG/PEEK laminates consisting of a PolyEther Ether Ketone (PEEK) thermoplastic matrix reinforced with a continuous carbon fiber fabric (Tenax®-E HTA40 3 K), and two outer plies of Glass/PEEK (used to provide electrical protection, and avoid the galvanic corrosion of aluminum and steel with carbon), as shown in Fig. 3. The laminates have 16 plies (average thickness is about 4.5 mm) with a layup of [(0/90)G, (0/90), (±45),(0/90),(±45),(0/90),(±45),(0/90)]s. The PEEK matrix characteristic temperatures (derived from the material technical data sheet and Ref. [19]) are given in Table 1, where T_g, T_m and T_d are, the temperature of glass transition, melting and decomposition respectively.

## 2.2. Experimental methods

### 2.2.1. Tensile mechanical testing under kerosene flame exposure

To provide in-situ tensile testing conditions under fire exposure, a home-made fire bench was used. The kerosene burner bench shown in Fig. 4 is a domestic device (Cuenod manufacturer). The heat flux and flame temperature of the burner were measured in Ref. [13] to ensure standard values (1100 ± 80 °C kerosene flame with a heat flux of 116 ± 10 kW/m² [2]) were met. In agreement with certification requirements [2], a maximum 900 s of fire exposure was applied. Three groups of open-hole tensile specimens (three specimens each group) were exposed to a kerosene flame exposure with an insulation layer for protective shielding, which was used to limit the thermal decomposition area (20 × 40 mm² as shown in Fig. 3b) of the specimen along the edges (see left part of Fig. 4). These tests consist of two stages:
<figure><img src="image_4.png" /><figcaption>Fig. 3. CG/PEEK material: (a) Laminate lay-up – (b) Open-hole tensile specimen dimension (in mm).</figcaption></figure>

<table><caption>**Table 1**

Characteristic temperatures of CG/PEEK laminates [19].</caption><thead><tr><th>T<sub>g</sub> / °C</th><th>T<sub>m</sub> / °C</th><th>T<sub>d</sub> / °C</th></tr></thead><tbody><tr><td>143</td><td>350</td><td>580</td></tr></tbody></table>

<figure><img src="image_5.png" /><figcaption>Fig. 4. Experimental set-up schematic of open-hole CG/PEEK laminates subjected to a kerosene flame during mechanical loading with an insulation layer for protective shielding.</figcaption></figure>

* A creep loading of 0.5 ± 0.05 kN during the flame exposure (300 s – 600 s – 900 s) to prevent the buckling of the thermally degraded specimen;

* A force-controlled tensile loading (0.5 kN/s) until failure of the specimen.

During the test, mechanical load was measured by a 100 kN load cell and the axial displacement was measured using a Digital Image Correlation (DIC) technique with a CCD camera (right part of Fig. 4), which will be further described in Section 2.2.2. InfraRed Thermography (IRT) was implemented to monitor the back surface (angle 45°) and through-thickness (angle 0°) 2D temperature field, which will be further described in Section 2.2.3. A fume extractor was used to reduce the detrimental effect of flame combustion smoke on the camera images' monitoring.

According to the European standard EN 6035 of Open-Hole Tensile (OHT) test [29], the equivalent axial stiffness *E* (with regard to unnotched specimen) and the ultimate open-hole tensile strength *σ*<sub>f</sub> were calculated as follows:

$$E = \frac{\Delta F}{S_0 \cdot \Delta \varepsilon_x} \quad \text{and} \quad \sigma_f = \frac{F^u}{S_0} \quad (6)$$

Where Δ*F* is the difference in the tensile loads corresponding to *ε*<sub>1</sub> =

0.05 % and ε<sub>2</sub> = 0.25 %, S<sub>0</sub> is the specimen initial cross section (disregarding the open hole), Δε = ε<sub>2</sub> - ε<sub>1</sub> is the difference in the axial strains obtained from the Digital Image Correlation (DIC) technique (described in the next section), F<sup>u</sup> is the maximum force borne by the open-hole specimen at failure.

