The fire performance of GFRP-RC beams based on FE thermal analysis

doi:10.21203/rs.3.rs-5000212/v1

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The fire performance of GFRP-RC beams based on FE thermal analysis

https://doi.org/10.21203/rs.3.rs-5000212/v1

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published 26 Oct, 2024

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GFRP (glass fiber reinforced polymers) rebars can be used in RC (reinforced concrete) structures as an alternative to conventional steel rebars. Their application offers advantages, especially in chemically aggressive environments, as they can reduce the C (concrete cover) thickness required and also optimize the cross-sectional dimensions of these structures as well as their durability and service life. However, based on FE (finite element) numerical analysis solved by Abaqus software, this research has shown that the reduction in C-thickness promotes a notorious incongruity: an improvement in the fire sensitivity of these structures. The time in which GFRP-RC structures failed in fire can be around 400% lower in relation to identical conventional steel RC structures. In some cases, the fire resistance rate (FRR) of GFRP RC structures can be less than 10 min when the critical temperature of these rebars (around 100°C) is taken into account. The interest in structural durability criteria is not sufficient to justify the application of this new type of reinforcement in some concrete buildings.

GFRP

Glass Fiber rebars

Glass Fiber Reinforcements

Glass Fiber Reinforced Concrete

Structures in fire

The service life of concrete structures has become increasingly important in recent years. The use of steel-reinforced concrete in areas with high chloride concentrations (e.g., coastal areas, seawater and de-icing salts) or chemical and radioactive wastes (e.g., in industrial facilities) can lead to corrosion of the steel rebars [1], resulting in concrete spalling [2] and weakening structural integrity [3], reducing the service life and structural performance [4]. Durability requirements must be considered in any structural design. When predicting service life, it is usually conservatively assumed that the service life ends when corrosion begins [5], which is typically solved in design by incorporating a minimum concrete cover thickness to protect the steel reinforcements [6], [7].

An alternative to solving the durability requirement is to use fiber reinforced polymer (FRP) rebars in place of the traditional steel [8]. The material is corrosion-resistant [9]. In addition, these composite materials are not only lighter than steel, but they also have higher tensile strength, electromagnetic transparency and low maintenance requirements [10]. FRP rebars enable thinner structural cross-sections with high reinforcement density by reducing the thickness of the concrete cover. Studies have demonstrated the application of FRP rebars in bridge decks [11], structural walls [12], industrial [13], maritime [14] and also infrastructure constructions [15] and others. Glass FRP (GFRP) is the most commonly used for concrete because it is less expensive and has acceptable mechanical properties for use in concrete structures [15], [16].

GFRP reinforcement is a composite material and are an attractive alternative to steel rebars to ensure long-term structural performance [16]. A low weight and high strength compared to conventional steel is an interesting advantage of this solution [17]. A high corrosion resistance, high strength-to-weight ratio, high stiffness-to-weight ratio, high degree of chemical inertia, excellent fatigue tolerance is another advantage to use the GFRP rebars (GFRP-R) [18]. The sustainability is a point of interest, measured by its associated economic and environmental impacts throughout its life cycle [16]. However, the benefits of GFRP-R instead of classical steel reinforcement are not exploited due to their vulnerability at high temperatures and under fire [14]. Mechanical properties of GFRP materials degrade significantly in this environment, leading to a serious and significant reduction in bearing capacity and brittle failure of GFRP-RC structures [19], mainly due to the tensile strength and elastic modulus damage [8].

Fire is one of the most serious potential risks for buildings and structures, and for this reason international codes provide specific guidelines to take account this scenario in the design of structures [18]. GFRP-R are more vulnerable to high temperatures than steel rebars [18], [20], [21], highlighting an issue to be addressed. The fire sensitivity of GFRP-R limits their wide use in construction. As seen in the steel reinforcements, the physical and mechanical properties of GFRP decrease at high temperatures. Material degradation is the key phenomenon, but it must meet the requirements of the relevant fire codes to be used as a structural solution. Most of these refer to the time available in a fire before the structural failure [22].

