# Evaluation of a new adhesively bonded sample configuration adapted for ELS test and creep frame setups

Arij Fawaz, Emilie Lepretre, Xavier Chapeleau, Sylvain Chataigner

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Arij Fawaz, Emilie Lepretre, Xavier Chapeleau, Sylvain Chataigner. Evaluation of a new adhesively bonded sample configuration adapted for ELS test and creep frame setups. CICE 2025 - 12th International Conference on FRP Composites in Civil Engineering, Jul 2025, Lisbon, Portugal. hal-05398814

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# EVALUATION OF A NEW ADHESIVELY BONDED SAMPLE CONFIGURATION ADAPTED FOR ELS TEST AND CREEP FRAME SETUPS

Arij FAWAZ<fnref n="1" />,<fnref n="2" />, Emilie LEPRETRE<fnref n="1" />, Xavier CHAPELEAU<fnref n="2" /> and Sylvain CHATAIGNER<fnref n="1" />

## Abstract

Despite the increasing use of structural adhesives for the repair or reinforcement of metallic structures (e.g., using adhesively bonded FRP systems), the durability of bonded joints under continuous loading remains a significant challenge. In fact, it affects their long-term performance and reduces their mechanical properties. Mechanical characterization tests are essential for evaluating the behavior of bonded joints before and after creep aging. However, the sample geometry required for mechanical characterization differs, in some cases, from that used in the creep testing setup. This study aims to propose and validate a new sample configuration reffered to as configuration 2, as an alternative to the conventional symmetrical sample geometry (noted as configuration 1), that integrates both the creep test frame and the mechanical characterization test setup. The study focuses specifically on fracture mechanics test in shear mode (mode II), which has been identified as the dominant failure mode in the literature. The End-Loaded Split (ELS) test, which is commonly used for mode II testing of bonded assemblies, was adapted in this work. The creep frame used in this study consists of a four-point bending setup allowing a constant stress state in the adhesive bondline. The mode II mechanical properties for both configurations were compared. First, mode II toughness was calculated at the initial state according to the ELS test standard. Subsequently, Digital Image Correlation (DIC) was used to assess the J-integral in mode II, and optical fiber sensors were employed to monitor crack growth during propagation.

## Keywords

Adhesively bonded assemblies; Mode II fracture mechanics test; ELS (End-Loaded Split) test; creep; J-integral; DIC (Digital Image Correlation).

## 1. Introduction

Adhesively bonded structures are usually exposed to harsh environmental conditions throughout their service life, which represents a major drawback for their large-scale implementation. Bonded joints are frequently exposed to creep aging conditions. Since adhesives are known for their viscoelastic behavior, they will present gradual deformation over time under constant stress. This deformation can affect the

<footnote n="1">¹MAST/SMC, Université Gustave Eiffel, Bouguenais 44344, France
email: arij.fawaz@univ-eiffel.fr</footnote>

<footnote>email: emilie.lepretre@univ-eiffel.fr (for author 2)</footnote>

<footnote>email: sylvain.chataigner@univ-eiffel.fr (for author 4)</footnote>

<footnote n="2">²COSYS/SII-I4S, Université Gustave Eiffel, Inria, Bouguenais 44344, France
email: xavier.chapeleau@univ-eiffel.fr (for author 3)</footnote>
mechanical properties of the bonded assembly and can lead, ultimately, to total failure of the joint. Therefore, investigating their long-term functionality and performance is crucial for extending their use in various civil engineering applications.

To evaluate the impact of creep on the durability of bonded joints, mechanical properties before and after exposure to creep aging have to be examined. Several methods are available in the literature, such as method based on the use of the Linear Fracture Elastic Mechanics (LFEM). This method has been proven robust and has been extensively used by researchers. For instance, Q. Sourisseau [1] validated its reliability through mechanical characterization tests on metal-to-metal bonded joints in mode I (opening mode), mode II (shear mode), and mixed-mode conditions.

This study focuses on investigating the mechanical properties in mode II, as mode II is the most common mode of solicitation and plays a more dominant role in the behavior of bonded joints compared to mode I. The mode II mechanical test employed in this research is the End-Loaded Split (ELS) test, which offers greater crack propagation stability compared to other mode II tests [2]. The ELS test involves clamping one edge of the sample and applying a vertical load or displacement at the opposite edge.