### 2.2.2. Digital image correlation (DIC) technique

A two-dimension DIC technique was applied on the back surface of specimens during tensile loading under flame exposure. The DIC pattern is painted on the specimens' back surface (out of the fire exposure zone, to avoid influencing on the temperature measurement) with a non-periodic pattern and without preferred orientation, as shown in Fig. 5 (a). A Grasshopper® high-speed monochrome camera enables the capture of digital images at full resolution (1920 × 1200 pixels), and at a rate of 5 fps. During the fire exposure, the implementation of DIC technique requires the usage of a high intensity light source conducted by an optic fiber to illuminate the patterns, and to mitigate the perturbation of fume. Ultimately, the strain measurement of fire exposure zone is derived from the 2D displacement field by means of the VIC-2D correlation software, which is provided by Correlated Solutions Company, as shown in Fig. 5(b). The gauge length of contactless extensometer is set at 70 mm, as the back-surface color within the exposed zone (Fig. 7) turns completely black (due to thermal degradation) after a 900 s of flame exposure, which may affect real-time measurements.

### 2.2.3. Temperature measurement

By means of a long-wave FLIR A655sc IR camera, the back surface and through-thickness 2D temperature field was monitored during the flame exposure, as shown in Fig. 6. With regard to the measured surfaces at elevated temperatures, an emissivity of 0.9 is adopted (based on the sensitivity analysis performed in Ref. [30]). This assumption is made regardless the testing period and the area of the specimen under consideration (Fig. 6), without significant loss of measurement precision [30]. In the case of back surface temperature, the measurement of open-hole area was not considered, given that the saturation level of the IR camera had been reached. These thermography analyses allowed to quantify the in-plane and through-thickness temperature gradients for an exposure time ranging up to 900 s. Specifically, the real-time swelling (i.e., the macroscopic change in thickness) of the specimens during the fire exposure could also be obtained via the pixel-wise segmentation of through-thickness temperature field.

### 2.2.4. Fractographic and structural health analyses

Fractographic analysis was performed using optical microscopy (Keyence VHX-5000 numerical confocal microscope) and X-ray tomography (RX Solutions EasyTOM-150) to reconstruct 3D images of the damaged areas. The analyses were carried out on specimens subjected to thermal aggression as well as after mechanical loading in order to visualize, classify and quantify the thermally- and mechanically-induced damages.
<figure><img src="image_6.png" /><figcaption>Fig. 5. Full field measurement during combined tensile testing and kerosene flame exposure: (a) Description of DIC measurement range – (b) Axial displacement field measurement.</figcaption></figure>

<figure><img src="image_7.png" /><figcaption>Fig. 6. Full field temperature measurement during combined tensile loading and flame exposure: (a) Back surface (angle 45° in Fig. 4) – (b) Though-the-thickness (angle 0° in Fig. 4).</figcaption></figure>

<figure><img src="image_8.png" /><figcaption>Fig. 7. Temperature evolution along the central x-axis (defined in Fig. 6.a) and in-plane thermal degradation extension on the back surface of open-hole CG/ PEEK laminates.</figcaption></figure>
# 3. Experimental results and discussion

## 3.1. Thermal response under burn-through kerosene flame exposure

As shown in Fig. 7, the temperature field measured on the back surface (along the central x-axis defined in Fig. 6a) during fire exposure is first examined, providing valuable information about the thermal degradation state of the PEEK matrix with respect to its characteristic temperatures (Table 1). At the beginning of the exposure, a notable steeper increase in temperature near the open hole was observed, during which the pyrolysis gases have been ignited (at about 180 s). After a 300 s flame exposure, the average temperature within the exposed zones stabilizes around 375 °C, indicating in-plane thermal equilibrium near the open hole. As expected, the extent of thermal degradation (illustrated by the darker region in Fig. 7) on the back surfaces expands as flame exposure time increases, exhibiting a similar evolution trend of the range of temperatures exceeding the melting temperature (350 °C).

Regarding the through-thickness direction, Fig. 8 shows the average temperature within the fire exposure area as a function of fire exposure time. Regardless the ply position within the laminates, the temperature rapidly increases in the initial 180 s, then tends to stabilize at different temperature levels and increases slowly after 300 s, indicating a thermal quasi-equilibrium through the thickness. It is observed that all the plies in the central zone have reached the melting temperature (350 °C) after being exposed to the flame for 300 s (Fig. 9a). The number of plies through the thickness that have reached the onset of thermal decomposition (550 °C) gradually increases with increasing flame exposure time. After 900 s, the plies having reached the pyrolysis temperature represent about 2/3 of the thickness, indicating a severe degradation of the structural integrity.