Extensive studies are available on the high temperature behavior of concrete reinforced with steel rebars, and their fire behavior is well known and also prescribed by many standardized procedures [23], [24], [25]. However, there is a gap in the literature regarding the analysis of concrete structures reinforced with GFRP-R subjected to fire. In the case of Brazil [23], as well as in the European standard [24], there are no proposed design procedures for the fire design of these structures. Due to the lack of knowledge, design guidelines (conservatively) do not recommend yet the use of GFRP rebars in structures where the fire action has to be considered at design (i.e., in buildings) [26].

There are many philosophies on the design of concrete structures in fire. The most usual is to limit the reinforcement to reach a critical temperature. This method is well accepted due to the low thermal diffusivity of concrete and makes the reinforcement the critical component of the structure. In the case of the steel rebars, this temperature is in order of 500°C, as proposed by fib Bulettin No.38 [27]. At critical temperature steel loses around half of its original strength, and the structure can no longer support the loads during a fire. An adequate fire design for reinforced concrete could be achieved by providing minimum dimensions and sufficient concrete cover thickness to protect the reinforcement [28]. Standards as CSA S806 [29] continue to use the temperature in the GFRP-R as a critical criterion to design the GFRP-RC structures, clearly based on the philosophy established to traditional RC structures.

The reduction in strength with increasing temperature varies considerably depending on the type of fiber and resin [30]. The tensile strength is significantly more affected by elevated temperature than the elastic modulus, but some authors highlight that the fire performance of these rebars is unclear [19]. Wang et al. [31] suggested a critical temperature based on 50% strength loss is 325°C. As the ultimate temperature, Abbasi and Hogg [32] proposed 462°C, while the experimental research done by Nigro et al. [33] highlight 460°C. Hajiloo et al. [13], Bilotta et al. [18], and Jin et al. [19] stablished 400°C as the ultimate temperature, but Rosa et al. [30] suggests 300°C to structural design (where the 50% of the ultimate condition is reached). Around 700°C there is a full decomposition of the resin [30].

However, the bond of the GFRP-R to concrete are damaged at 100°C, and this temperature need to be assumed as critical. Researches such as Katz et al. [34] concluded that the bond performance at ambient temperature, as well as the bond strength and stiffness degradation with temperature, depend mostly on parameters associated with the surface properties of the rebars. According to the authors, at 200°C the GFRP bond strength are reduced at least 80%, while in the steel rebars are only 38%. The literature about the reduction with temperature of the bond properties of GFRP-R to concrete is still very limited [8]. There are authors who believe that when continuous reinforcement from side to side of the RC element was used, not significant slips occurred at the end anchorage zones, and then the structural integrity was guaranteed [18].

Glass fibers can resist 800°C [35], but they do not perform their job perfectly in concrete elements without binding material (resin). In fact, according to Rosa et al. [14], the effects of elevated temperatures on the bond between GFRP rebars and concrete are more severe than the mechanical loss of these rebars. Studies have already shown that when the glass transition temperature of the polymeric matrix is reached (approx. 100°C), the bond strength between GFRP-RC decreases by 90 % [8], [36], [37]. Thee results raise concerns about the performance of GFRP-RC structures exposed to fire, as most of the bond strength of the reinforcements is likely to be lost in the early stages of a fire while temperatures are still moderate [14]. Studies show that GFRP-RC structures exposed to fire fail prematurely because the bond is lost along the overlapping length of the rebars [22].

The aim of the proposed study was to understand the fire behavior of GFRP-RC structures. A thermal analysis of a group of rectangular beam cross-sections was proposed. A group of RC beams with different width and concrete cover was assumed. The average temperature of the concrete and reinforcement was considered. The uniqueness of the proposed research is to compare a series of RC beams with different widths and heights and variable concrete cover thicknesses to determine and correlate the fire behavior and the FRR of GFRP and conventional steel reinforcements when exposed to fire.

2.1 Research overview

The study attempts to better understand the fire behavior of GFRP-RC structures and their correlation with steel RC structures. A thermal analysis of several rectangular beam cross-sections was presented. Abaqus software was used to solve a set of finite elements (FE) models. A variety of RC beam widths and heights and concrete cover thicknesses were assumed. The average temperature of the concrete and reinforcement (assuming conventional steel and GFRP) was identified as a crucial criterion. A comparison was done.

