Flexural Performance of CFRP-Strengthened RC Beams with Single and Hybrid Anchorage Systems

Reinforced concrete (RC) structures frequently require rehabilitation, retrofitting, or strengthening due to material aging, functional obsolescence, or changes in design requirements. Due to the high strength-to-weight ratio and durability characteristics of carbon fiber reinforced polymer (CFRP), such materials are currently preferred for such applications [1]. CFRP has recently emerged as a primary solution in structural strengthening applications compared to the traditional counterpart materials [2-6]. However, the problem of CFRP delamination from the substrate surface in regions of high shear and flexural stresses is the main limitation of utilizing the full capacity of CFRP strengthening [7-9]. Without proper incorporation of some confinement or anchorage systems, CFRP materials can only utilize 10–25% [7], or at best estimate only 50% of its strength [10].

In an effort to utilize the full tensile capacity of CFRP material by increasing the bonding strength, several confinement and anchoring techniques have been proposed and individually verified during the last few decades. As shown in Figure 1, carbon fiber jacketing (also known as U-Wrapping) and fiber-reinforced polymer (FRP) near-surface mounted (NSM) are among the most recognized solutions [2, 5, 8, 11-16]. It has been reported that such confinement systems can significantly improve the strength, ductility, stiffness, and overall performance of strengthened elements [6, 9, 11, 17]. However, each technique has its own drawbacks and limitations. For example, by implementing U-wrapping alone, debonding at the concrete-FRP interface near the termination point of the “U-anchors” is frequently reported as the dominant failure mode, unless mechanically anchored [18, 19]. NSM, on the other hand, requires deep grooves (typically in Figure 1, h_g = 15–30 mm deep [14]) to completely submerge a rigid FRP bar or strip. If the existing concrete cover is thin, this technique is practically disqualified because structural stirrups or tension steel might be cut due to the grooving process.

During the past few years, some researchers proposed geometric modification of the substrate surface by introducing shallow surface grooves to increase the bond area between CFRP and the substrate concrete, as shown in Figure 2 [20-24]. Introducing externally bonded fiber-reinforced polymer (EB-FRP) on/in grooves (externally bonded reinforcement on grooves (EBROG)/externally bonded reinforcement in grooves (EBRIG)) has resulted in a significant improvement in the flexural performance of strengthened RC elements due to the enhancement in the bond strength between the strengthening and substrate materials [21, 25, 26]. While grooving configurations are typically shallow, the technique becomes less viable for extensive surface applications or degraded RC elements. In such cases, the high labor intensity and prolonged installation time considerably reduce the efficacy of grooving techniques relative to alternative retrofitting strategies. In addition, it has been reported that such surface modifications do not completely prevent debonding, but can only postpone such failure due to better utilization of FRP tensile strength [26-28].

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It has recently been reported in some relevant publications that carbon fiber spike anchors (CFSA) can be a vital solution to the CFRP delamination problem by providing continuity to the load path and therefore directly transferring stresses to the substrate material [29-33]. The interfacial compatibility between the anchoring and strengthening materials—both being carbon fiber—optimizes the performance of the anchorage through material homogeneity. Furthermore, the implementation of CFSA is non-invasive, preserving the original dimensions, weight, and structural integrity of the strengthened structural elements. Beyond these structural benefits, the installation of CFSA is significantly more efficient than traditional alternatives, offering a rapid and less labor-intensive application.

While significant progress has been made towards understanding FRP-to-concrete bonding, the most recent version of the ACI PRC-440.2-23 design code [34] highlights a persistent need for “additional evidence to validate performance” in several key areas. A primary recommendation from the code involves the exploration of various anchoring configurations as a means to enhance structural integrity. This study responds to that mandate by addressing the ongoing challenges of debonding through the evaluation of hybrid anchorage strategies, an area that the ACI PRC-440.2-23 [34] code explicitly identified as requiring more robust experimental data. Moving beyond the study of individual anchorage methods that have been extensively investigated, this work investigates the combined application of spike anchors and surface grooving. This approach seeks to determine how and to what extent such a hybrid configuration enhances the structural performance indices such as flexural strength, ductility, stiffness, and energy absorption (toughness) of EB-CFRP-strengthened beams compared to traditional single-method benchmarks.