With the pixel-wise segmentation of through-thickness temperature field, the real-time swelling of specimens can be measured, as shown in Fig. 9(a). Similar to the evolution of temperature, the swelling ratio tends to stabilize after 180 s. This suggests the formation of char, which is typically porous and acts as an insulating layer that slows down the heat transfer through the thickness of the material [10]. As for the in-plane and through-thickness heat conduction, the temperature gradients are determined at the positions of the red arrows (with a length of 4.5 mm) illustrated in Fig. 6(a) (left and right parts of central x-axis) and Fig. 6(b) (through-the-thickness z-axis) respectively. In the assessment of the through-thickness gradient, the swelling of specimens (i.e., the thickness evolution) has been considered.

It is clearly shown that the through-thickness temperature gradient has a major contribution (nearly tenfold greater than that of the in-plane temperature gradient) to the thermal conduction within the open-hole specimens. There is a steep increase in the first 100 s and then a constant increase of about 40 °C /mm. The slight decrease (approx. 90 s) is mainly due to the significant swelling of the laminate (Fig. 9a), which is associated with thermal expansion as well as laminate structural

relaxation (including deconsolidation of the matrix and release of stored internal stresses, i.e., "fiber springback"). Along the central x-axis, the temperature gradient on both sides of the specimen shows the almost identical temporal evolutions, confirming a symmetrical exposure to the flame during the tests.

## 3.2. In-situ tensile behavior under burn-through kerosene flame exposure

The experimental results shown in Fig. 10 exhibit a high degree of consistency within each group of specimens (300 s – 600 s – 900 s). As flame exposure time increases, the dispersion becomes more noticeable. Under a prolonged burn-through flame exposure, the tensile behavior undergoes a transition from an elastic-brittle mode (as exhibited in the as-received states of notched and unnotched specimens tested at room temperature [31]) to an elastic-plastic mode, along with the gradual failure of the 0/90° plies. In particular, it was also observed that the strain at failure of the fire-damaged open-hole specimens is clearly higher than that of the notched virgin specimens. And after a prolonged flame exposure of 900 s, the strain at failure of the open-hole specimen nearly approached that of the unnotched virgin specimen. This indicates the manifestation of the specimens' plasticity, along with the less pronounced open-hole effect resulting from material degradation under flame exposure.

As illustrated in the right part of Fig. 10, the macroscopic open-hole tensile failures of specimens in Group 2 were recorded using a CCD camera under thermo-mechanical stress after different flame exposure times. The ablation of the outer plies of G/PEEK and the pull-out of carbon fibers were observed at failure. A large amount of fume was released in the 300 s case, as the unburnt resin on the fracture surface came into contact with the flame. The failure mechanisms become more pronounced in the case of 900 s, as evidenced by the significant pull-out of ±45° plies following the failure of 0° plies and the extensive delamination. This may be explained by the ever-elevated average through-thickness temperature of 570 °C, which is reached within the laminate after a 900 s flame exposure.

To provide a clear insight of the trend of in-situ tensile behaviors changing with different exposure times, one curve representative in each group was selected in Fig. 11(a) with a background of light color, representing the range of all three curves. As the exposure time increases, the strain at failure exhibits a marked increase, with an average value of 1.3 % at an exposure time of 600 s, whose value is 111.1 % higher than that of the virgin state (notched specimen). Compared to the virgin state of notched specimen, the increasing fire exposure time causes the gradual degradation of tensile properties, as shown in Fig. 11(b). It is noted that the progressive failure of the 0° plies occurred earlier in the 300 s case, where the tensile load is transferred from one ply to another (Fig. 11a). These non-linear characteristics can be attributed to the progressive erosion of the PEEK matrix's load-transfer capabilities along its multiple phase transitions under fire, ranging from the melting temperature (from the rubbery to the viscous state) up to the onset of resin pyrolysis, and ultimately to the formation of char.