Investigating the creep behavior of a bonded assembly using a mechanical characterization test requires a sample geometry compatible with both setups. Previous studies in the literature have primarily focused on either the mechanical test or the creep frame separately, without a combined configuration. Therefore, the aim of this study is to propose a new sample configuration that integrates these two setups. First, this paper introduces the newly modified sample geometry. Then, the conducted tests are presented. All samples were equipped with optical fiber to monitor crack propagation during the ELS test. Additionally, Digital Image Correlation (DIC) was employed to enhance the understanding of the mechanical behavior of the tested samples.

## 2. Experimental investigations

### 2.1. Bonded joints geometries

In this study, the specimen has to be submitted to a constant load and then tested mechanically using the ELS test. Consequently, the specimen's geometry and setup must be tailored to accommodate both the creep frame and the ELS test configurations.

For the ELS test, a symmetric sample is required to ensure that the neutral axis is located in the middle of the adhesive layer (see configuration 1 specimen in Figure 1). This adhesive layer is positioned between two adherends of the same material and equal thickness. Such a configuration is essential for promoting predominantly shear deformation within the adhesive layer.

Creep testing can be performed using either a tensile or a bending frame. In our case, bending was chosen due to its simpler experimental implementation. More specifically, A bending frame can employ either a three-point or four-point loading system [3]. In a three-point loading system, the bending moment varies linearly between the supports and the center of the sample (i.e., the loading point). On the other hand, in a four-point loading system, the bending moment remains constant between the loading points inducing a constant stress state along the adhesive bondline. Thus, a four-point bending setup was selected to ensure a constant bending moment between the loading points and thus a constant stress state along the adhesive bondline [4]. In order to derive a non-zero stress value along x-axis, $\sigma_{xx}$, within the adhesive layer, thanks to the four-point bendig setup, it was decided to use a non-symmetric design (i.e., different adherends' thicknesses with adherends made from the same material).

Then, the non-symmetric sample has to become symmetric in order to realize the ELS test. It corresponds to the configuration 2 specimen shown in Figure 1. This new geometry will be compared to the typical symmetric bonded assembly, named configuration 1.
<figure><img src="image_2.png" /><figcaption>**Figure 1.** Configuration 1 and configuration 2 specimens</figcaption></figure>

In configuration 2, the non-symmetric sample is thus first aged under creep conditions. Then, before conducting the ELS test, the sample is converted into a symmetric one by bonding a third plate onto the thinner lower adherend. The thickness of this additional plate is selected to ensure equal stiffness between the upper and the lower adherend (corresponding to the sum of lower adherend 1, adhesive 2 and lower adherend 2).

## 2.2. Materials properties and joint manufacturing

In this study, two commercial adhesives with different chemical compositions were selected. The first, referred to as adhesive 1, is a bi-component epoxy adhesive in which the resin is thoroughly mixed with the hardener. This epoxy-based adhesive has a Young's modulus of 3100 MPa and a glass transition temperature of 91°C, according to the corresponding technical data sheet. The second adhesive is methacrylate-based (adhesive 2), with a glass transition temperature of 85°C and a Young's modulus ranging between 1200 and 1700 MPa, as specified in its technical data sheets.

The primary adhesive tested in this study is the epoxy adhesive. The methacrylate adhesive is used only to achieve symmetry in the sample before conducting the ELS test, as this test requires a symmetrical sample. The crack always propagates in the epoxy adhesive layer, and only the properties of the epoxy adhesive are evaluated in this study.

To prevent plastic deformation during testing, all steel plates used for both configurations are made of high-yield steel. Plates with thicknesses of 4 mm and 5 mm have an elastic limit of 700 MPa (STRENX 700), while plates with a thickness of 0.63 mm have an elastic limit of 925 MPa.

Before applying the adhesive, the adherends were sandblasted and cleaned with acetone. The adhesives were then cured according to their respective technical data sheets: Adhesive 1 was cured at 80°C for 16 hours, while adhesive 2 was cured at 25°C for 24 hours.

## 2.3. ELS test and instrumentation

The ELS test is conducted using a test rig that complies with the specifications defined in the ISO 15114:2014 standard [5]. This test rig is placed in an electromechanical traction machine to apply the load vertically. The load application point remains fixed. The load is applied at the pre-cracked edge of the specimen, where an artificial crack is created using a Teflon insert with a length of $a_0$=75 mm. At the other edge, the specimen is clamped through a fixture that consists of a block sliding laterally on bearings. The initial free length, which is the distance from the clamping edge to the load application point, is set to $L$=100 mm.