2.2 Cross-sectional thermal analyses

The concrete cover (C) thickness criterion considered in structural fire design ($\:{\text{C}}_{\text{f}}$) is not the same as the durability and service life requirements. The $\:{\text{C}}_{\text{f}}$ is measured from the edge of the beam to the center of longitudinal reinforcement, as shown in Fig. 1. It’s a standardized criterion. The section width (W) and height (H) were assumed as variables for the study. The thickness of the slab ($\:{\text{t}}_{\text{s}}$) was 100 mm in all cases. The cross-section shown in Fig. 1a establishes the results to define the sagging bending resistant moment (+) of the RC beams (the concrete temperature above the neutral axis and the positive steel reinforcements as used in the conventional design procedure). The same in Fig. 1b, but to the hogging moment (-).

The temperature time history in concrete (average between pt1 to pt5 control points) and reinforcement (average between Rb1 to Rb3 control points) temperatures was monitored to determine the thermal field of each selected cross section (Table 1). For the concrete temperatures, since the concrete cross-section at the bearing capacity of the beam is above (in case of sagging moment) or below (hogging moment) the neutral axis, only measurement points in this region were assumed (i.e., 0.5 x H). The concrete temperature is the average of the control points pt1(+) to pt5(+) (in case of the sagging moment) and pt1(-) to pt5(-) (hogging moment). The beam cross-sections considered in the study are shown in Table 1. The usual width and height of the beams (as is usual in engineering practice) were taken into account.

Table 1

– Beams assumed in the research
Cross-section characteristics and nomenclatures
Beam number	Nomenclature	Width W (mm)	Height H (mm)	$\:{\text{C}}_{\text{f}}$ thickness (mm)
1	B1w15h50	150	500	20 to 50
2	B2w20h50	200	500	20 to 50
3	B3w25h50	250	500	20 to 50
4	B4w30h50	300	500	20 to 50
5	B5w35h50	350	500	20 to 50
6	B6w20h60	200	600	20 to 50
7	B7w20h70	200	700	20 to 50
8	B8w20h80	200	800	20 to 50
9	B9w20h90	200	900	20 to 50
10	B10w20h100	200	1000	20 to 50

As shown in Table 1, beams B1 to B5 were designed to investigate the effect of beam width on the average temperature of longitudinal reinforcements and concrete, while keeping the beam height fixed at 500 mm. Beams B6 to B10, on the other hand, fixed the beam width at 200 mm keeping the beam height as variable.

2.3 Numerical procedures

Conduction is the primary heat transfer process, while convection and radiation are the mechanisms for heat transfer from the environment to the structural surfaces exposed to fire [38]. Eq. 1 shows that the thermal diffusivity is related to the mass loss $\:\rho\:$, the thermal conductivity $\:\lambda\:$ and specific heat $\:{C}_{p}$. In the case of concrete, these parameters are those proposed in EN 1992 − 1.2 [24]. In the case of reinforcement, the thermal parameters are not required in the model, as described in section 2.4.

$$\:{\alpha\:}=\frac{{\lambda\:}}{{\rho\:}.{\text{C}}_{\text{p}}}$$

The governing equation for the convective and radiative heat transfer analysis are defined in Eq. 2 [38], where $\:{\text{n}}_{\text{y}}$ and $\:{\text{n}}_{\text{z}}$ are the components of the vector outward normal to the cross-sectional surface; $\:{\text{h}}_{\text{r}\text{a}\text{d}}$ and $\:{\text{h}}_{\text{c}\text{o}\text{n}}$ are the radiative and convective heat transfer coefficient, respectively; T is the initial temperature (20°C) and $\:{\text{T}}_{\text{E}}$ is the temperature of the environment (assumed as ISO 834 [39] according to the EN 1992 − 1.2 [24] design procedure). The radiation heat transfer coefficient is given by Eq. 3, where $\:{\sigma\:}$ is the Stefan-Boltzmann constant (σ = 5.67 x 10^− 8 W/m².$\:{^\circ\:\text{C}}^{4}$), and $\:{\epsilon\:}$ the emissivity factor (0.70 according to EN 1992 − 1.2 [24]). The convective heat transfer 25 and 9 W/m²°C in the fire-exposed and unexposed surface.