Depending on the pyrolysis degree of each ply, the load-bearing capacity of the different plies is subsequently reduced (as a function of the exposure time). With respect to the as-received notched specimen (not subjected to pre-heating and tested at room temperature), the considerable reductions in the open-hole tensile strength (–55 %) and the equivalent stiffness (–67 %) were observed after a 900 s flame exposure (Fig. 11b). With the introduction of an in-plane temperature gradient (due to open hole), contrary to the tensile test results in a preliminary study [17] conducted on post-fire CG/PEEK laminates (unnotched dog-bone specimens), the mechanical properties substantially decrease with a prolonged exposure time from 600 s to 900 s (even with an accelerated reduction in axial strength).

Nevertheless, as shown in Fig. 12 (obtained from a pre-test open-hole specimen without tensile loading), the residue of the PEEK matrix appears to have blocked the pathway of the flame after a 300 s flame

<figure><img src="image_9.png" /><figcaption>**Fig. 8.** Through-thickness temperature measurement as a function of flame exposure time.</figcaption></figure>
<figure><img src="image_10.png" /><figcaption>Fig. 9. Temperature profiles and gradients: (a) Through-thickness temperature evolution and swelling ratio of laminates – (b) Evolution of in-plane and through-thickness temperature gradients (as represented by red arrows in Fig. 2) and swelling ratio of laminates.</figcaption></figure>

<figure><img src="image_11.png" /><figcaption>Fig. 10. In-situ open-hole tensile behavior of CG/PEEK laminates and macroscopic observations of the thermo-mechanical failure (by CCD camera) under burn-through kerosene flame exposure.</figcaption></figure>

<figure><img src="image_12.png" /><figcaption>Fig. 11. In-situ open-hole tensile behavior of CG/PEEK laminates under burn-through kerosene flame exposure: (a) Macroscopic strain-stress responses of notched specimens – (b) Changes in open-hole tensile properties (axial strength and stiffness).</figcaption></figure>

exposure. This may provide an explanation for the relatively lower in-plane temperature gradient observed in Fig. 9(b), as the char layer acts as a thermal barrier that may slow down the planar heat transfer [19]. Additionally, higher axial thermal conductivity of the carbon fibers may promote heat spreading laterally along the laminate surface.

### 3.3. Failure mode under burn-through kerosene flame exposure

Fig. 13 shows the post-mortem through-thickness tomographic observations of the open-hole tensile failure modes of the CG/PEEK

laminates after different exposure times. As shown in Fig. 13, the red dotted lines represent the gradual extension of the char zone through the thickness, at the edge of the fire exposure area (Fig. 13a) and nearby the open hole (Fig. 13b), respectively, as a function of flame exposure time. The second region underlying the char exhibited an increasing proportion of delamination cracks between the plies as well as the fiber-matrix debonding within the plies up to 900 s, as a result of the conjunction of pressure rise and melting of the PEEK matrix.

From Fig. 13(b), the severe delamination of the outer G/PEEK ply on the back side near the open hole is observed. As expected, the plies
<figure><img src="image_13.png" /><figcaption>Fig. 12. Microscopic observation of the open hole on the exposed surface of post-fire CG/PEEK laminate after 300 s exposure to kerosene flame: 3D reconstruction in the right part, zoom view in the left part (the red dotted lines represent the edge of the initial open hole).</figcaption></figure>

<figure><img src="image_14.png" /><figcaption>Fig. 13. Post-mortem through-thickness tomographic observations of CG/PEEK laminates after open-hole tensile testing under burn-through kerosene flame exposure: (a) Cross-sections YZ at the edge of the fire exposure area (the yellow dotted lines in the right figure respectively) – (b) Cross-sections XZ along the central x-axis (the grey arrows indicates the loading direction and the yellow dotted line in the right part of Fig. 12).</figcaption></figure>

exposed to the kerosene flame are the first to fail. The tensile responses, depending on the thermal pyrolysis of the different plies, showed a gradual failure of the (0/90°) fibers. The breakage of 0° fibers is associated with a steep increase in axial strain until the tensile loading is transferred to the undamaged plies. This is followed by a discontinuity of the stresses between the plies, which causes extensive delamination within the laminates, as can be clearly seen in the cross-sections of the failed specimens (Fig. 13b).

