Experimental Study of the Interfacial Shear Behavior Between NRC and UHPC in UHPC-Jacketing Rehabilitation of Concrete Bridges

Abstract

Ultra-High-Performance Concrete (UHPC) jacketing is an effective and innovative strengthening method in the renovation projects of concrete bridges. In December 2021, the UHPC-jacketing method was first applied to rehabilitate a seriously damaged bridge in the Changzhou Bridge rehabilitation project in Guangzhou, China. However, the interfacial shear behavior between the Normal Reinforced Concrete (NRC) substrate and UHPC is a crucial factor for the effectiveness of the UHPC-jacketing strengthening method. Therefore, four push-out specimens were designed in this paper to investigate the effects of the embedded bolt diameter (12 mm and 16 mm) and construction method (cast-in-place UHPC layer (ZJ group) and precast UHPC panels with infilled high-strength mortar (GJ group)) on the shear behavior of the NRC–UHPC interface. The results indicated that with the increased bolt diameter from 12 mm to 16 mm, the first peak load (P₁) rose from 920.17 kN to 1048.07 kN (+13.9%) in the ZJ group and from 838.08 kN to 1204.20 kN (+43.7%) in the GJ group. The residual loads (P_r) of the GJ group were smaller than those of the ZJ group, at 41.9% and 30.2% lower for bolt diameters of 12 mm and 16 mm, respectively. The construction method of high-strength mortar filling was significantly influenced by the bolt diameter, with a diameter of 16 mm required to fully utilize its shear resistance. Predictions from ACI 318-19 underestimated experimental shear capacity by 70.6% on average, while AASHTO (2017) and Fib provided accurate estimations (within 9.8–10.9% of experimental values).

1. Introduction

With urbanization and economic development, road traffic pressure intensifies alongside natural aging and man-made factors, leading to varying degrees of deterioration in concrete bridges [1,2,3,4]. Many aging bridges experience reduced load-bearing capacity due to issues such as concrete cracking, rebar corrosion, and concrete spalling. Employing appropriate methods to strengthen and rehabilitate these aging bridges can not only avoid the inconveniences associated with bridge demolition but also yield significant social and economic benefits through the implementation of low-carbon practices.

Existing methods for strengthening reinforced concrete structures include external prestressing, fiber-reinforced polymer (FRP) wrapping, steel plate bonding, and section enlargement strengthening. Extensive research has been conducted on the advantages and disadvantages of these strengthening methods [5,6,7,8,9]. In the section enlargement method, when normal concrete (NC) was used as the strengthening material, the increased concrete layer thickness and additional reinforcement would impose extra loads on the bridge and compromise the vertical clearance beneath the bridge surface.

Ultra-High-Performance Concrete (UHPC) is an excellent cementitious material that has been widely used in various engineering structures in recent years [10,11,12,13]. This is because UHPC has advantages in higher compressive strength, tensile and flexural strength, and perfect ductility and durability as compared to other normal concretes [14,15,16]. As shown in Figure 1, the use of UHPC-jacketing to rehabilitate Normal Reinforced Concrete (NRC) has become a hot topic of research, due to higher self-weight and structural efficiency [5,6,17,18,19,20]. Moreover, there are several existing configurations for UHPC-jacketing techniques [17,21], such as bottom, lateral, and U-jacketing, all of which can significantly increase the load-bearing capacity of concrete beams, especially the U-jacketing type, which is the focus of this study.

A serious fire accident happened on the Changzhou Bridge on the Xinhua Expressway in Guangzhou, China (as shown in Figure 2). High-temperature flames mainly affected the girders, bent cap, and piers, which caused serious damage to the bridge’s localization. The rehabilitation project employed UHPC-jacketing of the section enlargement technique to strengthen localized seriously damaged regions, in order to restore the load-bearing capacity of the girder. In particular, the UHPC-jacketing was constructed in cast-in-place, and the thickness of the UHPC layer was initially designed to be 50.0 mm, while the embedment depth of the bolts in the UHPC layer was initially designed to be 30.0 m. The rehabilitation works were completed in December 2021. This represents China’s first application of UHPC-jacketing to rehabilitate a seriously damaged concrete bridge.