To prevent the impact of the artificial crack front on the results, a real mode II pre-crack was created via the ELS test. In this study, only the tests with a real pre-crack are presented, and those with an artificial pre-crack are not included. The real pre-crack length is denoted as $a_p$. A ratio of $a_p/L$ higher than 55% was selected to enhance the stability of crack propagation.
The calibration of the clamping fixture was performed prior to conducting the ELS test, in accordance with the procedure outlined in the ISO 15114:2014 standard [5]. This calibration accounts for the clamping effects and is primarily carried out to determine the clamp correction value $\Delta_{clamp}$ and the flexural modulus $E_I$ of the specimen.

Each specimen is equipped with a single-mode polyimide-coated optical fiber, a DIC system, and an inclinometer (Figure 2). The optical fiber was bonded to the center of the outer surface of the upper adherend. A Luna ODiSI-B Optical Distributed Sensor Interrogator was used. Along the optical fiber, multiple strain gauges are positioned at equal distances, enabling continuous strain measurement over time and along the entire length of the optical fiber. Through the obtained strain profile, the crack length can be determined during crack propagation. Furthermore, a DIC system was used to measure the rotation angles and the displacement at the S-S' section. This section is located sufficiently far from the clamping end and the initial crack tip, positioned between these two points (refer to [6] for more details). A black-on-white speckle pattern was applied using an airbrush along the edge of the specimen, positioned in front of two cameras. The cameras used were Basler acA2440-75um models, each featuring 5 MP sensors, paired with Kowa LR1015WM52 50 mm lenses and set at a stereo angle of 15 degrees. The mesh size was automatically calculated by the commercial software Correli, and was determined to be 16 × 16 pixels (Figure 3). Finally, an inclinometer was placed at the load application point to measure the rotation angle at this location, a parameter required for the J-integral calculation.

The mode II fracture toughness $G_{IIC}$ was determined for ELS tests using multiple data reduction techniques to ensure a comprehensive evaluation of fracture behavior. Three LEFM-based methods are described in the ISO 15114:2014 standard [5]: the Experimental Compliance Method (ECM), the Simple Beam Theory (SBT) method, and the Corrected Beam Theory using Effective crack length (CBTE) method. The CBTE method relies on calculating an effective crack length, $a_e$ instead of using the actual crack length to account for the clamp effect. The formula for $a_e$ is provided in (Eq. 1).

According to the ELS test standard, the CBTE method provides the most reliable values compared to the other methods outlined in this standard [5]. Therefore, this study focuses exclusively on calculating mode II toughness using the CBTE method, with its formula provided in (Eq. 2).

$$a_e = \left[ \frac{1}{3} \left( 2b \frac{C}{N} h^3 E_1 - (L - \Delta_{clamp})^3 \right) \right]^{1/3} \quad (1)$$

$$G_{IIC} = \frac{9P^2 a_e^2 F}{4 b^2 h^3 E_1} \quad (2)$$

Where P is the applied load, b represents the specimen width, h is the height of a single adherend, C is the specimen compliance, and N and F are correction factors. For detailed information on calculating the correction factors N and F, refer to the ISO 15114:2014 standard [5].

In cases where LEFM assumptions are not valid, the non-linear fracture mechanics approach, such as the J-integral method, should be employed. Unlike the $G_{IIC}$ calculation, the J-integral method does not require crack length measurements (values that are significantly affected by the large FPZ under mode II loading conditions).
The J-integral approach was also used in this study to obtain the critical toughness in mode II. A J-integral formula for an ELS test, which includes corrections for large displacements, was proposed and validated by Perez Galmez et al. [6] (Eq. 3):

$$J_{ELS} = \frac{3[P \cos(\theta_p)]^2}{5G_{12}b^2h} + \frac{P \cos(\theta_p)}{b} \tan(\theta_p - \theta_g) - \frac{E_{11}h}{3} \left( \frac{3P(L - L_s)}{E_1bh^2} \right)^2 \quad (3)$$

In this equation, $G_{12}$ is the shear modulus, and $\theta_p$ is the rotation angle measured directly by the inclinometer positioned at the load-displacement point. Additionally, $\theta_g$ is the rotation angle derived from the displacement profile at an S-S' section, based on DIC data.