Fig. 14 shows the post-mortem microscopic observations of the open hole from the exposed surface and back views of the damaged CG/PEEK laminates under tensile loading in fire. As a consequence of one-sided burn-through flame exposure, the through-thickness and in-plane temperature gradients lead to heterogeneous thermal degradation (porosity and carbonization). On the exposed surface (Fig. 14b), the pyrolysis of the PEEK matrix resulted in the formation of carbonaceous char. Melted PEEK matrix and porosity resulting from pyrolysis gas growth are observed on the back surface (Fig. 14a), indicating that the laminate degrades gradually with increasing exposure time. After the PEEK matrix melting, the bonding of the fiber bundles is no longer ensured by the matrix, and that causes a significant transverse contraction of the fibers as well as a more pronounced elongation (300 s case in Fig. 14a).

In terms of failure mode, as temperature increases through the thickness (Fig. 8 and Fig. 14a), it is observed that the role played by the different plies may vary from one to another. The structural integrity of the plies near the exposed surface is relatively maintained by the outer protective layer of char, with a clear bridging of 0° carbon fibers (900 s case in Fig. 14b). One must remind that the tensile load is mainly borne by the layers with (0/90°) fibers, accounting for 73 % of the total tensile load according to the Classical Laminate Theory (CLT) for unnotched laminates [8]. The elongation of the 0° fiber bundles near the back surface ultimately results in a marked misalignment of the 90° fiber

bundles (600 s case in Fig. 14a). Under force-controlled loading condition, the generalized necking leads to significant out-of-plane displacements, causing an extensive delamination (900 s case in Fig. 14a). Following the failure of 0° plies, the pull-out of fiber bundles initially oriented at ±45° was consequently observed (Fig. 14b). The complete ablation (due to pyrolysis of the matrix) of the outer G/PEEK ply after tensile failure is observed in the 900 s case (Fig. 14b). Furthermore, as PEEK matrix underwent rubbery and viscous transitions at high temperature, a reduced amount of notch tip damage was observed (300 s case in Fig. 14a). Similarly to preliminary studies [5,6] conducted on the notched C/PPS laminates, these mechanisms are found to be effective in releasing overstress, thereby reducing open-hole sensitivity.

# 4. Analytical modeling results and discussion

## 4.1. Effect of thermal decomposition on in-situ mechanical properties

As indicated in Section 1.2, the knowledge of the mechanical properties at different isothermal temperatures is the fundamental input parameter for modeling purposes. The isothermal in-situ tensile tests have been performed in the previous studies [8,17,31] on the CG/PEEK laminates. To extend these input thermo-mechanical laws P(T) up to the pyrolysis of PEEK, the influence of pyrolysis on the in-situ mechanical properties is to be investigated in a first step. A set of complementary tests (based on the same experimental protocol as in Section 2.2) was performed to better account for the pyrolysis of the PEEK matrix, which requires the in-situ tensile properties of the char, as shown in Fig. 15. The unnotched and notched specimens were first exposed to a one-sided kerosene flame for 300 s, followed by a cooling sequence to room temperature. Then, another side of the partially thermally-degraded specimens was exposed to a 900 s of flame exposure for a second
<figure><img src="image_15.png" /><figcaption>**Fig. 14.** Post-mortem microscopic observations of the in-situ open-hole tensile failure modes of CG/PEEK laminates under burn-through kerosene flame exposure condition: (a) Back surface – (b) Exposed surface.</figcaption></figure>

sequence, after which the quasi-charred (completely-degraded) specimens underwent force-controlled tensile loading until failure. The selection of the 300 s for the first exposure sequence is driven by the practical consideration of ensuring that the specimen's integrity is not overly damaged for the subsequent sequence of 900 s flame exposure.

According to the temperature profile shown in Fig. 8 and the TGA test results from a previous study [19], these properties can be approximated as the corresponding in-situ tensile behavior at 625 °C, where the pyrolysis of the PEEK matrix is almost complete. In the exposed section of the specimens, an average mass loss of 37.2 % was achieved, the value of which is close to the matrix content ratio of 42 wt. % (from the technical data sheet of the material manufacturer). Hence,

