concrete

Construction and Building Materials 227 (2019) 116385

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Construction and Building Materials journal homepage: www.elsevier.com/locate/conbuildmat

Influence of curing conditions on mechanical behaviour of glued joints of carbon fibre-reinforced polymer composite/concrete A.T. Lee a, M. Michel a, E. Ferrier a, Brahim Benmokrane b a b

LMC2, University Claude Bernard LYON 1, 82 bd Niels Bohr, 69622, Villeurbanne, France University of Sherbrooke 2500 Boulevard de l’Université, University of Sherbrooke, Sherbrooke, Québec J1K 2R2, Canada

h i g h l i g h t s
Carbon fiber reinforced polymer.
Tensile strength.
Glass transition temperature.
Curing conditions.
Bond on concrete.

a r t i c l e

i n f o

Article history: Received 17 November 2018 Received in revised form 11 July 2019 Accepted 15 July 2019

a b s t r a c t Polymeric adhesives are widely used for the bonding of carbon fibre-reinforced polymers (FRPs) on a concrete structure. Structural reinforcement using FRPs is a very efficient strategy owing to the good mechanical characteristics of FRPs and the implementation methods, which are applied onsite and make it possible to reduce the work time. Efficiency is limited by FRP debonding and adhesive has a key issue to play in the strengthening. Working onsite assumes that the conditions for implementation are highly dependent on the climate. In this context, it should be noted that the resins used are cold-curing epoxy and that their mechanical and physical properties are temperature- and humidity-dependent. In this study, the impact of the application temperature on the behaviour of composite reinforcement is investigated. The objective is to analyse the mechanical and physical behaviour of epoxy polymers used for reinforcement by measuring the mechanical properties of composite adhesion on a concrete support via double-shear and pull-out tests, as well as the glass transition temperatures Tg. The Tg measurements are performed using differential scanning calorimetry and thermomechanical analysis. Double-shear tests are performed to establish the local composite-to-concrete shear laws for adhesive joints according to the nature of the polymers and the curing conditions. The test results allow for determination of the bonding length (Le) and its dependence on the curing conditions. The physical measurements of Tg support the study by elucidating the influence of the curing conditions on these properties. Ó 2019 Elsevier Ltd. All rights reserved.

1. Introduction External bonded reinforced composite materials are widely used in civil engineering for structural reinforcement and building repair. Composites generally exhibit better mechanical properties than steel, in addition to advantageous such as light weight, durability, non-corrosive characteristics, and easy transportation to building sites. Carbon fibres are commercially applied in building reinforcement [1] owing to their high stiffness (Young’s modulus of 200–750 GPa) and strength (failure stress of 1.5–6 GPa). Several kinds of fibres have been used for different requirements according

E-mail address: [email protected] (E. Ferrier) https://doi.org/10.1016/j.conbuildmat.2019.07.111 0950-0618/Ó 2019 Elsevier Ltd. All rights reserved.

to their properties such as glass fibres (failure deformation up to 5%) and natural fibres (low environmental impact for sustainable development). The mechanical properties of composites depend on not only the fibre characteristics but also the matrix characteristics. The physical and mechanical properties of thermoset resins such as epoxy change according to the curing conditions. Benedetti et al. [2] investigated the evolution of the elastic modulus (E-modulus) and pullout force of epoxy resin for Near Surface Mounted–carbon fibrereinforced polymer (CFRP) systems at different curing temperatures (20, 30, and 40 °C). To accelerate the curing process, the composites can be exposed to an elevated temperature in an autoclave. Michels et al. [3] compared the glass transition temperature (Tg)—measured

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via dynamic mechanical analysis (DMA)—of resin specimens depending on curing conditions. They divided the specimens into two groups as those cured at room temperature and at an elevated temperature. The Tg consistently increased with the curing temperature owing to the crosslinking. The Tg depends on external factors such as product type, maximum curing temperature, specimen age, and heating rate. Although many studies on the influence of a high curing temperature on the evolution of Tg have been performed, few [4,5] have investigated the influence of low temperatures on Tg and consequently on the mechanical properties of epoxy. These studies considered the properties of the matrix and not the composite FRP as a reinforcement for a concrete structure. Many researchers focused on the physical aging of polymers and its influence on the glass transition temperature. The authors deal with the evolution of the mechanical and physical properties under different curing conditions and establish the correlation between the stiffness (Young’s modulus E) and the resistance (ultimate tensile force Fult) of the epoxy. A kinetic chemical analysis was also performed to determine the degree of curing as a function of time, as well as the cure rate according to the degree of curing. Michels et al. [3] described an accelerated curing method. DMA is used to determine the value of the glass transition temperature, which has several definitions. Similar to other studies, the authors investigated the influence of the curing conditions on the physical and mechanical properties of the material. The definition of the service temperature was presented at the end of the article, and the authors introduced the determination of this temperature under different standards to ensure the mechanical properties of the polymer. Lapique et al. [6] examined the development of the mechanical properties of glue, which is used in construction, under different curing conditions. First, the authors studied the curing procedure and explored the rheology (especially the viscosity) of the glue under different temperatures. The degree of curing and the minimum duration for ensuring a satisfactory degree of curing were also discussed. Finally, they studied the effects of the aging of the glue in a humid environment. Matsui et al. [7] reported that the shear modulus Ga and the shear stress sb of an adhesive depend on the curing temperature, curing time, test temperature, and resin–hardener ratio. In construction, it is important to define the temperature at which the external bonded CFRP reinforcement should not be applied to avoid the sharp decrease in strength of the strengthened structure. This section presents different definitions of the service temperature Ts [3] according to different standards. For example, according to fib-Bulletin 14 published in 2001 [8], the Tg of the epoxy must be >45 °C, or the temperature in the shade (shade temperature) must not exceed (Tg-20) °C. According to the Italian standard CNR-DT 200 R1/2013 [9] and the American standard ACI 440 [9], Ts = (Tg-15) °C. Ferrier et al. [10] demonstrated that the Tg of the epoxy must be >55 °C; otherwise, Ts is limited to (Tg-15) °C. However, according to the respective recommendations, the onset temperature (ACI 440, [11]) and the inflection temperature (TR 55 [12]) must be considered when evaluating Tg. During crosslinking, each hydrogen atom from an amine group opens an epoxide ring. This polymerisation occurs over time, and the molar mass of the polymers increases, increasing the Tg of the epoxy [13]. In general, the Tg of epoxy–amine mixture is not a constant, because the crosslinking is not perfect. When the curing temperature exceeds the Tg, the crosslinking starts again, and the remaining amines in the epoxy continue to open the remaining rings in the epoxides. As a result, the Tg continues to increase for a period of time. Undoubtedly, crosslinking plays an important role in the mechanical performance of the material. It ensures the mechanical strength, rigidity, and adhesion between the outer surface of the concrete and the reinforcement.