3.1 Critical loads and corresponding deflections

The cracking, flexural steel yielding, and ultimate moments during the test, and the corresponding deflections have been carefully recorded. Table 7 shows the recorded loads and deflections of all tested specimens. Depending on the data recorded, the percentage increase/decrease in the recorded critical loads and midspan deflections has been determined as compared to the CB and to the EBR beam as listed in Tables 8 and 9, respectively. In both tables, the positive sign indicates an increase, and the negative sign indicates a decrease in the calculated index.

Examining Tables 8 and 9, it can be concluded that either of the strengthening systems employed has resulted in a marginal positive effect regarding the load-carrying capacity at all critical loading stages.

Table 7. Critical loads and midspan deflections

Table 8. Increase/decrease of critical loads and midspan deflections compared to CB

Table 9. Increase/decrease of critical loads and midspan deflections compared to EBR

Traditional CFRP strengthening, as in specimen EBR, has resulted in a noticeable increase in the cracking, yielding, and ultimate loads by more than 45, 25, and 23%, respectively, compared to the CB specimen. However, such an increment in the ultimate load-carrying capacity was at the expense of midspan deflection, where the corresponding ultimate deflection has been decreased by 33%, which will ultimately result in a reduced ductility of EBR compared to CB, as will be discussed later.

Specimens FG and SA, where either full grooving or spike anchors, respectively, were incorporated, led to an increase in the critical loads recorded at all loading stages as compared to CB and EBR specimens. For specimen FG, the critical loads, in sequence, have been increased by 62, 76, and 83% compared to CB, and by 11, 41, and 49% compared to the EBR beam. Similarly, specimen SA showed an increase in the load-carrying capacity during the critical loading stages by 52, 77, and 82% compared to CB, and by 5, 41, and 48% compared to the EBR specimen. Both beams, especially the one with FG, have provided a comparable ultimate deflection compared to the CB, where the ultimate deflections have been reduced by about 5.5 and 16.5% for specimens FG and SA, respectively. This probably will result in an enhancement in the beam's ductility compared to the EBR beam, where only externally bonded CFRP has been incorporated.

With respect to the FGSA specimen, which, in addition to the EBR strengthening, has been additionally anchored by means of full-length grooving and an SA at each CFRP layer termination point, it has shown superior behavior and therefore loading capacity compared to either of the specimens discussed above. For example, the ultimate load of this beam has been increased by more than 94 and 57% compared to CB and EBR, respectively. In addition, the difference in the ultimate deflection of this beam specimen was only 1.29% smaller than CB and about 48% larger than the ultimate deflection in comparison with the EBR specimen.

The last three beams that were tested to evaluate the interaction of grooving length and spike anchor quantity with the presence of intermediate U-wrapping also exhibited superior performance compared to the CB and EBR specimens, as shown in Tables 8 and 9. Specifically, the specimen featuring double spike anchors with 600 mm grooving (G6DAU) and the specimen with a SA and 800 mm grooving (G8SAU) achieved yielding and ultimate loads comparable to those of the FGSA. In addition, these specimens nearly doubled the capacity of the CB. This demonstrates that intermediate U-wrapping can successfully compensate for a 40% and 20% reduction in grooving length when utilizing double and single anchors, respectively. The G6SAU specimen likewise demonstrated considerable load-carrying capacity at all three loading stages. Nonetheless, its overall results were marginally lower than the other two specimens in this category. This difference was particularly pronounced in the cracking load, which increased by only 62% compared to the 100% improvement achieved by both G6DAU and G8SAU.

3.2 Modes of failure

Failure of RC beams is not a sudden event at a single load point; it is a gradual process that begins with cracking and evolves through several stages until ultimate collapse. Therefore, failure mode can be considered a critical flexural performance aspect, particularly when CFRP strengthening is employed. CFRP debonding is considered a highly unfavorable failure mechanism because it occurs abruptly and without warning once the ultimate strain capacity of either the CFRP-epoxy or epoxy-concrete interface is reached [4, 6, 37]. The modes of failure of all tested beams, as shown in Figure 8, have been carefully monitored and discussed.