However, the key to the effectiveness of the UHPC-jacketing is the interfacial properties of the NRC substrate and the UHPC layer. Many scholars have investigated the interfacial bonding properties of NRC and UHPC in recent years [22,23,24,25,26,27]. Harris and Carbonell et al. [22,23] investigated the effect of concrete substrate roughness and wetness, curing conditions, and other factors on the interfacial bond between UHPC and normal-strength concrete (NSC) by means of splitting tensile tests, pullout tests, and oblique shear tests. The results showed that the interfacial bonding between UHPC and NSC was excellent, and the early strength of UHPC has a large influence on the interfacial bond strength. According to diagonal shear tests by Hussein et al. [24], they investigated the bonding and friction at the UHPC–NRC interface with different roughness levels and obtained the tensile strength and friction coefficient of the bonded interface. Shuo et al. [25] reported that the bonding of UHPC to concrete substrates of different strengths was superior to normal concrete. He et al. [28] investigated the bond strength between UHPC and normal concrete (NC) through tensile tests, and the results indicated that the bond strength of UHPC to NC was approximately 2.0 MPa. Weng et al. [26] applied 3D laser scanning to quantify UHPC–NC interfacial roughness via fractal dimensions in diagonal shear tests, establishing a correlation between bond strength and slip for retarded scoured interfaces. In real engineering, bridges endure cyclic fatigue loading that may accelerate the interface between the UHPC layer and the NRC degradation, which may affect the durability of the UHPC-jacketing and compromise long-term safety [29,30,31].

In addition to the experimental methods, many scholars have also used numerical simulation methods to predict the properties of the NRC–UHPC interface [32,33,34,35]. It should be noticed that the UHPC studied above was poured as a repair material on the NRC substrate. The results indicated that the strengthening effect of the UHPC layer depended on the strength of the concrete substrate when no grooves and interfacial rebars were provided, and the roughness of the substrate had a significant effect on the strengthening effect.

The properties of the NRC–UHPC interface can be enhanced by embedding the steel bars in the NRC substrate [36,37]. Gao et al. [36] investigated reinforced concrete (RC) beams strengthened with thin layers of UHPC and embedded steel rebars in the RC substrate. The results indicated that the embedded steel rebar could effectively improve the cracking resistance of the interface. At the same time, bolts were widely used as high-strength connectors in a variety of structures [37,38,39,40]. Jiang et al. [41] conducted single shear tests on “Z”-shaped specimens with different interface treatments and grooves on the NRC interface and proposed formulas to calculate the direct shear strength of the NRC–UHPC interface.

However, though many scholars have studied the NRC–UHPC interface, the effect of the shrinkage of UHPC materials has not been taken into account. Shrinkage of UHPC is difficult to control in real engineering, and for UHPC-jacketing, not only cast-in-place UHPC but also prefabricated UHPC panels can be used. Therefore, prefabricated UHPC panels can be used as an alternative and connected to the existing T-beams by completing them with high-strength mortar. Therefore, for practical UHPC-jacketing applications, factors such as UHPC shrinkage, cover thickness, the interface reinforcement ratio, anchor detailing, etc., need to be considered. Specific design parameters should be determined case-by-case based on the actual bridge conditions.

While previous studies have investigated the NRC–UHPC interface through tensile tests and diagonal shear tests [22,23,24,25,26,33], there are few references to the study of interfacial properties through push-out tests [7,8,18], especially the interfacial shear behavior between NRC and precast UHPC panels with infilled high-strength mortar. In order to fully reveal the interfacial shear behavior of NRC–UHPC and further enrich the database of related studies, a more in-depth study is necessary.

In summary, based on the design and requirements of the Changzhou Bridge rehabilitation works, this study designed two groups of specimens: (1) the ZJ group, which consists of specimens with a cast-in-place UHPC strengthening layer around the NRC beams, and (2) the GJ group, which involves prefabricated UHPC with high-strength mortar filled between the UHPC and the NRC substrate. Through these two groups of specimens, the influence of the bolts’ diameter and different casting methods on the shear capacity of the UHPC strengthened NRC substrate and the structural ductility at interface failure was investigated. The interfacial shear strength was predicted by mainstream formulae. The study aims to identify key factors affecting the interfacial connection performance and validate the applicability of existing formulas. The findings provided a reference for UHPC-jacketing applications, enabling designers to balance structural performance against cost efficiency within project constraints. In addition, after clarifying the interfacial shear behavior of NRC–UHPC, the authors will conduct further research on the shear resistance of UHPC-jacketing composite beams.