Amorphous materials (polymers, glasses, metals) have common characteristics specific to their vitreous state. In particular, they undergo changes in their structural state when they are maintained at a temperature below Tg. These changes, which are commonly known as physical aging or structural relaxation, affect the properties of the material (mechanical, dielectric, and thermal behaviour). The first studies related to these phenomena [14,15] concerned the evolution of thermodynamic quantities (specific volume, enthalpy) as a function of time. When the vitreous material is out of equilibrium, these modes of configuration are frozen. However, over time, the properties of the system evolve, and the material passes through intermediate energetic states, approaching the metastable equilibrium situation. Thus, for a temperature below Tg, the enthalpy of the system changes over time to the state corresponding to the supercooled liquid. The structural units of the molecular chains move to lower energy configurations. This physical phenomenon is called structural relaxation or physical aging. It is nonlinear, as demonstrated by epoxy experiments [16]. Kovacs showed that the direct kinetics are slowed by structural crosslinking (decrease in enthalpy). This is well-explained qualitatively: when the system evolves towards states of lower energy, its molecular mobility decreases. On the other hand, the inverse kinetics are self-accelerated because when the system evolves towards states of higher energy, the molecular mobility increases. This cannot be described by a simple exponential function. The first reason is the distribution of relaxation times that leads to a stretched exponential (exp (-t / s) n). The second explanation comes from the changes in the state of the material during relaxation. The relaxation times depend on these states. It is not thermo-rheologically simple: the relaxation function describing the kinetics of physical aging at a given temperature follows an extended exponential Fv = exp (t/s) n. n depends on the temperature at which the kinetics are performed, and under these conditions, the material is said to be thermorheologically simple. The evolution of the system out of equilibrium towards a more stable state depends on the aging temperature and the thermal history, which determines the state of the material at the beginning of the kinetics. The vitreous state cannot be described by a single order parameter. The experiments of Kovacs [16], which involved ‘‘memory effects,” showed that the evolution of a system having a distribution of relaxation times depends on its entire thermomechanical history. Thus, a vitreous state cannot be characterised by a single parameter (volume or enthalpy), because the effects can only be explained by a wide distribution of relaxation times. These results indicate that the vitreous transition temperature is an important characteristic for an epoxy but that this value strongly depends on the thermal history of the material. It is therefore important to clarify the interaction between the curing conditions of the material and the thermomechanical properties of the material. In this study, four commercial adhesives were used. The samples were cured isothermally at different temperatures (5, 10, 20, and 40 °C) for a period of 14 days to reproduce variable environmental conditions such as those for onsite applications in different seasons. The objective was to examine not only the influence of the curing conditions on the mechanical and physical properties of the epoxy resin but also the mechanical behaviour of the adhesively bonded CFRP-to-concrete joint.

2. Solution method Four structural cold-curing epoxy adhesives from four manufacturers were investigated. The properties of the adhesives are presented in Table 1. The epoxy system was mixed at room temperature respecting ratio by weight of the respective con-

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A.T. Lee et al. / Construction and Building Materials 227 (2019) 116385 Table 1 Adhesives used in the study* Adhesive

Manufacturer

Type

Density (kg/m3)

Mixing ratio

Pot life (min)

Tg (°C)

Resin Resin Resin Resin

A B C D

epoxy epoxy epoxy epoxy

– 1,300 1,150 –

2.5:1 04:01 2.33:1 2.33:1

100–130 90 (10 °C); 30 (35 °C) 45 80

80–86 >40 53.5 64

1 2 3 4

*According to manufacturer.

stituents (resin and hardener). All the adhesives were used to impregnate FRP fabrics that are employed to strengthen existing concrete or steel structures. The concrete surface was sanded before the FRP bonding. Specimens of each resin were preserved for a curing time of 14 days in four storage chambers, which had the following curing temperatures: 5 °C (curing condition mode 1), 10 °C (curing condition mode 2), 20 °C (curing condition mode 3), and 40 °C (curing condition mode 4). After 14 days, all the remaining specimens are preserved at 20 °C, that to say in the same condition than mode 3 (Fig. 1). To evaluate the physical properties (Tg) and the mechanical properties (stiffness, strength, normal adhesive ability, shear adhesive ability, and mechanical behaviour of the bonded joint in shear), tests were performed after 7, 14, 21, 28, 42, and 220 days. 2.1. Glass transition temperatures Tg The determination of the glass transition temperature Tg is a common method for describing the curing propagation and relating it to the mechanical properties: as curing advances, the molecular network mobility decreases, and the Tg increases [17,18]. The glass transition temperature Tg is defined as the point approximately at the middle of the temperature range in which the glass transition occurs. Glass transition is usually investigated via differential scanning calorimetry (DSC). In this study, another thermal analysis technique is used: thermomechanical analysis (TMA). Because they involve different testing methods, DSC and TMA provide slightly different results for identical curing conditions.

- A Mettler Toledo TMA/SDTA 1 IC/600 instrument equipped with a quartz glass probe with a ball point of 3 mm was used for the TMA. Samples 5  5  2 mm3 in size were subjected to a constant force of 0.02 N, under nitrogen purging at a flow rate of 20 mL/min, while the temperature increased from 5 to 150 °C at a heating rate of 10 °C/min. The glass transition temperature obtained via thermodilatometry (TMA), TgTMA, is defined as the intersection of the tangents with the length/temperature curve before and after the glass transition phase according to the standards ASTM E 1545-00 [19] and ISO 11359-2 [20] (Fig. 2a). - The experiments were performed using a Mettler Toledo TGADSC3 + differential scanning calorimeter. Epoxy samples of 10–15 mg were placed in an aluminium pan covered with a lid. An empty aluminium pan was used as a reference during every scan. Each DSC experiment was performed under nitrogen purging at a flow rate of 20 mL/min, while the temperature was increased from 5 to 150 °C at a heating rate of 10 °C/min. Data acquisition was performed using the accompanying software (STARe Evolution). The Tg is observed as an endothermic stepwise change in the DSC heat flow. As shown in Fig. 2b, three values according to the standards ASTM E 1356 [21] and NF EN 12614 [22] were considered for the glass transition temperature: TgONSET is defined as the temperature corresponding to the intersection of tangents from the baseline and initial slope of the step, TgMIDPOINT the middle of the step measured as half of the step height, and TgINFLECTIONPOINT is the inflection point of the step [23].