these thermally-decomposed laminates can also be regarded as the charred material (PEEK matrix residue with dry carbon fibers). Fig. 15 shows the tensile response of the unnotched and open-hole quasi-charred specimens. Compared to the virgin states, their tensile properties have been dramatically reduced (shown in Fig. 16), with 18.9 % and 8.2 % of axial strength remaining, and 7.8 % and 7.4 % of (equivalent) stiffness remaining, respectively. The results suggest that the charred laminates still retain a small fraction of mechanical properties, and their (equivalent) stiffness exhibit a much lower sensitivity to the open hole, as the material contains little polymer matrix content. This highlights the importance of exposure time (longer than 900 s) it takes to completely degrade the material, and the possibility of having a
<figure><img src="image_16.png" /><figcaption>Fig. 15. In-situ open-hole and unnotched tensile behavior of quasi-charred laminates and macroscopic observations of the thermo-mechanical damage under burn-through kerosene flame exposure.</figcaption></figure>

<figure><img src="image_17.png" /><figcaption>Fig. 16. Analytical model inputs: (a) Thermo-mechanical laws (unnotched specimens) as a function of isothermal heating temperature – (b) Through-thickness temperature profiles at 300–600–900 s.</figcaption></figure>

laminate exposed to fire for a certain amount without losing most of its structural integrity. During the gradual tensile failure, extensive delamination, ablation of the outer G/PEEK plies and pull-out of the carbon fiber bundles were also observed, indicating a low-level of structural integrity.

## 4.2. Modeling input: mechanical properties as a function of temperature

As mentioned earlier, the core of the analytical modeling is based on the thermo-mechanical laws of the in-situ mechanical properties as a function of isothermal testing temperatures, over a wide range of multiple phase transitions. In a second step, the modeling approach presented in Section 1.2 is adapted to the CG/PEEK laminates in order to represent the in-situ tensile behavior in fire. For this purpose, the curves fitting the in-situ mechanical property-temperature relationships proposed by Gibson et al. [21,27] (Eq. (3)) and Mahieux et al. [25,28] (Eq. (4)) are considered. As shown in Fig. 16(a), compared to the former (take example of axial strength), the latter assumes that the mechanical properties are negligible after the final phase transition temperature (pyrolysis of the matrix into char). As discussed in Section 4.1, the residual char resulting from the pyrolysis of the PEEK matrix retains non-negligible mechanical properties. The first curve-fitting model [21] is thus employed in this study, adapted with an additional term considering the decomposition phase transition temperature ($T_d'$). The properties $E(T)$ and $\sigma_f(T)$ can be calculated from an extended version of Eq. (3):

$$ 
\begin{aligned}
E(T) = & E_0 - \frac{E_0 - E_g}{2} [1 + \tanh(k_g(T - T_g'))] \\
& - \frac{E_g - E_m}{2} [1 + \tanh(k_m(T - T_m'))] \\
& - \frac{E_m - E_d}{2} [1 + \tanh(k_d(T - T_d'))]
\end{aligned}
\qquad (7) $$

$$ 
\begin{aligned}
\sigma_f(T) = & \sigma_{f,0} - \frac{\sigma_{f,0} - \sigma_{f,g}}{2} [1 + \tanh(k_g(T - T_g'))] \\
& - \frac{\sigma_{f,g} - \sigma_{f,m}}{2} [1 + \tanh(k_m(T - T_m'))] \\
& - \frac{\sigma_{f,m} - \sigma_{f,d}}{2} [1 + \tanh(k_d(T - T_d'))]
\end{aligned}
\qquad (8) $$

Where $T$ is the average temperature (exposure area) of each position through the thickness, the phase transition temperatures (both 20 °C lower than those measured by thermal techniques in this study) are the glass transition temperature ($T_g' = 123$ °C), the crystalline transition temperature ($T_m' = 330$ °C) and the decomposition temperature ($T_d' = 560$ °C) respectively, $E_i$ and $\sigma_{f,i}$ is the instantaneous property (stiffness and strength) after each phase transition, $k_i$ is the constant characterizing the range of phase transition, $T_0$ refers to the room temperature