$$\:\text{k}\bullet\:\left(\frac{\text{d}\text{T}}{\text{d}\text{y}}\bullet\:{\text{n}}_{\text{y}}+\frac{\text{d}\text{T}}{\text{d}\text{z}}\bullet\:{\text{n}}_{\text{z}}\right)=-\left({\text{h}}_{\text{r}\text{a}\text{d}}+{\text{h}}_{\text{c}\text{o}\text{n}}\right)\bullet\:\left(\text{T}-{\text{T}}_{\text{E}}\right)$$

$$\:{\text{h}}_{\text{r}\text{a}\text{d}}=4\bullet\:{\sigma\:}\bullet\:{\epsilon\:}\bullet\:\left({\text{T}}^{2}+{\text{T}}_{\text{E}}^{2}\right)\bullet\:\left(\text{T}+{\text{T}}_{\text{E}}\right)$$

Heating was applied to both lateral surfaces and one lower surface of the beam, as shown in Fig. 2.

The 3D model using thermal transient non-linear analysis is used to solved the problem. The concrete was modeled as DC3D8 finite element with classical integration from the ABAQUS library. A mesh analysis was performed and the mesh is rather refined, the size of the elements is approximately 1.0 x 1.0 x 1.0 mm. The time increment used was 0.001s and a convergence criterion of 0.1ºC.

The DC3D8 assumes a single degree of freedom (temperature) at each node along the x, y, and z axis. The solution used to solve the thermal problem is shown in Eq. 4, where the coefficients follow the same criteria and descriptions detailed in Eq. 1 to 3, where “t” is the time. The temperature distribution in the cross-section is obtained through conduction thermal analysis, and the heat transfer equation is formulated by applying the principle of conservation of the energy to a three-dimensional differential control volume.

$$\:\frac{\partial\:}{\partial\:\text{x}}\left({\lambda\:}\frac{\partial\:{\theta\:}}{\partial\:\text{x}}\right)+\frac{\partial\:}{\partial\:\text{y}}\left({\lambda\:}\frac{\partial\:{\theta\:}}{\partial\:\text{y}}\right)+\frac{\partial\:}{\partial\:\text{z}}\left({\lambda\:}\frac{\partial\:{\theta\:}}{\partial\:\text{z}}\right)={\rho\:}\bullet\:{\text{C}}_{\text{p}}\bullet\:\frac{\partial\:{\theta\:}}{\partial\:\text{t}}$$

Where $\:\frac{\partial\:{\theta\:}}{\partial\:\text{x}}$, $\:\frac{\partial\:{\theta\:}}{\partial\:\text{y}}$ and $\:\frac{\partial\:{\theta\:}}{\partial\:\text{z}}$ is the temperature $\:{\theta\:}$ gradient in the x, y and z-directions, and the $\:{\rho\:}\bullet\:{\text{C}}_{\text{p}}\bullet\:\frac{\partial\:{\theta\:}}{\partial\:\text{t}}$ is the time rate of change of the thermal energy of the medium per unit volume. In the case of the thermal boundary conditions, the exchange of convection and radiation during the heating process was considered.

2.4 Reinforcements assumptions

No reinforcement was used in the cross-section as its area is much smaller than that of concrete. The study focuses on the average temperature of the concrete cross-section and defines the reinforcement temperature assuming the concrete temperature in the same layer where the reinforcement is placed. Previous numerical FE studies by the authors, confirmed by full-scale experimental results (see [40], [41]), have already shown that the lack of reinforcement in the thermal model has little effect on the thermal field in the cross-section. The small cross-sectional area of the reinforcements compared to the total area of the concrete and the high thermal conductivity of the steel may contribute to the explanation.

2.5 Validation of the FEA model and numerical procedure

The FE model proposed in this research and the numerical procedure to solve the heat transfer problem (mesh size, steps, boundary conditions, integration process, governing equations, etc.) were validated by the authors through an extensive full-scale test (heating according to ISO 834 [39]) which they carried out in various studies. Some of these can be found in research papers such as [41], [42], [43]. The FEA model and numerical simulation were also based on the approaches highlighted by the authors in [44].

The temperature history in the concrete and in the positive and negative reinforcement is shown above. The temperature field in the cross-section and the fire resistance rate (FRR) of each beam, which are listed in Table 1, are also shown.