Fig. 1. Curing temperatures for the experimental campaign.

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A.T. Lee et al. / Construction and Building Materials 227 (2019) 116385

a) Evolution of the length with the increasing temperature according to TMA

b) Evolution of the heat flow with the increasing temperature according to DSC (ASTM E 1356) Fig. 2. Illustration of the TMA and DSC responses.

2.2. Durometer hardness test The hardness of the material represents the level of indentation of the surface against the penetration of a harder body. There are 12 scales of hardness. For elastomers, two widely used hardness scales are Shore A hardness (flexible elastomer) and Shore D hardness (rigid elastomer). The result of the durometers is in the interval of [0, 100]. This result is determined by the penetration depth of the needle at the scale of hardness studied immediately after force application or after 15 s of force application. A stiffer material yields a higher hardness result. The hardness scale used in this study is the Shore D hardness because the epoxy resin is rigid after crosslinking. The load applied to the penetration needle of the Shore D scale is 44.64 N (corresponding to a mass of 4.55 kg), and the applied load for the Shore A scale is 8.064 N. The Shore D hardness result depends on the depth of the indenter. If the needle penetrates 2.54 mm or more, the durometer shows that the hardness of the Shore D material is 0. If the needle cannot penetrate the surface of the material, the hardness on this scale is 100. This is why there are several scales of hardness; a material may be too hard (or too soft) for one scale but suitable for another.

2.3. Mechanical properties Via tensile tests of a dog-bone-shaped sample according to the standard NF EN ISO 527-1 [24], two mechanical properties of

epoxy are investigated: the Young’s modulus and the tensile strength. The tensile load, which increases at a speed of 1 mm/ min, is controlled during displacement. The strain at the middle of the specimen is measured by a strain gauge, which has a resistance of 120 O. The Young’s modulus is determined by the slope of the stress–strain curve, and the tensile strength is calculated using the ultimate load and measured width (around 10 mm) and thickness (around 3.8 mm) of the specimens. For each set of specimens, five specimens are tested. The adhesive strength of a cylindrical specimen with a diameter of 50 mm is measured via a pull-out test according to the French standard P18-852 [25]. Here, only polymer 4 is tested. The concrete is a C40/50, meaning that the minimal compressive strength measured on cylinder is 40 MPa. Once the FRP is cured, a core drill is used to core the concrete, as shown in Fig. 8. Then, an aluminium disc is stuck to the FRP, and after 24 h, a dynamometer is used for applying direct tension to the aluminium disc at a rate of 0.06 MPa/ s. The value and type of failure are the main parameters obtained. Five specimens are tested from each set. Polymer 4 is used, with the same carbon fibre sheets. The carbon is a bi-directional fabric 70/30 (70%/30% for weft and warp weight ratio) with a weight by square meter of 520 g/m2. At rupture, tensile stress and tensile strain are respectively 4900 MPa and 2.10%. Young modulus E is 230 GPa. The double-shear test allows the mechanical behaviour of the CFRP–concrete joint, as well as the adhesion ability of the resin, to be studied according to the curing conditions [25,26]. The tests are performed using a universal tensile machine Zwick 1475 at a cross-head speed of 1 mm/min until failure. The test was carried out at 20 °C. The dimensions of a test specimen are presented in Fig. 3. According to previous studies [25], an unbonded length (L2) of 40 mm is retained. During the test, two linear variable differential transformers (LVDTs) are set up at two opposite sides. The average value provided by the LVDT indicates the relative distance between the two concrete blocks (DL1). The strain gauges (J1–J7) are bonded along the composite strip in order to determine the strain of the studied points of the joint. The layout of the gauges is presented in Fig. 3. The gauges used in this test have a length of 10 mm, and the maximum captured strain is 1.5%. Two gauges called J1 and J2 bonded to two opposite sides in the middle of the composite strip allow the lengthening of the composite in the loading state to be calculated. e1 and e2 represent the strains recorded by the first and the second gauges, respectively. The lengthening of the composite (DL2) during the loading is expressed as

DL2 e1 þ e2 ¼ e ) DL 2 ¼  L0 L0 2

ð1Þ

where L0 represents the initial length (initial distance between the two concrete blocks). The total spacing of the concrete blocks (DL1) is due to the lengthening of the composite (DL2) during the loading state and the slip of the resin (DL). The slip of the joint can be determined using the data provided by the LVDT and strain gauges:

DL ¼ DL1  DL2

ð2Þ

The other gauges are distributed along composite strip (from J3 to J7). The distance between 2 consecutive gauges is 25 mm. The determination of the strain throughout the joint allows the effective bonding length (the length of the transfer of the load) to be determined. The effective zone of anchoring is the zone in which almost all of the local shear stress of the adhesive bonded joint is distributed [27]. The local bond stress s is given as

s¼


Ef :Def :ðbf :t f Þ Dr:Af DF ¼ ¼ Abond b f : Dx b f : Dx

ð3Þ

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A.T. Lee et al. / Construction and Building Materials 227 (2019) 116385

Fig. 3. Shear-test specimen and its dimensions. Table 2 Resins and composites tests layout. Tests

Curing conditions Mod 1 7d

14d

DSC R3 R4

TMA

R3 R4

Durometer Hardness R4 R4 Tensile Test

Pull-out Double lap shear test

R1 R2 R3 R4

Mod 2 21d

28d

42d

210d

R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4 R2 R3 R4

R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4

R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4

R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4

7d

14d R3 R4

R3 R4

R4 R4

R1 R2 R3 R4

R4 R4

Mod 3

Mod 4

21d

28d

42d

210d

7d

14d

21d

28d

42d

210d

7d

14d

21d

28d

42d

210d

R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4 R2 R3 R4

R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4 R4 R4

R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4

R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4

R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4

R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4 R4 R4

R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4

R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4 R4 R4

R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4

R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4

R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4

R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4 R4 R4

R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4

R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4 R4 R4

R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4

R1 R2 R3 R4 R1 R2 R3 R4 R1 R2 R3 R4

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A.T. Lee et al. / Construction and Building Materials 227 (2019) 116385

Fig. 4. Heat flow with respect to the reference temperature at 28 d with different curing temperatures for resin 1.