<table><caption>**Table 2**
Analytical modeling input parameters of CG/PEEK laminates.</caption><thead><tr><th>Phase transitions</th><th></th><th>σ<sub>f,i</sub>(T)/MPa</th><th>E<sub>i</sub>(T)/GPa</th><th>k<sub>i</sub>/K<sup>-1</sup></th></tr></thead><tbody><tr><td>T<sub>0</sub></td><td>RT = 20 °C</td><td>791</td><td>52.5</td><td></td></tr><tr><td>T<sub>g</sub></td><td>T'<sub>g</sub> = 123 °C</td><td>639</td><td>45.3</td><td>0.02</td></tr><tr><td>T<sub>m</sub></td><td>T'<sub>m</sub> = 330 °C</td><td>278</td><td>27</td><td>0.03</td></tr><tr><td>T<sub>d</sub></td><td>T'<sub>d</sub> = 560 °C</td><td>57</td><td>18</td><td>0.025</td></tr></tbody></table>
(RT), the subscripts g, m and d refer to each phase transition, as shown in Table 2.

From Table 2, the input thermo-mechanical laws of axial strength and stiffness as a function of temperature (based on Eq. (7)) up to the pyrolysis of resin have been identified, as shown in Fig. 16(a). It is worth noting that all of these properties were obtained with the unnotched specimens, and the notch sensitivity α is introduced to account for the effect of the open hole. Notch sensitivity is defined as the ratio of the notched to the unnotched ultimate strength of the laminate, which depends on the ability of the laminate to accommodate overstress near the hole [3,6]. In this study, the value of α is set to 0.9 [31].

As illustrated in Fig. 17(b), the through-thickness average stress distribution of QI specimens under one-sided flame exposure is ideally depicted as a function of temperature at different positions along the z-axis. It is clear that, in reality, stress is not constant throughout the thickness of the laminate, but rather varies depending on the ply direction. This dependence can be readily evaluated using classical laminate theory. Nevertheless, the objective of the proposed analytical model is to consider the temperature-induced effects on the average stress of a composite laminate, assuming it to be a homogeneous material. In this regard, contrary to the isothermal cases of a homogeneous laminate (Fig. 17a), the fracture is characterized by a gradual progression from the exposed surface (the highest temperature) to the back surface (the lowest temperature).

As depicted in Fig. 2(b), via the predicted values of E(T) and σf(T) at a particular temperature level, the stress-strain curves of each position σ(T, ε) can be modeled as follows:

$$ \sigma(T, \varepsilon) = \sigma_{max} \left[ 1 - \exp\left(\frac{-E\varepsilon}{\sigma_{max}}\right) \right] \quad \text{and} \quad \sigma_{max}(T) = 1.25\sigma_{f}(T) \quad (9) $$

Where the asymptotic yield stress σmax(T) is set at values that exceed the value of σf(T) by a margin of 25 % in this study (according to the isothermal tests [8]). The failure of each position through the thickness occurs at the point of its ultimate tensile strength, α·σf(T), with consideration for the effect of open hole. At a given strain value ε, the average laminate stress σ̄ after different exposure times (with the different temperature profiles, Fig. 16b) is determined by integrating the stress over the specimen's thickness as follows:

$$ \bar{\sigma} = \frac{1}{h} \int_{0}^{h} \sigma(T, \varepsilon) dz \quad (10) $$

Where h is the thickness of specimen. It is worth noting that the variance in the y-axis direction was not considered here, since the in-plane temperature gradient is much less important than the through-thickness one. Finally, the calculation of the average stress for a range of strains enables the formulation of a “laminate” stress-strain curve after exposure times of 300–600–900 s respectively, as shown in Fig. 18 (a).

<figure><img src="image_18.png" /><figcaption>Fig. 17. Average stress distribution under different heating conditions: (a) Isothermal heating – (b) One-sided flame exposure.</figcaption></figure>

### 4.3. Modeling results of in-situ tensile properties in fire

Fig. 18 shows the prediction results of the in-situ open-hole tensile behavior using the proposed modeling method and the input data presented in Section 4.2. As shown in Fig. 18(a), the modeled curves exhibit the gradual failure of laminate resulting from a gradual decrease in temperature-dependent mechanical properties through the thickness. Compared to the experimental results in light colors, the predicted strain-stress curves are found to well represent the in-situ thermo-mechanical nonlinear behavior as a function of exposure time up to 900 s. The modeled strain at failure displays a decreasing tendency with prolonged flame exposure time, given the increasing average temperature within the laminate. This outcome is primarily due to the adoption of a constant ratio between σmax and σf, whose value undergoes a decreasing trend at higher temperatures (indicating enhanced non-linearity) [21]. Furthermore, given that the proposed model is formulated in terms of average stress (under the assumption of laminate homogeneity), the bending induced by the failure of fire-exposed plies is not considered. The laminate is regarded as being loaded in membrane, a premise that is invalid once the mirror symmetry of the laminate is no longer satisfied.