The CB exhibited flexural failure as flexural cracks initiated once the cracking load reached 21 kN. As the applied load increased, these cracks propagated extensively and penetrated deeper into the section. Ultimately, the beam failed when the applied load and corresponding deflection reached 56 kN and 16.34 mm, respectively.

Similarly, the failure process of the EBR specimen also started once the cracking load reached 30.5 kN (about 45% of CB). When the applied load slightly increased, some shear cracks were also initiated. Flexural and shear cracks penetrated deeper into the section, but at a lower rate compared to CB, as the applied load increased up to the sudden failure by end-CFRP debonding, as shown in Figure 8, which is encountered at the ultimate load and corresponding deflection of 69 kN and 10.92 mm, respectively.

The other six beams tested displayed the same failure process as EBR, where flexural and some shear cracks initiated once the applied load reached a specific value ranging from 32 to 42 kN. These cracks have been widening and penetrating deeper into the section, and some other new cracks have been initiated as the applied load increased. Ultimately, either of the beam specimens in this category has failed abruptly once the ultimate load has reached a specific value depending on the strengthening and anchoring configuration employed, as shown in Table 7.

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Figure 8. Mode of failure of the tested beams

Except for the FG specimen, which failed due to concrete cover separation, the five specimens reinforced with additional spike anchors exhibited a shift in failure mode from CFRP debonding to CFRP rupture followed by flexural failure. As illustrated in Figure 8, this transition to a more desirable failure mechanism, along with an increased load-carrying capacity, confirms the effectiveness of the spike anchoring technique. From the current results, it can be concluded that the anchored specimens sustained the CFRP contribution to exhibit delayed debonding and higher load levels in comparison with the control specimen. This behavior, combined with the enhanced load-deflection relationship, provides additional evidence of more effective stress transfer provided by the spike anchors.

3.3 Load-deflection responses

The individual deflection response of the tested beams as the applied load is increased has been recorded using a high-accuracy midspan dial gauge. Then, the load-deflection relationship has been constructed, discussed, and compared against each other in some main categories.

The first category, as shown in Figure 9, contains the CB, the beam with EBR, and the beams with single anchor techniques FG and SA. As shown in Figure 9, the CB specimen showed the expected elastic and inelastic behavior through the loading stages. Once the ultimate load of 56 kN was reached, the recorded load started to drop gradually, and the deflection continued to increase until failure of the specimen with an ultimate deflection of 16.43 mm as stated previously. In contrast, all other three beams in this category have shown only the elastic part, and all have failed abruptly, but with higher loads and lower deflections as compared to CB. Examining Figure 9, it is clear that incorporating either surface grooving or spike anchors has resulted in an important enhancement in the flexural capacity of the anchored beams compared to the beam without any sort of anchoring (i.e., EBR).

In the second category, the load-deflection response of the specimen incorporating a hybrid anchoring technique of full grooving and a SA at each of the CFRP layer termination points combined (FGSA) will be compared against the specimens that incorporated a single anchoring technique of either FG or a SA. The purpose of this category is to determine whether combining the two proposed anchoring techniques will result in any advancement in the flexural performance of the strengthened beams. The load-deflection curves of this category are shown in Figure 10. It is evident that specimen FGSA has resulted in a slightly better performance compared to either FG or SA that contains a single anchoring technique since it provided a larger ultimate load and deflection.

Examining Figures 9 and 10 as well, one can clearly see that the performance of the FG specimen was better than that of specimen SA in terms of the ultimate deflection, as it sustained roughly the same ultimate load of 102 kN but with about 2 mm larger deflection. This will probably affect the ductility of the specimen with grooving as compared to the one with a spike anchor, as will be discussed later.

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In the last category, several comparisons will be made. The first one will be made between G6SAU and G6DAU to isolate and examine the effect of the number of spike anchors employed. The load-deflection responses of both specimens, in addition to the EBR specimen, are shown in Figure 11. As can be concluded from this figure, both specimens greatly outperformed the EBR specimen in terms of ultimate load capacity and the corresponding deflections. In addition, these P- $\Delta$ graphs show that G6DAU has sustained a larger ultimate load and deflection compared to specimen G6SAU. This figure emphasizes the effect of the number of spike anchors in strengthening applications.