Fig. 5. Glass transition temperature Tg evaluated via DSC and TMA for resin 1.

s ¼ Ef :t f :

Def Dx

ð3’Þ

where s represents the local bond stress (MPa), Ef represents the CFRP Young’s modulus (MPa), t f ¼ 0:48mmrepresents the thickness of the CFRP, Def represents the difference in strain between the two strain gauges, and Dx ¼ 25mm represents the distance between the two strain gauges. Polymer 4 is used with the same as composite 3 carbon fibre sheets. The carbon is the bi-directional fabric 70/30 previously described. Three specimens are tested for each curing condition. 3. Results and discussion All the tests were done on the four resins under four different curing conditions with at least four deadlines. Table 2 summarizes all the tests carried out.

3.1. Glass transition temperatures Tg The Tg is determined using two different devices. The first one (DSC) determines the Tg by observing the profile of the heat flow during the change of state, and the second one (TMA) identifies the Tg by benefitting the coefficient of thermal expansion (CTE) at the time of transition. The objective of this part is to find the correlation between the Tg value determined via DSC and TMA. In an experimental way, the Tg measured via TMA having the nearest value with Tg inflection point measured by the DSC. For example, for resin 1, the heat flow and Tg after 28 d of curing at various curing temperatures are presented in Figs. 4 and 5, respectively. Theoretically, the Tg inflection point, which is measured via DSC, is the transition temperature at which the speed of the variation of the calorific flow according to the temperature is maximised. In this case, the first heating run is presented in Fig. 4. Furthermore, the Tg measured via TMA is the temperature at which the rate of increase of the CTE is in transition. At the time of the transition, the polymer strongly absorbs energy for the

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A.T. Lee et al. / Construction and Building Materials 227 (2019) 116385

a) resin 1

c) resin 3

b) resin 2

d) resin 4

Fig. 6. Evolution of the glass transition temperature Tg with respect to time under different curing conditions.

80

70

polymer 2 polymer 3 polymer 4

Tg (°C)

60

50

40

30 30

40

50

60

70

Shore D hardness value Fig. 7. Tg versus the Shore D hardness value.

endothermic process, which involves a considerable change of the mechanical properties, including the thermal dilation coefficient. In Fig. 6, the Tg inflection point is used to draw the evolution curve of the glass transition temperature Tg for the four resins with respect to time under different curing conditions. For all the studied resins, no results are obtained at 5 °C and 7 d of curing, because

the samples do not receive sufficient energy to activate the chemical reaction for crosslinking. After 14 d of curing at 5 and 10 °C, some resins (resin 1 and resin 2) are not cured yet (Tg < 30 °C), while the two other resins (resin 3 and resin 4) are cured (Tg > 30 °C). Thus, the minimum curing time at a low temperature (Tcure  10 °C) differs among the resins. For the category

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A.T. Lee et al. / Construction and Building Materials 227 (2019) 116385

4000 3500

Young Modulus (MPa)

3000 2500 2000

polymer 1 polymer 2 polymer 3

1500

polymer 4

1000 500 0 5 °C 28 days

10 °C 21 days

10 °C 28 days

20 °C 14 days

20 °C 28 days

40 °C 14 days

40 °C 28 days

Curing conditions

(a) Young modulus 70

60

Strength (MPa)

50

40 polymer 1 polymer 2

30

polymer 3 polymer 4

20

10

0 5 °C 28 days

10 °C 21 days

10 °C 28 days

20 °C 14 days

20 °C 28 days

40 °C 14 days

40 °C 28 days

Curing conditions

(b) Tensile strength Fig. 8. Polymere tensile properties.

‘‘low curing temperature” (where Tcure does not exceed 10 °C), Tg increases with time: the minimum required curing time at 5 and 10 °C is longer than 2 weeks (Fig. 6). For the categories ‘‘medium curing temperature” and ‘‘high curing temperature” (Tcure of 20 and 40 °C, respectively), after the seventh day, the Tg stabilises over time. This phenomenon can be explained by the crosslinking of the epoxy resin at the ambient temperature. A higher ambient temperature yields a larger chemical reaction between the epoxide group and the amine group in the resin. This involves an evolution

of Tg according to the curing temperature. B. Ellis [28] explained that the reaction rate decreases and becomes diffusion-controlled when the curing temperature is lower than Tg. Furthermore, F. Lapique [6] suggested that with a curing temperature lower than Tg, the adhesive is not fully cured, and the evolution of Tg reaches a plateau-like regime as early as 7 d. Importantly, for all the tested polymers, the value of the glass transition temperature at a curing temperature of 40 °C is never obtained, even after 240 d of post-curing at 20 °C when the initial

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curing temperature is 20 °C or lower, this result is important to know when Tg properties is given for a polymer. Otherwise, this study has also allowed to show that if the initial curing temperature is lower than 20 °C (even if it is 5 °C), the final glass temperature exhibits the same plateau for a longer curing period. For example for Tg of polymer 1 an initial value of 48 °C is obtained for 5 °C during 14 days plus 14 days at 20 °C while a Tg value of 50 °C is obtained when the curing conditions are 28 days at 20 °C (Fig. 5). These results are similar for all polymer [29]. Curing conditions decrease the final Tg value if an perfect curing is not done during the first hours. 3.2. Results of durometer hardness test The hardness of the four resins under different curing conditions is studied. In some cases of handling, there is a large difference in hardness between t = 0 and t = 15 s because the resin does not have time to cure at low temperatures (7 d at 5 °C, for example). If the difference between the instantaneous hardness and the 15-s hardness is >10 units, we can intuitively conclude that the value of the hardness is unreliable. The hardness of the resin is measured using the Shore D durometer, according to the hardness scale for the rigid polymers. The hardness measurement results are presented in the Fig. 7 and in Tables 3 and 4. Some hardness values are missing, especially for the first resins (resin 1 and resin 2), because of the low degree of crosslinking, which leads to excessive variability of the measurements (difference of >10 units). The hardness values of different resins subjected to the same curing condition are close but different. This is shown by comparing the hardness of four resins studied at 20 °C 42 d after the curing. In general, the curing temperature influences the hardness of the resins. A higher curing temperature yields greater evolution of the hardness (for the same curing duration). For example, the evolution of the hardness for resin 2 is presented in Table 3. The hardness also depends on the curing time. For samples left at a low temperature (5 and 10 °C), the hardness value is low for a short curing time (7 d and possibly 14 d), and there is a large difference between the instantaneous hardness and the hardness recorded at 15 s. This difference can reach 20 units in some cases, which is explained by the flexibility of the material (under severe curing conditions) allowing the Shore D durometer needle to penetrate easily, causing the penetration to increase over time.