2. Experimental Investigations

2.1. Specimen Design

In this paper, the push-out tests were designed to capture primary failure modes and load-slip behavior, quantify the effects of bolt diameter and construction method, and provide baseline data for subsequent composite-beam tests.

In real UHPC-jacketing engineering, the NRC substrate was covered by a UHPC layer, so in this paper, we adopted a double shear push-out specimen, where the UHPC was on the outer side. Moreover, the connecting bolts were embedded in the NRC substrate and crossed the interface of NRC and UHPC. Connecting bolts with diameters ranging from 12 and 16 mm were employed as shear connectors at the interface. The design of the shear connectors was based on real engineering cases where 12 mm diameter ribbed steel bars were used for through-thickness reinforcement of an existing beam in the rehabilitation project by the section-enlarging method. This experiment was designed with a total of two groups of specimens, with three specimens in each group. The first group constitutes UHPC poured directly over the NRC substrate, and the second group consists of pouring high-strength mortar between precast UHPC panels and the NRC (see Figure 3); the parameters of the specimens are detailed in Table 1. In every group, the parameter was the bolt diameter, which was set as 12 mm or 16 mm.

The naming convention of the specimens follows the format of “construction method + bolt diameter”. For example, “ZJ-12” denotes a specimen reinforced with 12 mm diameter bolts and cast with a UHPC layer, whereas “GJ-12” refers to a specimen reinforced with 12 mm diameter bolts, strengthened with a prefabricated UHPC layer, and filled with high-strength mortar between the UHPC and NRC substrate. Additionally, HRB 400 steel bars with diameters of 12 mm and 25 mm were used as stirrups and the main reinforcement for the NRC substrate, respectively (see Figure 3).

2.2. Specimen Fabricating

The specimen fabrication process is illustrated in Figure 4. Initially, 4 specimens of the normal concrete structural portion were cast and cured, and their surfaces were roughened at an age of 3 days. Subsequently, the UHPC layer for the ZJ group specimens and the prefabricated UHPC layer for the GJ group specimens were cast. The specimens in the ZJ group and prefabricated UHPC panels in the GJ group were demolded the following day and cured using hot water at a temperature exceeding 80 °C. For the specimens in the GJ group, when the UHPC reached an age of 3 days, the concrete block and two side UHPC panels were assembled, while the high-strength mortar was installed in the gap between NRC and UHPC. The specimens were then cured until the high-strength mortar reached an age of 28 days. The GJ group was tightened with nuts and washers. Additionally, to ensure sufficient anchorage strength, the nuts and washers were fixed to the ends of the bolts prior to casting the UHPC layer for the ZJ group specimens (Figure 4b).

2.3. Material Properties

The mix proportions of the three relevant materials are provided in Table 2. The UHPC and high-strength mortar utilized P·II 52.5R cement, while the C50 concrete used P·II 42.5R cement. Hooked-end steel fibers with a diameter of 0.22 mm, a length of 13 mm, and a tensile strength of 2600 MPa were employed. The high-strength mortar was optimized based on the UHPC mix proportion by adjusting the fine aggregate ratio. The fine aggregate consisted of 30% quartz sand with particle sizes between 0.6 mm and 0.3 mm and 70% quartz sand with particle sizes less than 0.3 mm. In order to ensure the quality of concrete materials, the concrete materials in this paper meet the requirements of the GB/T 17671-2021 standard [42].

On the day of specimen loading (when the high-strength mortar reached an age of 28 days), mechanical property tests were conducted on the three types of concrete used in the specimens in accordance with relevant standards [43,44,45]. For NRC, the compressive strength was determined by a 150 mm × 150 mm × 150 mm cube, and the splitting tensile strength and elastic modulus were measured by a φ100 mm × 200 mm cylinder. The compressive strength of the UHPC and high-strength mortar was tested on 100 mm × 100 mm × 100 mm cubes, whereas the splitting tensile strength and elastic modulus were tested on φ100 mm × 200 mm cylinders. The results are presented in Table 3. According to the manufacturer’s report, the HRB 400 steel bars had an elastic modulus of 198 GPa, a yield strength of 460.85 MPa, and an ultimate tensile strength of 579.44 MPa. The bolts had an elastic modulus of 202 GPa, a yield strength of 680.48 MPa, and an ultimate tensile strength of 732.33 MPa