Similarly, another comparison will be made between specimens G6SAU and G8SAU to determine the effect of the length of grooving in such applications. P- $\Delta$ graphs of both specimens, in addition to EBR, are shown in Figure 12. This figure, as well, emphasizes the role of grooving length in sustaining more load where specimen G8SAU has sustained about 14 kN larger force compared to G6SAU. As will be discussed later, the stiffness of the specimen with 800 mm grooving will probably be larger than the specimen with 600 mm grooving.

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Another comparison can be performed between G6DAU and G8SAU to determine which anchoring configuration can provide better flexural performance. The load-deflection responses of both specimens can be seen in Figure 13. It can be seen that both specimens have provided approximately the same ultimate loads. However, specimen G8SAU will indeed result in a larger stiffness, as the ultimate deflection is lower than that of G6DAU.

Finally, all the specimens in this category, namely G6SAU, G6DAU and G8SAU, can be compared against the FGSA specimen to determine the effectiveness of the incorporated intermediate u-wrapping and whether such wrapping can effectively compensate for the reduced length of grooving. The results of this comparison can be found in Figure 14. The results clearly suggest that full grooving with SA as in specimen FGSA has provided a comparable performance compared to the specimens incorporated intermediate u-wrapping in terms of the ultimate load sustained, but with possible lower stiffness, as will be discussed in the next section.

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3.4 Ductility and stiffness

It is noteworthy that the strengthening effectiveness of RC elements is often achieved at the expense of ductility, which defines the capacity of the element to deflect inelastically up to failure without a significant loss of strength [38, 39]. Conversely, the stiffness of the RC elements can be greatly boosted by CFRP strengthening of such elements [40]. Such indices, in fact, can be considered among the most critical performance indicators that must be assessed for a safe strengthening process. Since these two indices are inversely related, they will be calculated for all beams tested and compared against each other in this research.

The stiffness (k) has been calculated by considering the elastic part of the load-deflection curve. Eq. (1), which is equivalent to the slope of the elastic part of the load-deflection curve, can be used to calculate the bending stiffness of the tested beam:

$k=\frac{P_y-P_{c r}}{\Delta_y-\Delta_{c r}}$              (1)

On the other hand, since the behavior of strengthened beams is not perfectly elasto-plastic, as such beams employ concrete, steel and carbon fiber at the tension side of the cross-section, ductility $\left(\mu_{\Delta}\right)$ has been calculated based on the total and elastic absorbed energy [38]. This definition was first suggested by Namaan and Jeong [41]. This index can be calculated based on Eq. (2). For perfectly elasto-plastic behavior, this definition will result in exactly the same value as the conventional ductility definition proposed by Park [42] that uses the ultimate and yield deflections.

$\mu_{\Delta}=\frac{1}{2}\left(\frac{E_{\text {Total }}}{E_{\text {Flastic }}}+1\right)$            (2)

where:

E_Total is the area under the load-deflection curve up to the ultimate load, and

E_Elastic is the area under the elastic part (i.e., up to the yield load) of the load-deflection curve.

Table 10. Stiffness and ductility indices of the tested beams

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Figure 15. Percentage increase or decrease of the stiffness and ductility indices of tested beams as compared to the control beam (CB) specimen

The inputs used to calculate the stiffness and ductility indices are listed in Table 7, whereas Table 10 presents the final results of both indices. It should be noted here that E_Total listed in the second column of Table 10 represents the toughness index (in units of kN·mm) used frequently as a performance index in some publications [43]. Figures 15 and 16, on the other hand, illustrate the percentage of increase or decrease in such indices compared to CB and EBR, respectively. Examining Table 8 and Figure 15, it can be concluded that, in general, the gain in the flexural capacity of the strengthened and anchored specimens, and therefore the stiffness, has been at the expense of ductility. Conversely, by comparing the ductility and the stiffness of all tested beams to the EBR specimen, it can be concluded that the anchoring techniques incorporated have restored part of the lost ductility due to the strengthening.

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Figure 16. Percentage increase or decrease of the stiffness and ductility indices of tested beams as compared to the externally bonded reinforcement (EBR) specimen