The specimens placed under severe conditions have a low degree of crosslinking, leading to low hardness and a large difference between the instantaneous measurement and the measurement at 15 s. For the samples left at a high temperature (20 and 40 °C), the hardness is stabilised after 7 d of curing. This means that the minimum time required for curing is less than one week, while the minimum curing time at 10 °C may be less than 2 weeks. Table 3 shows the stabilisation of the hardness for resin 4 at 20 °C over time. The hardness depends on the thermal history of the resin. This observation can be obtained based on the results obtained on the hardness of the four resins under three different curing conditions: (20 °C for 14 d), (10 °C for 14 d first and 20 °C for 14 d later) and (40 °C for 14 d first and 20 °C for 14 d later). In general, the 15-s hardness at (10 °C for 14 d first and 20 °C for 14 d later) is identical to the 15-s hardness at (20 °C for 14 d). The first curing condition requires more time (4 weeks) to obtain a hardness that is equal to that obtained via curing at 20 °C for 2 weeks. The hardness obtained at a low temperature (10 °C) is equal to a value obtained at higher temperature (20 °C) for a longer time. In addition, the 15-s hardness at (40 °C for 14 d first and 20 °C for 14 d later) is significantly higher than the 15-s hardness at (20 °C for 14 d). This means that the hardness is an irreversible property size and depends on the thermal history of the resin, especially the highest temperature in its history. 3.2.1. Correlation between Shore D Hardness and Tg The objective of determining the hardness–Tg correlation is to estimate one value when the other value is known. Specifically, we can estimate the glass transition temperature of the resin in the field using a Shore D durometer, which is relatively easy to use compared with other devices for determining the Tg. We compare the Tg determined using the durometer with the service temperature in civil engineering and then decide on the commissioning of the joint. Table 4 presents the hardness and Tg (inflection point) of three resins at 14 d of curing. In the temperature range (5, 40 °C), the Tg of a resin can be determined via interpolation (Fig. 7). For example, the Tg of resin 2 subjected to a curing temperature of 30 °C for 14 d (Shore D hardness is 56) is given as follows.

T g ¼ 1:68  Shore D hardness value  39:7 

¼ 1:68  56  39:7 ¼ 54:4 C

Table 3 Results of the hardness test. Durometer hardness D (15 s)

R1

R2

R3

R4

20 °C 14 d 20 °C 42 d 10 °C 14 d + 20 °C 14 d 10 °C 14 d + 20 °C 14 d 40 °C 14 d + 20 °C 14 d 5 °C 14 d 10 °C 14 d 20 °C 14 d 40 °C 14 d

44 55 48 48 54 – – – –

53 58 53 53 62 37 43 53 62

52 53 45 45 66 – – – –

57 61 59 59 64 – – – –

Table 4 Comparison between Tg and hardness, after 14 d.

Hardness Tg (inflect) (°C) Hardness Tg (inflect) (°C) Hardness Tg (inflect) (°C)

5 °C

10 °C

20 °C

40 °C

! ! 45 38.3 ! !

43 32.3 ! ! 56 43.7

53 49.9 52 53,9 57 52.6

62 64.2 65 72.2 62 65.8

resin 2 resin 3 resin 4

ð4Þ

Obviously, the Tg and hardness depend on several factors (curing time, humidity, etc.). This formulation is developed to obtain a more precise value of Tg = f (hardness, cure, curing time, humidity, etc.). 3.3. Mechanical properties: Tensile properties of polymer The results of the tensile tests, which are presented in Table 5, indicate the evolution of the Young’s modulus and the strength with time under different curing conditions. Three samples are tested for each curing condition and time. After 14 d at 5 and 10 °C, no resin is strong enough to be demoulded and tested. After 21 d, only resins 2, 3, and 4 can be tested. The faster kinetics of the increase in strength for resins 2, 3, and 4 is in agreement with the evolution of Tg. For all the resins, after 28 d, the strength increased with the curing temperature. The development of the epoxy E-moduli obtained through tensile tests is also presented in Table 5. Regarding the Young’s modulus, no obvious conclusion can be drawn (Fig. 8). As shown by Savvilotidou et al. [30], the Emodulus mainly increases over the first 20 d. Benedetti et al. [31] reported only a slight increase of the E-modulus with the increase

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A.T. Lee et al. / Construction and Building Materials 227 (2019) 116385

Table 5 Young’s modulus and strength evolution with respect to the curing conditions and time. Resin 1

Resin 2

Resin 3

Resin 4

Curing temperature [°C]

Age [d]

Strength [MPa]

Modulus [MPa]

Strength [MPa]

Modulus [MPa]

Strength [MPa]

Modulus [MPa]

Strength [MPa]

Modulus [MPa]

5 °C 5 °C 10 °C 10 °C 20 °C 20 °C 40 °C 40 °C

21 28 21 28 14 28 14 28

– 13.3 – 23.2 16.6 31.9 31.5 44.3

– 3,288 – 2,436 3,047 2,977 2,811 2,888

47.2 41.7 53.3 52 52.6 53.7 60.0 57.3

2,691 2,807 2,492 2,656 2,695 2,752 2,734 2,906

20.4 27.1 25.7 34.0 36.2 44.2 41.7 55.9

3133 3479 3443 3411 3317 3405 3351 2646

23.8 19.4 22.7 21.4 24.4 25.7 20.9 29

2,975 3062 2520 3143 3227 3301 2849 3055

of the curing temperature (20, 30, and 40 °C) at 144 h. Same observation is obtained with our study, 10% decrease in mechanical properties are obtained if the curing conditions are lower than 20 °C, strength is more influence than Young modulus by the curing conditions, but the difference is low. 3.4. Mechanical properties: Pull-out test For the adhesion of a CFRP made with resin 4 and concrete, the results obtained from the pull-out test in all the curing conditions indicate an identical failure mode: failure in the concrete support