2.4. Experimental Setup and Loading Protocol

The loading setup is illustrated in Figure 5. During the test, four displacement transducers with a range of 50 mm were positioned at the mid-span of the strengthening interface to measure the relative slip at the interface. Both the load and displacement readings were recorded by a data acquisition system throughout the test. At the beginning of the test, a preload of 50 kN was applied to the specimen, followed by unloading to 0 kN. Subsequently, formal step-by-step loading commenced, with each step increasing by 20 kN until the specimen failed.

3. Experimental Results and Discussions

The first peak load (P₁), the corresponding interfacial shear strength at the first peak, the second peak load (P₂), and the slips (d₁ and d₂) corresponding to the first and second peaks are summarized in Table 4. We took the load at a slip of 2 mm as the residual load (P_r). In particular, the S side was taken for the ZJ-16 specimen and the N side for the GJ-16 specimen. The interfacial shear stress (τ₁) at the peak was calculated using the formula τ = P/(2ab), where a is the length of the interface (a = 600 mm) b is the width of the interface (b = 350 mm).

3.1. Failure Modes

A constant loading rate was kept throughout the testing process. As shown in Figure 6, significant differences in the failure modes between the ZJ group and GJ group were observed during loading. It should be noted that the details in Figure 6b were captured in the NRC surface, whereas Figure 6d was captured in the UHPC surface. For the ZJ group, no obvious cracks were observed on either NRC or UHPC, which indicated that shear force transfer in cast-in-place construction relied largely on the bonding effect at the interface. The bonding force was mainly determined by the strong adhesion of UHPC and the bonding area [28]. However, for the GJ group specimens, the failure surface near the loading point occurred at the interface between the mortar layer and the NRC substrate, while at the bolt positions far away from the loading point, the shear transfer path deviated, and the failure surface shifted to the interface between the mortar layer and the UHPC layer. Specifically, as shown in Figure 6c, the high-strength mortar layer in specimen GJ-12 exhibited diagonal cracks traversing both the NRC and UHPC panels, at an angle of approximately 20° to the UHPC panel, which was typical of tensile shear failure. Since the load on the NRC was transmitted to the UHPC panel through the mortar layer and bolts, the contribution of the bolts at the location of the diagonal cracks was minimal, which led to the load transmitted by the mortar layer exceeding its ultimate limitation.

3.2. Load-Slip Curves

Figure 7 presents the load-slip curves for specimens. In order to fully reveal the failure process of the specimens, the load–slip curves of the specimens were distinguished by the S side and the N side (see Figure 5 for the positions of S and N), and we moved the curves of the N side 4 mm to the right. Based on Figure 7 and the observations during the loading process, it can be concluded that both groups of specimens underwent the following sequence: load reaching the first peak (P₁) and resulting in cracks and failure at one interface; experiencing a sudden drop in load-bearing capacity; load re-climbing to the second peak (P₂) and resulting in cracks and failure at the other interface; another decrease in load-bearing capacity; and complete failure at both interfaces. The occurrence of two peaks is attributed to the fact that the NRC–UHPC interfaces on both sides of the specimens did not fail simultaneously during loading. Instead, one interface failed first as the load increased, reaching the first peak. As loading continued, the other interface also failed, leading to the second peak in load.

The load–slip curves of the ZJ group specimens exhibited highly similar trends. Taking specimen ZJ-16 as an example, the load initially experienced a linear stage before reaching the second peak, with the load at the end of the linear stage being approximately 0.65P₁. This was followed by a nonlinear stage and a descending stage. For specimens ZJ-12 and ZJ-16, the ratios of the second peak load (P₂) to the first peak load (P₁) were 0.93 and 0.99, respectively, with corresponding slip values of 0.99 mm and 3.36 mm, respectively. This indicates that, compared to the first peak, the increase in bolt diameter enhanced both the second peak load and the corresponding ductility of the ZJ group specimens.