Fig. 18(b) presents the outputs of the model regarding the changes in open-hole tensile properties as a function of flame exposure time. The modeling results show a good agreement with the experimental ones, most of the predicted values are located in the experimental data domains (in light colors). The notch sensitivity α was found to be acceptable to account for the open-hole effect. Compared to the notch sensitivity at RT (α = 0.55) and at 150 °C (α = 0.61) [31], the overstress accommodation effect near the open hole may be reduced under severe temperature conditions. The observed data scatter can be attributed to the identical input properties of all the plies, whose differences may be accentuated by prolonged fire exposure. In brief, this model provides a practical tool for the design of composite structural parts to be used under burn-through flame exposure.

## 5. Conclusions

To investigate the thermo-mechanical coupling of CG/PEEK laminates with an open hole under burn-through flame, a fire bench test has been specifically designed to provide a small-scale set-up reproducing critical conditions in service (fire scenario in aircraft engine). In accordance with fire certification standards, the open-hole specimens were exposed to a kerosene flame (1100 °C and heat flux of 116 kW/m²) under tensile loading, and temperature profiles of the specimen were measured. Property modeling through multiple phase transition temperatures of PEEK matrix has been successfully implemented to represent and predict the tensile behavior of open-hole specimens under fire conditions.

The following conclusions can be drawn: (i) the burn-through flame exposure induces through-thickness and in-plane temperature gradients, with the former playing a more important role in heat conduction (nearly 10 times higher than the latter); (ii) after the melting of the PEEK
<figure><img src="image_19.png" /><figcaption>**Fig. 18.** Modeled in-situ open-hole tensile behavior of CG/PEEK laminates under burn-through kerosene flame exposure: (a) Modeled macroscopic stress-strain responses – (b) Changes in open-hole tensile properties as a function of flame exposure time.</figcaption></figure>

matrix, the carbon fiber bundles are no longer bonded to the matrix, leading to significant transverse fiber contraction and significant elongation near the open hole; (iii) the tensile behavior shifts from an elastic-brittle mode (virgin state) to an elastic-plastic mode with gradual failure, the ablation of the outer plies of G/PEEK and the pull-out of carbon fibers were observed; (iv) after a 900 s exposure, dramatic decreases in the axial strength (–55 %) and the equivalent stiffness (–67 %) were measured; (v) based on the in-situ isothermal heating and quasi-charred specimen mechanical tests, the input thermo-mechanical laws representing the mechanical properties as a function of temperature up to the PEEK pyrolysis were obtained; (vi) the predicted strain-stress curves show a good agreement with the experimental results, proving the ability of the model to accurately simulate the thermo-mechanical behavior of composite structures under burn-through flame exposure.

To better understand the thermo-mechanical coupling, the model should be improved by combining a thermal model accounting for the through-thickness temperature profiles and the pyrolysis degree. Furthermore, it would be relevant to better assess the effect of the open hole, concerning the burn-through kerosene flame exposure conditions.

### CRediT authorship contribution statement

**Lanhui Lin:** Writing – original draft, Visualization, Software, Methodology, Investigation. **Benoît Vieille:** Writing – review & editing, Validation, Supervision, Resources, Project administration, Methodology, Data curation, Conceptualization. **Christophe Bouvet:** Writing – review & editing, Validation, Methodology, Conceptualization. **Tanguy Davin:** Writing – review & editing, Validation, Supervision, Methodology.

### Declaration of competing interest

On behalf of my co-authors, I confirm that there is no conflict of interest regarding the publication in Composites Part A: "In-situ Open-hole Tensile Testing and Modeling of Hybrid PEEK Thermoplastic Laminates under Burn-through Kerosene Flame Exposure".

### Acknowledgments

The financial support from China Scholarship Council (CSC) is greatly appreciated (CSC No. 202208120011).

### Data availability

The data that support the findings of this study are available from the corresponding author upon reasonable request.

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