(Fig. 9). For the studied curing conditions (temperature ranging between 10 and 40 °C for 14 and 28 d), a cohesive concrete failure mode is obtained for all the tests (Fig. 9). The tensile strength of the concrete obtained via a splitting (Brazilian) test is fctm = 2.62 MPa. This result is close to those obtained via the adhesive test, which vary between 2.47 and 2.80 MPa on average (Tables 6 and 7). We discuss the results as follows. First, the results obtained under all the curing conditions indicate an identical mode of failure: cohesive failure in the support with tensile concrete failure. Thus, for the curing conditions

Fig. 9. Pull-out strength at different curing temperatures and the failure mode for resin 4.

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A.T. Lee et al. / Construction and Building Materials 227 (2019) 116385 Table 6 Pull-out test results after 14 d of curing for resin4. Pull-out strength [MPa]

40 °C 20 °C 10 °C

Pull-out strength [MPa]

1

2

3

4

5

Average

Standard deviation

2.72 2.89 2.46

2.26 3.32 2.84

2.56 3.05 2.16

2.45 2.34 2.38

2.72 2.39 2.64

2.54 2.80 2.50

0.20 0.42 0.26

Table 7 Pull-out values after 14 d of curing and 14 d at 20 °C, for resin 4. Pull-out strength [MPa] 1

2

3

4

5

Average

Standard deviation

2.14 2.66 3.18

2.08 2.72 2.62

2.80 2.77 3.00

2.82 2.65 2.22

2.53 2.94 2.70

2.47 2.75 2.75

0.35 0.12 0.37

(a) 14 d 6

5

Local Bond stress (MPa)

40 °C 20 °C 10 °C

Pull-out strength [MPa]

4

3

10°C 20°C 40°C

2

1

0

0

0.1

0.2

0.3

0.4

0.5

0.6

Slip (mm)

(b) 28 d Fig. 10. Local bond stress–slip curves for resin 4.

0.7

0.8

0.9

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A.T. Lee et al. / Construction and Building Materials 227 (2019) 116385

Table 8 Double lap shear test results for resin 4. Pu (kN) 14 d + 14 d at 20 °C

Pu (kN) 14 d

Sample 1 Sample 2 Sample 3 Average Standard deviation

10 °C

20 °C

40 °C

10 °C

20 °C

40 °C

32.1 30.45 29.42 30.66 1.35

27.675 32.45 30.40 30.18 2.40

36.55 34.40 32.30 34.42 2.13

31.175 32.93 29.42 31.17 1.75

30.15 31.12 29.15 30.14 0.99

30.45 31.24 29.65 30.45 0.80

studied (temperature ranging between 10 and 40 °C for durations of 14 and 28 d), the adhesion capacity of the polymer is guaranteed, and there is no decohesive failure at the composite–support interface. Second, the strength capacity is not affected by the curing. The pull-out strength depends on the concrete quality and varies between 2.47 and 2.80 MPa on average. 3.5. Mechanical properties: Double lap shear test For three curing conditions, double lap shear tests are performed on resin 4. Three samples are tested for each series. Six specimens are prepared for each curing condition: three are tested immediately after each 14 days and specific curing conditions, and three are kept at 20 °C for 14 d after each curing conditions for post-curing and then tested. The ultimate load capacity is close to 30 kN for all the specimens. The standard deviation ranges from 0.80 to 2.40 kN depending on the curing conditions. The initial 40 °C curing conditions yield an ultimate force capacity of 34 kN, which is approximately 10% higher than that for the other curing conditions (Table 8). This value is not sufficient for drawing a conclusion regarding the effect of the curing conditions on the ultimate capacity. For all the specimens, concrete failure occurred along the adhesive joint. The strain distribution along the bonded joint allows the local shear bond stress to be calculated. According to the difference of deformation between the two consecutive gauges, the mechanical properties, and the dimensions of the joint, the s-s profile of the bonded joint can be established. The s-s relationship is presented in Fig. 10. The results are analysed according to the curve of the local bond stress versus the slip displacement in the adhesive joint. Fig. 10 clearly shows that the curing temperature influences the mechanical behaviour of the joint. First, by observing the first slope of s-s, which represents the stiffness of the joint, one can conclude that the lap joint is stiffer if it is exposed to a high curing temperature (40 °C). In addition, the local stress at the loaded end of the polymer exposed to a high or low curing temperature is not significantly modified. Fig. 10a shows that the maximum stress at 40 °C at the loaded end is 5.5 MPa, while that at 10 °C is only 3.8 MPa. Fig. 10b shows the behaviour of the joint 28 d after pouring. The mechanical behaviour is identical even if the pre-curing temperature is different. The mechanical properties of the joint are poor in the short term but can improve over time. In conclusion, although if the local shear stress distribution is affected by the curing conditions, the global behaviour is not significantly affected, and the adhesive strength remains the same regardless of the curing conditions. The obtained results are very similar to results previously obtained by several authors [31–34] under static loading and ambient conditions with fully cured polymer. 3.5.1. Bonding length Fig. 11 presents some of the results obtained from the doubleshear test: for a load level of 20 kN, the evolution of the local strain

along the bonded joint is shown. After 14 d of curing, the curing temperature clearly influences the mechanical behaviour of the joint. First, the local strain, which represents the stiffness of the joint, increases with the curing temperature, as shown by Caggiano & Schicchi [32] and Hoult [33]. Obviously, the joint is less stiff if it is exposed to a low curing temperature (10 °C). The effective bonding length Le is very similar regardless of the curing temperature, as shown in Fig. 11a: Le is approximately 75 mm at 10 and 20 °C and reaches 80 mm at 40 °C. Importantly, even if the local strain depends on the curing temperature at 14 d, failure occurs in the concrete layer under all conditions. After 28 d, regardless of the curing temperature, the strain distribution along the bonding length is similar. We can conclude that the bonding length Le—approximately 75 mm—does not depend on the curing conditions (Fig. 11a and b). The analytical behaviour of the bonded joint is established according to the test results. Zhou and al. [35] proposed an analytical model for calculating the bond stress–slip relationship. The results compared are those for the specimen subjected to the most favourable curing conditions (40 °C pre-curing for the first 14 d + 20 °C curing for 14 d). To establish the s-s relationship for the joint, we must first determine all the parameters in the equation for s(s):

sðsÞ ¼

Ef :tf a sa s
: :e : 1  e a 1 þ q b2

ð5Þ

Here, Ef ¼ 74; 26 GParepresents the E-modulus of the composite, which is calculated using the data obtained from the test (deformation of the gauges between the two concrete blocks, force exerted over time, section of the joint), and tf ¼ 0:48mm is the thickness of the composite.