For the GJ group specimens, specimen GJ-12 exhibited no change in slip after reaching the first peak. When the load reached the second peak, both peak loads and slips were identical, after which the curve entered the descending stage. This was because the load was transmitted to the UHPC panels through the mortar layer. After the mortar layer cracked (see Figure 6c), the bearing capacity rapidly decreased, the interfacial slip suddenly increased, and the bolts almost carried all the load at that moment. In contrast, specimen GJ-16 showed a change in the curve slope after the load reached 1000 kN due to cracking in the high-strength mortar of the grouting layer before reaching the first peak. Unlike the other specimens, the second peak load of GJ-16 was only 0.59P₁. This may be attributed to the larger load borne by this specimen, causing slight damage to the other interface, even though no visible cracks were observed when the first peak load was reached.

3.3. The Influence of Bolt Diameter

The comparison of the first peak load (P₁) and residual load (P_r) between the ZJ group specimens and the GJ group specimens is shown in Figure 8.

From Figure 8 and Table 4, it can be observed that the load-carrying capacity of the ZJ and GJ groups increased by 13.9% and 43.7%, respectively, when using 16 mm bolts compared to 12 mm bolts. This indicates that increasing the bolt diameter enhances the shear capacity of the specimens.

After complete failure of both interfaces, all specimens retained some load-bearing capacity due to the presence of the bolts, which had not yet sheared off. This phenomenon occurred because both the interface and the bolts shared the load during the direct shear test. The connection between the NRC substrate and the strengthened UHPC layer via the bolts provided some residual resistance, allowing the strengthened specimens to continue bearing load even after complete interface failure. From Figure 8b, it can be observed that with the increase in the bolt diameter, the residual load decreased by 30.7% and 16.7% for the ZJ and GJ groups, respectively. It could be possible that the slip required for yielding has already occurred for the bolt with a diameter of 16 mm at P1 (see Figure 6a,b), resulting in a smaller contribution in the residual stage.

The above analysis demonstrated that the presence of bolts not only improved the load-bearing behavior of the specimens after interface failure but also increased the slip capacity at peak with larger bolt diameters. However, the larger the bolt diameter, the smaller the contribution of the residual load capacity at a later stage. Furthermore, it is worth noting that despite fixing nuts and washers to the bolt ends before the test to enhance the pull-out resistance of the bolts in the thin UHPC layer, the bolt head of ZJ-16 was found in a state between shearing and pull-out after the test (see Figure 6b). For specimen GJ-16, desquamation was observed in the high-strength mortar layer (see Figure 6d).

3.4. The Influence of the Construction Methods

As shown in Figure 9a, for specimens using 12 mm diameter bolts, specimen GJ-12, which was filled with high-strength mortar between the prefabricated UHPC layer and the NRC substrate, exhibited an 8.9% decrease in the first peak load (P₁) compared to specimen ZJ-12, which used a cast-in-place UHPC layer. At the point of failure of both interfaces, the interfacial slip of specimen GJ-12 was less than 0.1 mm for specimens using 16 mm diameter bolts. However, the first peak load (P₁) of specimen GJ-16 was 14.9% higher than that of specimen ZJ-16.

As shown in Figure 9b, for residual loads, the GJ group was lower than the ZJ group, where GJ was 41.9% lower than ZJ when the bolt diameter was 12 mm and 30.2% lower than ZJ when the bolt diameter was 16 mm. It should be noted that residual load is mainly composed of interlock and interface friction and requires the bolt to contribute to the pressure at the NRC and UHPC interface. However, higher shear forces may cause the bolt to be sheared off or to be unbonded, which may be the reason for the lower residual force of larger diameters.

The experimental results indicated that, compared to the cast-in-place UHPC layer, the method of prefabricating the UHPC layer and then filling high-strength mortar between it and the NRC substrate could effectively enhance the load-bearing capacity and stiffness of specimens when larger diameter bolts were used. Additionally, the use of 12 mm diameter bolts resulted in a negative effect with this method. This may be attributed to the fact that, after the peak load, the specimens with cast-in-place UHPC maintained a tight connection between the UHPC and the bolts, forming a constraint that kept the bolts in a triaxial stress state.