<figure><img src="image_3.png" /><figcaption>Figure 2. ELS test setup with the associated instrumentation</figcaption></figure>

<figure><img src="image_4.png" /><figcaption>Figure 3. 16x16 pixels meshing in the ROI</figcaption></figure>

# 3. Results and discussion

## 3.1. Load-displacement curves

Stable crack propagation was observed in all samples for both configurations. The corresponding load-displacement curves for various $a_p/L$ ratios are shown in Figure 4. Similar initial stiffness was obtained for both configurations, but a significant difference in the maximum load reached is evident. On average, the maximum loads for configuration 2 are 15% higher than those for configuration 1 samples. This may be due to the compressive forces applied to adhesive 1 following the sandblasting procedure before bonding the third steel plate
<figure><img src="image_5.png" /><figcaption>**Figure 4.** ELS load-displacement curves for configuration 1 and configuration 2 samples.</figcaption></figure>

## 3.2. G<sub>IIC</sub> and J-integral calculation

As previously described, a clamp calibration should be performed before conducting the ELS test. Following this calibration, the mean values of the flexural modulus and the calibration correction can be determined. Table 1 summarizes the obtained values for both configurations.

<table><caption>**Table 1.** Calibration correction and flexural modulus for configuration 1 and configuration 2 samples.</caption><thead><tr><th></th><th>Configuration 1</th><th>Configuration 2</th></tr></thead><tbody><tr><td>Δ<sub>clamp</sub> (mm)</td><td>39.41</td><td>38.09</td></tr><tr><td>E<sub>1</sub> (GPa)</td><td>238.85</td><td>245.17</td></tr></tbody></table>

The optical fiber provides the strain profile at each time step along the length of the specimen. The maximum strain is expected to correspond to the crack tip [7]. Consequently, the crack length can be determined during crack propagation at each time increment defined by the optical fiber parameters (i.e., acquisition frequency). This enables the calculation of the *ap/L* ratio for each performed test. Moreover, DIC is primarily used to measure the rotation angle and deformation at the S-S' section. These values are crucial for determining the J-integral.

The mean values of the energy release rate in mode II and the J-integral, calculated using (Eq. 2) and (Eq.3), respectively, are presented in Table 2. Six tests were conducted for configuration 1 samples, while seven tests were conducted for configuration 2 samples.

<table><caption>**Table 2.** Mean values of G<sub>IIC</sub> and J-integral for ELS samples of configurations 1 and 2.</caption><thead><tr><th></th><th>Configuration 1</th><th>Configuration 2</th></tr></thead><tbody><tr><td>G<sub>IIC</sub> CBTE (N/mm)</td><td><strong>0.94</strong><br>± 0.16</td><td><strong>1.11</strong><br>± 0.20</td></tr><tr><td>J-integral (N/mm)</td><td><strong>0.84</strong><br>± 0.15</td><td><strong>0.97</strong><br>± 0.14</td></tr></tbody></table>

The mean values of G<sub>IIC</sub> and the J-integral are higher for configuration 2 samples compared to configuration 1. This result aligns with the load-displacement curves obtained, as these values are proportional to the applied load *P*. Furthermore, G<sub>IIC</sub> based on the CBTE method exceeds the J-integral value by 12% and 15% for configuration 1 and 2 respectively. The variation between the J-integral and G<sub>IIC</sub> values primarily arises from the LEFM assumptions, which are used to calculate G<sub>IIC</sub> based on the CBTE method.

Indeed, the CBTE method, being LEFM-based, assumes that the Fracture Process Zone (FPZ) is negligible compared to the specimen's key dimensions, such as width, thickness, or crack length. However, for mode II loading tests, the FPZ is relatively large. As a result, the bonded assembly may
exhibit non-linear behavior, which can sometimes invalidate the LEFM assumptions for analyzing the fracture behavior of these bonded joints.

This discrepancy between $G_{IIc}$ and J-integral mean values is consistent with findings in the literature. For instance, Sun et al. [8] demonstrated that even when LEFM-based methods are coupled with the effective crack length approach (where the crack length is calculated from beam compliance), they remain less accurate for an ELS test. In this test, damage occurs not only within the FPZ but also away from it, such as near the clamping region. Consequently, the toughness formulas in the ELS test standard, which are LEFM-based, require revision to address these inaccuracies.