q¼

Ef :t f :bf 74; 26  0; 48  50 0 ¼ 35  70  140 Ec :t c :bc

ð6Þ

The coefficients a and b are determined using the test results, according to the load–slip profile. From the bilinear curve, the coefficient K0, which corresponds to the slope of the first line, is given as follows.

a ¼ sm ¼ K0 ¼

b¼

Pmax 30000 ¼ 0:40 ¼ 2  75000 K0

Pmax 30000 ¼ 75000 ¼ 2  0:20 sm

K0 75000 ¼ 40:85 ¼ Ef  t f  bf 85  0:48  2  50

ð7Þ

ð8Þ

ð9Þ

The calculated slope has uncertainty. In this case, K 0 varies in an interval of 70,000 to 80,000. The coefficients a and b are therefore determined in the intervals according to the uncertainty of the slope K 0 Table 9. The coefficients a = 40 and b = 40.85 obtained via regression conform well to the behaviour of the joint. The beha-

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A.T. Lee et al. / Construction and Building Materials 227 (2019) 116385

8000 20 kN 40°C 20 kN 20°C

Strain (µm/m)

6000

20 kN 10°C

4000

2000

0

0

20

40

60

80

100

120

140

Bonded length (mm)

b) 14 d 20 kN 40°C at 14 days + 20°C at 14 days

8000

20 kN 20°C at 28 days 20 kN 10°C at 14 days + 20°C at 14 days

Strain (µm/m)

6000

4000

2000

0

0

20

40

60

80

100

120

140

Bonded length (mm)

b) at 28 d Fig. 11. Evolution of the local strain with respect to the distance from the loaded end for a load of 20 kN for resin 4.

Table 9 Double lap shear test model parameters. K0

75; 000

a

0.40 40.85

b

viour of the bonded joint is similar to a previously reported analytical model [36] (Fig. 12). The curing conditions tested in this study do not modify the existing bond-slip model under static loading. 4. Conclusions According to the results presented in this paper, the following conclusions are drawn with regard to the resins cured at low temperatures. - The glass transition temperature Tg measured via TMA had the closest value to the Tg inflection point measured via DSC.

- At a low temperature (Tcure  10 °C), the kinetics of the evolution of Tg differed according to the resins up to 14 d. - There was a relationship between the glass transition temperature and the hardness. - The evolution of the Tg according to time and the curing temperature was in agreement with the evolution of the tensile strength. The molecular network mobility, which depended on the curing time and curing temperature, guided the evolution of Tg and thus the evolution of the mechanical properties. More precisely, a greater degree of crosslinking, which was obtained at a high curing temperature (40 °C), provided a more complete reaction (Tg_tcure = 40 °C > tg_tcure = 20 °C) and consequently improved the mechanical properties (tensile strength). - A major application of the cold-curing resin examined in this study is to bond CFRP strips and sheets to concrete. In a basic pull-out test, regardless of the resin used, failure occurred on concrete. This shows that the CFRP–concrete interface is not the weak link. The bonding ability, which was indicated by the bonding length Le obtained in a double-shear test, was very

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A.T. Lee et al. / Construction and Building Materials 227 (2019) 116385

6 40°C at 14days + 20°C at 14 days 20°C at 28 days

Local bond stress (MPa)

5

10°C at 14days + 20°C at 14days Analycal model

4

3

2

1

0 0

0.2

0.4

0.6

0.8

Slip (mm) Fig. 12. Comparison between the model and experimental results for resin 4.

similar regardless of the curing conditions. The curing conditions tested in this study do not modify the local bond slip model proposed in the literature. In a future study, it would be interesting to evaluate the effect of the curing conditions on the durability of the material. Even if the Tg value is influenced by the curing conditions, the present study did not highlight the significant modification of the mechanical properties, instead focusing on physical values (Tg and hardness). Declaration of Competing Interest The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper. References [1] C. Lawrence Bank, Properties of FRP Reinforcements for Concrete, FiberReinforced-Plastic (FRP) Reinforcement for Concrete Structures, 1993, Pages 59-86. [2] A. Benedetti, P. Fernandes, J.L. Granja, et al., Influence of temperature on the curing of an epoxy adhesive and its influence on bond behaviour of NSM-CFRP systems, Compos. Part B Eng. 89 (2016) 219–229. [3] J. Michels, R. Widmann, C. Czaderski, et al., Glass transition evaluation of commercially available epoxy resins used for civil engineering applications, Compos. Part B 77 (2015) 484–493. [4] O. Moussa, A. Vassilopoulos, T. Keller, Effects of low-temperature curing on physical behaviour of cold-curing epoxy adhesive in bridge construction, Int. J. Adhes. Adhes. 32 (2012) 15–22. [5] M. Savvilotidou, A.P. Vassilopoulos, M. Frigione, T. Keller, Effects of aging in dry environment on physical and mechanical properties of a cold-curing structural epoxy adhesive for bridge, Constr. Build. Mater. 140 (2017) 552–561. [6] F. Lapique, K. Redford, Curing effect on viscosity and mechanical properties of a commercial epoxy resin adhesive, Int. J. Adhes. Adhes. 22 (2002) 337–346. [7] K. Matsui, Effects of curing conditions and test temperatures on the strength of adhesive-bonded joints, Int. J. Adhes. Adhes. 10 (1990) 277–284. [8] Externally bonded FRP reinforcement for RC structures, N° 14. Externally bonded FRP reinforcement for RC structures. Technical report (138 pages, ISBN 978-2-88394-054-3, October 2001) [9] Guide for the Design and Construction of Externally Bonded FRP Systems for Strengthening Existing Structures Materials, RC and PC structures, masonry structures, CNR-DT 200 R1/2013, ROMA – CNR October 10th 2013 – release of May 15th 2014

