- Open Access
A Review on Structural Behavior, Design, and Application of Ultra-High-Performance Fiber-Reinforced Concrete
© The Author(s) 2016
- Received: 28 February 2016
- Accepted: 11 April 2016
- Published: 3 May 2016
An overall review of the structural behaviors of ultra-high-performance fiber-reinforced concrete (UHPFRC) elements subjected to various loading conditions needs to be conducted to prevent duplicate research and to promote its practical applications. Thus, in this study, the behavior of various UHPFRC structures under different loading conditions, such as flexure, shear, torsion, and high-rate loads (impacts and blasts), were synthetically reviewed. In addition, the bond performance between UHPFRC and reinforcements, which is fundamental information for the structural performance of reinforced concrete structures, was investigated. The most widely used international recommendations for structural design with UHPFRC throughout the world (AFGC-SETRA and JSCE) were specifically introduced in terms of material models and flexural and shear design. Lastly, examples of practical applications of UHPFRC for both architectural and civil structures were examined.
- ultra-high-performance fiber-reinforced concrete
- bond performance
- structural behavior
- design code
Ultra-high-performance fiber-reinforced concrete (UHPFRC), which was developed in the mid-1990s, has attracted much attention from researchers and engineers for practical applications in architectural and civil structures, because of its excellent mechanical performance, i.e., compressive strength is greater than 150 MPa and a design value of tensile strength is 8 MPa (AFGC-SETRA 2002), durability, energy absorption capacity, and fatigue performance (Farhat et al. 2007; Graybeal and Tanesi 2007; Yoo et al. 2014c; Li and Liu 2016). In particular, its very high strength properties result in a significant decrease in the structural weight, i.e., the weight of UHPFRC structures is about 1/3 (or 1/2) of the weight of general reinforced concrete (RC) structures at identical external loads (Tam et al. 2012). Consequently, slender structures, which are applicable for long-span bridges, can be fabricated with UHPFRC, leading to low overall construction costs.
In order to apply such a newly developed innovative material to real structures, numerous studies have been carried out in many countries in Europe, North America, and Asia. Since UHPFRC was first developed by France’s research group, the first technical recommendation on UHPFRC for both material properties and structural design was introduced in France in 2002 and was called the AFGC-SETRA recommendation (AFGC-SETRA 2002). Thereafter, a state-of-the-art report on UHPFRC covering all material and design aspects was published in Germany in 2003 (DAfStB 2003). Then, in 2004, the Japan Society of Civil Engineers (JSCE) published their own design recommendations for UHPFRC based on Ductal® (Orange et al. 1999), commercial UHPFRC available in the world, (JSCE 2004). Lastly, in recent years, the Korea Concrete Institute (KCI) also developed a design code for UHPFRC (KCI 2012), similar to those in France and Japan, by using K-UHPC, another UHPFRC material developed by Korea Institute of Civil Engineering and Building Technology (Kim et al. 2008).
Due to the superb fiber bridging capacities of UHPFRC at cracked surfaces, leading to a special strain-hardening (or deflection-hardening) response with multiple micro-cracks, many researchers have focused on using it to the structures dominated by flexure, shear, and torsion. Furthermore, UHPFRC has also been considered as one of the promising materials for impact- and blast-resistant structures, because of its enhanced strength and energy absorption capacity, along with strain-hardening cementitious composites containing polymeric fibers (Astarlioglu and Krauthammer 2014; Choi et al. 2014). These properties can help to overcome the brittle failure of plain concrete, which has poor energy absorption capacity for impacts and blasts. Since the structural behaviors of UHPFRC under flexure, shear, and torsion, and when subjected to high-rate loadings, such as impacts and blasts, are highly sensitive to numerous factors, i.e., structural shape, loading condition, strain-rate, casting method, reinforcement ratio, etc., it is necessary to synthetically review the scattered studies.
The purpose of this research is to analyze the current state of knowledge of the structural behavior, design techniques, and applications of UHPFRC under various loading conditions. As explained above, the attention of this paper is focused on (1) bond performance between UHPFRC and various reinforcements, which is basic information needed for the design of reinforced structures, (2) structural behavior of UHPFRC under flexure, shear, torsion, and high-rate loading, (3) the most widely used UHPFRC design recommendations in the world, and (4) examples of practical applications in both architectural and civil structures.