3.1 Concrete

The temperature history in the concrete is shown in Fig. 3.

The temperature above (Fig. 3a) and below (Fig. 3b) the neutral axis is described. The temperature refers to the average of the points pt1 to pt5 (see Fig. 1). The critical temperature of the concrete (approx. 350°C) is also highlighted in the curves. The temperature above and below the neutral axis is elementary for the determination of the sag or hogging moment, if the usual analytical expressions for determining the bending capacity of these structures are taking into account.

The average temperature in the cross-section of the beams naturally depends on their geometry. The most important factor influencing the average temperature of the concrete (above or below the neutral axis) is the width of the beam.

Above the neutral axis analysis: at the end of the numerical analysis, the temperature difference between B1 and B5 is about 300°C, as shown in Fig. 3a. Assuming the height of the beam, the difference between B6 and B10 is no more than 50°C. Only B1, B2 and B6 to B10 reached the temperature that characterizes critical mechanical damage of the concrete, i.e., only the beams with 15 and 20 cm width. Beams with a width of more than 20 cm did not reach the critical temperature of the concrete.

Below the neutral axis analysis: it is most affected by the high temperatures, as the three sides of this area of the beam are exposed to fire (see Fig. 2). As described in Fig. 3a, Fig. 3b also shows that the width of the beam has less influence on the temperature field in the cross-section than the height. At the end of the analysis, the difference between B1 and B5 (with a width of 15 and 35 cm respectively) is almost 400°C, while between B6 and B9 (with a height of 60 and 90 cm respectively) is no more than 50°C. It is also important to note that the critical temperature of the concrete is reached in all cases. B1 is the most critical condition, while B5 is the most stable in the fire. With regard to the concrete temperature, the width of the beams is decisive for their fire behavior.

The critical temperature of the concrete must be discussed in parallel with the reinforcement to predict the structural fire behavior. The positive and negative reinforcement temperatures are discussed below.

3.2 Positive reinforcements

Figures 4a, 4b, 4c and 4d show the temperature history in the reinforcements with C = 20, 30, 40 and 50 mm. The critical temperature of steel and GFRP reinforcements is also highlighted. The temperature history in the positive reinforcements is required to define and apply the analytical expressions for determining the sagging bending resistance moment in fire.

The average temperature of the reinforcement is directly dependent on the thickness of the concrete cover. In general, the temperature of the reinforcement decreases in the order of 100°C to each 10 mm increase in concrete cover. Since the critical temperature of conventional steel reinforcement is five times higher than that of GFRP reinforcement, the FRR of these beams naturally tends to be higher. Assuming that C = 10, 20 or 30 mm, the GFRP beams do not reach 20 min to standard fire, while the beams with conventional steel rebars reach about 40, 60 and 70 min, respectively. The FRR of the conventional steel RC beam is at least 80 min and can reach a maximum of 110 min, depending on the beam width. At 50 mm, the FRR of the GFRP RC beam is approximately 30 min.

It is important to note that in the most critical condition, the FRR of steel RC beams can exceed 80 minutes compared to GFRP RC structures. This condition highlights the safety of conventional steel RC structures in terms of fire performance and also the shortcomings that need to be discussed in order to apply GFRP reinforcement in conventional multi-storey buildings when fire safety requirements need to be met

3.3 Negative reinforcements

The temperature in negative reinforcements with C = 20, 30, 40 and 50 mm is shown in Figs. 5a, 5b, 5c and 5d respectively. These rebars are more thermally protected compared to the positive rebars due to the adjacent slab. These temperatures are necessary to define and apply the analytical expressions required to determine the hogging bending resistance moment in case of fire.

Figure 5 shows that high temperatures have no effect on steel reinforcement. This indicates that with the proposed cross-sectional dimensions of the beams (as given in Table 1), these reinforcements don't shown mechanical damage when the beam is exposed to fire. In contrast, the GFRP reinforcements with C = 20, 30, 40 and 50 mm reached their critical temperature at 90, 85, 80 and 75 min of fire exposure respectively. It's important to remember that the thickness C is counted for the surface of the beam. As the C thickness increases, the rebars are exposed to higher temperatures, which reduces their FRR tine.