4. Comparison of Interface Shear Capacity Calculation Results from Existing Codes

4.1. Calculation Results of Interface Shear Capacity According to Different Codes

According to Section 22.9.4 “Nominal Shear Strength” of ACI 318-19 (2019) [46], based on the shear-friction theory, for cases where steel reinforcement or similar materials are used as shear connectors and are oriented perpendicular to the shear plane, the nominal shear strength of the interface should be calculated using Equation (1):

(1)$V_n=\mu A_{vf}{f}_y$

In the equation, μ is the friction coefficient, which is taken as 1.4 for cases of monolithic specimens, a high interface condition by sufficient surface roughening, and subsequent high-strength mortar or UHPC casting on the existing NC structure [41]. A_vf is the cross-sectional area of the shear connectors within the shear plane (mm²), while f_y is the yield strength of the shear connectors (MPa).

The formula provided in Section 5.7.4.3 “Interface Shear Resistance” of the AASHTO (2017) [47] specification takes into account the shear resistance contributed by both the concrete interface and the shear connectors through the shear-friction effect. The nominal shear strength of the interface should be calculated using Equation (2):

(2)$V_n=c{A}_{cv}+\mu (A_{vf}{f}_y+P_c)$

In the equation, c is the concrete interface coefficient, taken as 2.80 MPa for cases of monolithic specimens, and μ is the friction coefficient, taken as 1.4 [41]. Due to the fully chiseled NC substrate interface in this study, A_cv is the area of the shear plane (mm²). A_vf and f_y have the same definitions as in Equation (1), and P_c is the horizontal confinement force (kN). In this experiment, since the nuts for the GJ group specimens were only fixed to the surface of the UHPC, P_c is taken as 0.

Similar to the AASHTO (2017) [47] specification, the Fib Model Code for Concrete Structures (2013) [48] in Section 6.3 “Concrete to Concrete” also considers the shear resistance provided by the shear-friction theory. However, this specification additionally accounts for the dowel action of shear connectors. The nominal shear strength of the interface τ_u and the nominal shear force V_u should be calculated using Equations (3) and (4), respectively:

(3)${\tau }_u={\tau }_a+\mu \left({\sigma }_n{+K}_1\rho f_y\right)+K_2\rho \sqrt{f_y{f}_{cc}}$

(4)$V_u={\tau }_u{A}_c$

In the equations, τ_a is the shear strength provided by the bond on both sides of the interface. According to reference [49], τ_a = β × τ_c × A_c, where β = 0.454 for the NRC–UHPC interface and β = 0.483 for the interface of high-strength mortar and NRC. τ_c = 0.75 (f_cc1 × f_ct)^0.5, where f_cc1 is the axial compressive strength of C50 concrete cylinders and f_ct is the splitting tensile strength of C50 concrete cylinders (MPa). A_c is the area of the shear plane (mm²). μ is the friction coefficient, taken as 1.385 according to reference [49]. σ_n is the horizontal compressive stress at the interface (MPa), taken as 0 in this experiment. The coefficient K₁ = 0.5. ρ is the reinforcement ratio of the shear connectors in the shear plane (ρ = A_s/A_c, where A_c is the area of the shear connectors and A_c is the area of the shear plane (mm²)). According to reference [49], the coefficient K₂ = 0.9. f_y is the yield strength of the shear connectors (MPa) and f_cc is the axial compressive strength of the NRC structural concrete cylinders (MPa).

In addition, the requirements and limitations for the use of each formula are shown in Table 5.

4.2. Comparison of Formula Calculation Results

By substituting the experimental data into the above formulas, the theoretically calculated values of the shear capacity for each specimen in this test can be obtained. The ratio of the calculated values to the experimental values for shear capacity is shown in Figure 10, and the detailed contributions of each factor are shown in Table 6.

The average ratios of the calculated results to the experimental results are 0.294 for the ACI specifications, which underestimated the shear capacity by 70.6%. Due to the ACI formula focusing primarily on the contribution from the shear-friction effect of the shear connectors, the contribution of interface adhesion was ignored. However, as shown in Table 6, the calculation results of AASHTO (2017) and Fib specifications showed that the interface adhesion force V_cc contributed significantly to the overall result, which led to the ACI calculation results being too small. The average values for AASHTO (2017) and Fib were 0.89 and 0.90, respectively. It is worth noting that the average and standard deviation calculated by AASHTO (2017) and the Fib specifications were nearly the same. The AASHTO (2017) specification takes into account the bonding effect of the interface compared to the ACI specification. The Fib specification simultaneously considers the shear-friction effects of both the concrete interface and the shear connectors, as well as the dowel action of the shear connectors.