**4. Conclusions and perspectives**

In this study, a new sample configuration (denoted as configuration 2) is proposed, combining two test setups: the four-point bending creep test and the ELS test setups. This newly configured is compared to a typical symmetrical bonded assembly (referred to as configuration 1). Both configurations demonstrated the same initial stiffness in their load-displacement curves; however, configuration 2 samples reached higher loads at greater displacement values. All samples from both configurations were equipped with strain-distributed optical fiber sensors, enabling real-time crack length measurement based on the derived strain profiles. Additionally, a DIC system was used to capture deformation fields and the rotation angles at the S–S’ section.

Using data obtained from DIC, optical fiber measurements, and ELS test results, the $G_{IIc}$ value was calculated based on the CBTE method following the ELS test standard [5]. Furthermore, the J-integral for the ELS test samples was calculated using a slope-based equation proposed and validated by Perez Galmez et al. [6]. A difference of 12% and 15% between the $G_{IIc}$ values from the CBTE method and the J-integral was observed for configurations 1 and 2, respectively. This variation is primarily attributed to the LEFM assumptions of the CBTE method. The toughness formulas in the ELS test standard require revision to account for the non-linear behavior of the bonded assembly. In this context, slope-based equations, such as those for the J-integral, are recommended.

Given the observed differences between the two configurations, only one configuration will be selected for future investigations. Specifically, configuration 2 samples will be subjected to creep, humid and coupled aging conditions, then tested at various time intervals to assess the variation in mechanical properties of adhesively bonded joints under these conditions. Moreover, cohesive zone models can be derived for the different aging coniditons from the J-integral.

**Acknowledgments**

The authors would like to acknowledge the technical teams of the MAST/SMC laboratory and COSYS/SII laboratory at Université Gustave Eiffel, as well as the 'Région Pays de La Loire' for co-funding this study.

**References**

[1] Sourisseau, Q. (2022). *Development of a robust methodology for the design assessment of bonded reinforcements on steel structures* « Use of cohesive zone model and distributed optical fiber for characterization » [PhD thesis]. Nantes Université.

[2] Pérez-Galmés, M., Renart, J., Sarrado, C., Brunner, A. J., & Rodríguez-Bellido, A. (2018). Towards a consensus on mode II adhesive fracture testing: Experimental study. *Theoretical and Applied Fracture Mechanics*, 98, 210-219. https://doi.org/10.1016/j.tafmec.2018.09.014

[3] Hollenberg, G. W., Terwilliger, G. R., & Gordon, R. S. (1970). *Calculation of Stresses and Strains in Four-Point Bending Creep Tests*.
[4] Arao, Y., Koyanagi, J., Utsunomiya, S., Takeda, S., & Kawada, H. (2009). Analysis of time-dependent deformation of a CFRP mirror under hot and humid conditions. *Mechanics of Time-Dependent Materials, 13*(2), 183-197. https://doi.org/10.1007/s11043-009-9081-1

[5] ISO 15114:2014 standard : Fibre-reinforced plastic composites—Determination of the mode II fracture resistance for unidirectionally reinforced materials using the calibrated end-loaded split (C-ELS) test and an effective crack length approach (Version First edition). (2014).

[6] Pérez-Galmés, M., Renart, J., Sarrado, C., Rodríguez-Bellido, A., & Costa, J. (2016). A data reduction method based on the J -integral to obtain the interlaminar fracture toughness in a mode II end-loaded split (ELS) test. *Composites Part A: Applied Science and Manufacturing, 90*, 670-677. https://doi.org/10.1016/j.compositesa.2016.08.020

[7] Sourisseau, Q., Lepretre, E., Chataigner, S., Chapeleau, X., Mouton, L., & Paboeuf, S. (2022). Use of high spatial resolution distributed optical fiber to monitor the crack propagation of an adhesively bonded joint during ENF and DCB tests. *International Journal of Adhesion and Adhesives, 115*, 103124. https://doi.org/10.1016/j.ijadhadh.2022.103124

[8] Sun, F., Lißner, M., Petrinic, N., & Blackman, B. R. K. (2024). Universal slope-based J-integral methods for characterization of the mode I, mode II and mixed mode I/II fracture behaviour of adhesively bonded interfaces. *Composites Science and Technology, 252*, 110611. https://doi.org/10.1016/j.compscitech.2024.110611