3.4 Temperature field in the cross-section

Figures 6 and 7 show the temperature field in cross-section for two cases of beams when B5w35h50 and B7w20h70 are described, respectively. These isotherms confirm the results shown in the previous sections of the study and also illustrate the heterogeneity of the temperature distribution in the cross-section. When increasing exposure time to the ISO 834 temperature, the thermal field and thus the average temperature in the concrete and in the reinforcement also improves. These temperatures cause mechanical damage to the materials (i.e., concrete and steel or GFRP reinforcements), which can lead to failure of the beam and any RC structure.

3.5 Ultimate condition in fire

Figure 8 shows the FRR for each beam proposed in Table 1 when the critical concrete temperature (350°C) is reached. The concrete above and below the neutral axis, which is required to define the sagging and the hogging bending resistant moment in case of fire with the analytical expressions, respectively, has been considered. As expected, the compressive concrete below the neutral axis, which is required to determine the hogging moment, is most damaged by the fire. The concrete below the neutral axis develops a lower average temperature, which is partly due to the adjacent slab in this area. Comparing Fig. 8 with Figs. 9 and 10, it is clear that the temperature of the reinforcement is crucial to justify the failure of the beams exposed to fire.

Figure 9 and Fig. 10 show the FRR time assuming the critical temperature in the positive and negative reinforcement for conventional steel and GFRP rebars, respectively. Regarding negative reinforcement, the beams with conventional steel rebars have an FRR > 180 min. This time is not exactly determined because the numerical analysis ends at 180 min (as standardized building requirements often do not require FRR more than 180 min, which justify the authors criteria).

Assuming the positive reinforcement temperatures shown in Fig. 9, which represents the critical case of the fire-exposed beams, the FRR of the conventional steel RC beams was 401.1, 419.5, 375.0, and 348.4% higher than that of the identical GFRP RC beams at C = 20, 30, 40, and 50 mm. It is to be expected that the GFRP-RC beams will fail quickly in fire, which shows a major inconsistency in the technical application of these rebars. Of course, there is great interest in these reinforcements due to the aggressiveness of the environment (especially the chemical environment), but since the building must also meet a variety of performance requirements, the use of these rebars must be reconsidered in some cases.

The main conclusions of this research were:

The width is the main factor influencing the average temperature of the concrete in the cross-section of the beam;
GFRP RC beams with a concrete cover thickness of 20, 30 and 40 mm do not reach the 20 min of the ISO 834 fire. It is to be expected that these beams will fail very quickly when subjected to high temperatures;
In terms of fire resistance time, concrete beams with conventional steel reinforcement may have a fire performance - in some cases - four times higher than identical reinforced concrete structures with GFRP rebars;
Taking into account typical cross-sectional dimensions of the RC beams, the negative GFRP reinforcement was damaged by the fire, while identical beams with steel reinforcement were not damaged;
The study shows a contradiction: the lower concrete cover thickness of GFRP rebars compared to traditional steel rebars is advantageous in terms of environmental aggressiveness, but undesirable in terms of fire safety criteria. The susceptibility of GRP rebars to fire may help to explain this;
In certain circumstances, the use of GFRP RC structures instead of the usual steel RC structures must be investigated and discussed, especially if the building has to meet fire safety requirements;
As future research, the authors suggest considering the fire sensitivity of these reinforcements and finding a technical solution to improve the thermal and mechanical properties of the GFRP rebars at high temperatures.

Author Contribution

FLB: conceptualization, data curation, formal analysis, numerical analysis, writing original text; DBS: conceptualization, formal analysis; ECP: conceptualization, formal analysis.

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The fire performance of GFRP-RC beams based on FE thermal analysis

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Abstract

Figures

1 INTRODUCTION

2 METHODS

2.1 Research overview

2.2 Cross-sectional thermal analyses

2.3 Numerical procedures

2.4 Reinforcements assumptions

2.5 Validation of the FEA model and numerical procedure

3 RESULTS AND DISCUSSIONS

3.1 Concrete

3.2 Positive reinforcements

3.3 Negative reinforcements

3.4 Temperature field in the cross-section

3.5 Ultimate condition in fire

4 CONCLUSIONS

Declarations

Author Contribution

References

Additional Declarations

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