[10] E. Ferrier, G. Lagarde, P. Hamelin, Concrete beams reinforced by fibrereinforced plastics: the effect of temperature on the adhesive layer, Compos. Sci. Technol. 61 (3) (2001) 425–431, https://doi.org/10.1016/S0266-3538(00) 00129-9. [11] Guide for the Design and Construction of Externally Bonded FRP Systems for Strengthening Concrete Structures, ACI 440.2R-08, ISBN 978-0-87031-285-4, July 2008 [12] C. Arya, J.L. Clarke, E.A. Kay, P.D. O’Regan, TR 55: Design guidance for stengthening concrete structures using fibre composite materials: a review TR 55: Design guidance for stengthening concrete structures using fibre composite materials: a review, Eng. Struct. 24 (7) (2002) 889–900, https:// doi.org/10.1016/S0141-0296(02)00027-5. [13] Y. Jeong, M.M. Lopez, C.E. Bakis, Effects of temperature and sustained loading on the mechanical response of CFRP bonded to concrete, Constr. Build. Mater. 124 (15) (2016) 442–452. [14] R. Quison, J. Perez, M. Rink, A. Pavan, Components of non-elastic deformation in amorphous glassy polymers, J. Mater. Sci. 31 (1996) 4387–4394. [15] L.C.E. Struik, Physical aging in amorphous polymers and other materials, Elsevier science Ltd., 1978, p. 229. P. ISBN 0444416552. [16] A.J. Kovaacs, J.J. Aklonis, J.M. Hutchinson, A.R. Ramos, Isobaric volume and anthalpi recovery of glasses. II. A transparent multi-parameter theory, J. Polym. Sci.: Polym. Phys. Ed. 17 (1979) 1097–1162. [17] J. Perez, J.Y. Cavaille, L. David, New experimental features and revisiting the a and b mechanical relaxation in glasses and glass-forming liquids, J. Mol. Struct. 479 (1999) 183–194. [18] O. Moussa, A. Vassilopoulos, T. Keller, Experimental DSC-based method to determine glass transition temperature during curing of structural adhesives, Constr. Build. Mater. 28 (2) (2012) 263–268. [19] ASTM E1545-00, Standard Test Method for Assignment of the Glass Transition Temperature by Thermomechanical Analysis, ASTM International, West Conshohocken, PA, 2000. www.astm.org. [20] ISO 11359-2:1999, Plastics – Thermomechanical analysis (TMA) – Part 2: Determination of coefficient of linear thermal expansion and glass transition temperature, Publication date : 1999-10, ICS : 83.080.01 [21] ASTM E1356–08(2014), Standard Test Method for Assignment of the Glass Transition Temperatures by Differential Scanning Calorimetry, ASTM International, West Conshohocken, PA, 2014. www.astm.org. [22] E.N. Nf 12614, Products and systems for the protection and repair of concrete structures - Test methods - Determination of glass transition temperatures of polymers ICS 91 080 April 2005 40 [23] Sidney H. Goodman, Book Chapter, 6: Epoxy Resins, Handbook of Thermoset Plastics (Second Edition), 1998, Pages 193-268. [24] E.N.I.S.O. Nf 527–1, Plastics - Determination of tensile properties - Part 1: general principles – Plastiques ICS 83 080 April 2012 01 [25] P18-852 April 1993, Special products for hydraulic concrete constructions. Synthetic resin based on hydraulic binders based products or systems for superficial applications on hardened concrete. Traction adhesive strength test on surface sawn support slab. [26] E. Ferrier, M. Quiertant, K. Benzarti, P. Hamelin, Influence of the properties of externally bonded CFRP on the shear behavior of concrete/composite adhesive joints, Compos. Part B: Eng. 41 (5) (2010) 354–362.

A.T. Lee et al. / Construction and Building Materials 227 (2019) 116385 [27] Huy Binh Pham, Riadh Al-Mahaidi, Modelling of CFRP-concrete shear-lap tests, Constr. Build. Mater. 21 (4) (2007) 727–735. [28] B. Ellis, Chemistry and Technology of Epoxy Resins, Chapman & Hall, 1993. [29] F. Ania, J. Martinez-Salazar, F.J. Baltá Calleja, Physical ageing and glass transition in amorphous polymers as revealed by microhardness, J. Mater. Sci. 24 (8) (1989) 2934–2938, https://doi.org/10.1007/BF02385650. [30] Maria Savvilotidou, Thomas Keller, Anastasios P. Vassilopoulos, Fatigue performance of a cold-curing structural epoxy adhesive subjected to moist environments, Int. J. Fatigue 103 (2017) 405–414. [31] A. Benedetti, P. Fernandes, J.L. Granja, et al., Influence of temperature on the curing of an epoxy adhesive and its influence on bond behaviour of NSM-CFRP systems, Compos. Part B Eng. 89 (2016) 219–229.

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[32] A. Caggiano, D.S. Schicchi, A thermos-mechanical interface model for stimuling the bond behavior of FRP strips glued to concrete substrates exposed to elevated temperature, Eng. Struct. 83 (2015) 243–251. [33] N.A. Hoult, J.M. Lees, Time-dependent behaviour of RC beams retrofitted with CFRP straps, ASCE J. Compos. Constr. 15 (1) (2011) 75–84. [34] J. Dai, T. Ueda, Y. Sato, Development of the nonlinear bond stress–slip model of fiber reinforced plastics sheet–concrete interfaces with a simple method, J. Compos. Constr. 9 (2005) 52–62. [35] Y.-W. Zhou, Y.-F. Wu, Y. Yun, Analytical modeling of the bond–slip relationship at FRP-concrete interfaces for adhesively-bonded joints, Compos. Part B Eng. 41 (2010) 423–433. [36] J.F. Chen, J.G. Teng, Anchorage strength models for FRP and steel plates bonded to concrete, J. Struct. Eng. 127 (2001) 784–791.