Based on the experimental values and the predictions from the specifications, it is recommended to use the AASHTO (2017) or Fib specifications for calculating the shear capacity of reinforced NRC structures strengthened with UHPC in design.

5. Conclusions

This paper investigates the interfacial shear performance of reinforced concrete structures strengthened with UHPC. By analyzing the experimental results of specimens with different bolt diameters and construction methods, and comparing them with existing code formulas, the following conclusions are drawn:

(1) The ZJ group showed significant brittle damage with almost no slip at the interface before reaching the first peak load. The slip at the first peak load for the GJ group was 0.09 mm and 0.16 mm, whereas the ZJ group was 0.02 mm and 0 mm. The construction method of filled high-strength mortar had better ductility. For ZJ-12 and ZJ-16, the ratio of the second peak load (P₂) to the first peak load (P₁) was 0.93 and 0.99. For GJ-12 and GJ-16, the ratio was 1.00 and 0.59, respectively.

(2) Increasing the bolt diameter to some extent enhanced the interface load-bearing capacity, and with an increased bolt diameter from 12 mm to 16 mm, the first peak load of ZJ and GJ groups increased by 13.9% and 43.7%, respectively; however, the residual loads decreased by 30.7% and 16.7%, respectively.

(3) Compared with the construction method, the residual load (P_r) of the GJ group was generally smaller than that of the ZJ group; when the bolt diameter was 12 mm and 16 mm, respectively, the residual load (P_r) of the GJ group was 41.9% and 30.2% smaller compared with that of the ZJ group.

(4) For the construction method of cast-in-place UHPC-jacketing, it is recommended to use bolts with diameters ranging from 12 mm to 16 mm. For the construction method of UHPC-jacketing, it is recommended to use bolts with a diameter of 16 mm.

(5) The average and standard deviation of the ratio of calculated-to-experimental values for the AASHTO (2017) formula were 0.89 and 0.06, respectively, while Fib was 0.90 and 0.08. It is recommended to use the AASHTO (2017) or Fib specifications for calculating the interfacial capacity between NRC substrates and the UHPC layer with bolts in design.

(6) In real-world engineering, concrete bridges can suffer different levels of damage (e.g., fire damage, acid corrosion, and fatigue damage), so it is recommended that future research consider the impact of the extent of damage to the concrete substrate.

Author Contributions

Conceptualization, H.J.; methodology, K.W., F.M., R.S. and H.J.; software, K.L. and K.W.; validation, H.J., K.W., H.F. and B.X.; formal analysis, H.F., K.W. and K.L.; investigation, H.J. and K.W.; resources, H.J.; data curation, H.F. and B.X.; writing—original draft preparation, H.F. and B.X.; writing—review and editing, K.W., R.S. and H.J.; visualization, K.W., F.M., K.L. and R.S.; supervision, H.J.; project administration, H.J.; funding acquisition, H.J. All authors have read and agreed to the published version of the manuscript.

Data Availability Statement

The original contributions presented in the study are included in the article; further inquiries can be directed to the corresponding author.

Acknowledgments

This project was undertaken at the plant of the Structural Laboratory of Guangdong University of Technology in Guangzhou City, China. The authors would like to acknowledge their generous support.

Conflicts of Interest

The authors declare no conflict of interest.

Footnotes

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Figures and Tables

Figure 1 UHPC jacket-reinforced NRC T-beam bridge.

Figure 2 UHPC-jacketing rehabilitation of seriously damaged concrete girder (Xinhua Expressway, photograph taken by Haibo Jiang).

Figure 3 Geometric details of specimens (Unit: mm): (a) dimensions of specimen ZJ-X; (b) dimensions of specimen GJ-X; (c) reinforcement layout of the ZJ group specimens; (d) reinforcement layout of the GJ group specimens.

Figure 4 Fabrication procedure of test specimens: (a) specimen preparation process; (b) NRC matrix interface treatment.

Figure 5 Push-out test setup.

Figure 6 Failure modes of push-out specimens: (a) Specimen ZJ-12; (b) Specimen ZJ-16; (c) Specimen GJ-12; (d) Specimen GJ-16.

Figure 7 Load–slip curves of push-out specimens: (a) Group ZJ; (b) Group GJ.

Figure 8 Comparison under the same construction method: (a) comparison of first peak load (P₁); (b) comparison of residual load (P_r).

Figure 9 Comparison under the same bolt diameter: (a) comparison of first peak load (P₁); (b) comparison of residual load (P_r).

Figure 10 Comparison between calculated values and experimental values.

Table 1

Test parameters.

Number	Specimens	Bolt Diameter (mm)	Bolt Position	Construction Method	Interface Treatment Method
1	ZJ-12	12	UHPC cover thickness of20 mm	Cast-in-place UHPC	Conventional chiseling, exposing NRC coarse aggregate
2	ZJ-16	16
3	GJ-12	12	Extending 50 mm over UHPC layer	Prefabricated UHPC panels	Conventional chiseling, exposing NRC coarse aggregate, and filling high-strength mortar between NRC and UHPC
4	GJ-16	16

Table 2

Mix proportions for concretes.

Concrete Type (kg/m3)	Cement	Microsilica Fume	Nano CaCO3	Coarse Aggregate	Fine Aggregate	Water	Superplasticizer	Steel Fiber
C50	408	/	/	1124	663	200	/	/
High-strength mortar	828	214	33	/	1024	192	28	/
UHPC	828	214	33	/	1024	192	28	156

Table 3

Mechanical properties of the types of concrete.

Concrete Type	Compressive Strength (MPa)	Splitting Tensile Strength (MPa)	Elastic Modulus (GPa)
C50	52.34	2.41	34.6
High-strength mortar	103.55	4.03	34.2
UHPC	138.28	10.77	43.7

Table 4

Experimental results of push-out specimens.

Specimens	P1 (kN)	τ1 (MPa)	d1 (mm)	P2 (kN)	d2 (mm)	P2/P1	Pr (kN)
ZJ-12	920.17	2.19	0.02	910.62	0.99	0.99	594.12
ZJ-16	1048.07	2.50	0	975.53	3.36	0.93	411.92
GJ-12	838.08	2.00	0.09	838.08	0.09	1.00	345.13
GJ-16	1204.20	2.87	0.16	713.99	0.39	0.59	287.45

Table 5

Requirements and limitations for formulae.

Shear Friction	ACI	AASHTO (2017)	Fib
Requirements	If using equation to calculate required Avf, then Vn ≤ μAvf fy	Avf > 0.05 Acv/fy	/
Limitations	0.2fc’ × Avf and 11.03 MPa and 3.31 + 0.08 fc’	0.25fc’ × Avf and 10.34 MPa	0.3 v × fc’ × Avf,where, v = 0.55 (30/fc’)1/3 < 0.55

Noted: All symbols are consistent with those above.

Table 6

Comparisons between experimental results and theoretical results.

Standard	Specimen	Vcc (kN)	Vbc (kN)	Vb (kN)	Shear Capacity		Calculated/Experimental
					Calculated Value (kN)	Experimental Value (kN)	Ratio	Average	Standard Deviation
ACI [46]	ZJ-12	/	215.38	/	215.38	920.17	0.234	0.294	0.052
	ZJ-16	/	382.90	/	382.90	1048.07	0.365
	GJ-12	/	215.38	/	215.38	838.08	0.257
	GJ-16	/	382.90	/	382.90	1204.20	0.318
AASHTO [47]	ZJ-12	588.00	215.23	/	803.23	920.17	0.873	0.891	0.058
	ZJ-16	588.00	382.63	/	970.63	1048.07	0.926
	GJ-12	588.00	215.23	/	803.23	838.08	0.958
	GJ-16	588.00	382.63	/	970.63	1204.20	0.806
Fib [48]	ZJ-12	676.87	106.54	34.33	817.74	920.17	0.889	0.902	0.080
	ZJ-16	676.87	189.40	61.04	927.31	1048.07	0.885
	GJ-12	720.11	106.54	34.33	860.98	838.08	1.027
	GJ-16	720.11	189.40	61.04	970.54	1204.20	0.806

Note: In the table, V_cc represents the shear capacity provided by the bond between the interface of concrete layers on both sides of the interface (kN), V_bc represents the shear capacity provided by the shear-friction effect of the bolts (kN), and V_b represents the shear capacity provided by the dowel action between the bolts and the concrete (kN).

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