Roy et al. (1972) and Yudenfreund et al. (1972) first introduced ultra-high-strength cementitious paste with low porosity in the early 1970s. With special curing methods using heat (250 °C) and pressure (50 MPa), Roy et al. (1972) achieved a cementitious paste with almost zero porosity and a compressive strength of approximately 510 MPa. On the other hand, Yudenfreund et al. (1972) obtained a cement paste having a compressive strength of about 240 MPa with normal curing (25 °C) for 180 days. To do this, Yudenfreund et al. (1972) provided a special treatment on-ground clinker, used the low water-to-cement ratio of 0.2, and Blaine surface areas ranging from 6000 to 9000 cm2/g. After nearly 10 years, Birchall et al. (1981) and Bache (1981) could develop two types of ultra-high-strength paste (or concrete) with very low porosity, such as densified with small particles (DSPs) concrete and macro-defect free (MDF) paste, by developing a pozzolanic admixture and a high-range water-reducing agent. Birchall et al. (1981) achieved the development of the cement pastes with compressive strength over 200 MPa and flexural strengths of 60–70 MPa, by removing macroscopic flaws during material preparation without using fibers or high-pressure compaction. Bache (1981) also successfully developed the concrete that was DSPs and had a compressive strength of 120–270 MPa. The key technique to densely pack the spaces between the cement particles was to use ultra-fine particles and an extremely low water content, with a large quantity of high-range water-reducing agent. In the mid-1990s, Richard and Cheyrezy (1995) first introduced the concept of and mixing sequence for reactive powder concrete (RPC), which was the forerunner of UHPFRC. To obtain a very high strength, the granular size was optimized by the packing density theory, by excluding coarse aggregate and by providing heat (90 and 400 °C) and pressure treatments. In addition, 1.5–3 % (by volume) of straight steel microfibers, with a diameter of 0.15 mm and a length of 13 mm, were added to achieve high ductility; consequently, the RPC developed by Richard and Cheyrezy (1995) showed compressive strengths of 200–800 MPa and fracture energies up to 40 kJ/m2.
3.1 Bond Behavior Between UHPFRC and Reinforcements
In addition, CMR model (Cosenza et al. 1995), which sets τ = τ max × (1 − e−s/sr ) β , was found to be appropriate for simulating the ascending bond stress versus slip behavior of steel bars embedded in UHPFRC, and the parameters were proposed as τ max = 5.0f c ′0.5, s r = 0.07, and β = 0.8, where s r and β are coefficients based on the curve fitting of test data.
Schäfers and Seim (2011) performed experimental and numerical investigations on the bond performance between timber and UHPFRC. The glued-laminated timber was bonded to sandblasted and ground UHPFRC with the “Sikadur 330” epoxy resin. Regardless of the bond length and surface treatment, most of specimens showed failure of the bond in the timber close to the bond-line. Based on the Volkersens theory, Schäfers and Seim (2011) suggested a bond length of 400 mm for standard test method to evaluate the bond strength of timber-concrete composites and noted that the effect of tensile stresses, orthogonal to the bond-line, can be neglected when the bond length is beyond 300 mm.
3.2 Flexural Dominated Reinforced UHPFRC Beams, Girders, and Composite Structures
Due to its excellent post-cracking tensile performance with multiple micro-cracks occurred, UHPFRC has attracted attention from engineers for application in structural elements subjected to bending. Several international recommendations (AFGC-SETRA, JSCE, and KCI) from France, Japan, and South Korea (AFGC-SETRA 2002; JSCE 2004; KCI 2012) thus provide stress–strain models for compressive and tensile stress blocks in the cross-section, as well as the detailed process of predicting the ultimate capacity of UHPFRC elements under flexure. Since strain (and stress) distribution in the cross-section varies according to the curvature of a beam, multilayer sectional analysis (Yoo and Yoon 2015) is required to calculate an appropriate neutral axis depth and moment capacity at a certain curvature level.
In order to establish reasonable design codes for UHPFRC, Yang et al. (2010) carried out several four-point flexural tests for UHPFRC beams having reinforcement ratios less than 0.02. Test variables were the amount of steel rebar and the placement method. From their test results, placing concrete at the ends of the beams yielded better performance than when concrete was placed at the mid-length because of better fiber orientation to the direction of beam length at the maximum moment zone. In addition, they reported that all test beams showed a ductile response with the ductility index ranging from 1.60 to 3.75 and were effective in controlling cracks. However, the meaning of ‘ductile response’ could be incorrectly delivered to readers because no test results of reinforced UHPC beams without fiber were reported. In accordance with the test results by Yoo and Yoon (2015), reinforced UHPFRC beams exhibited lower ductility indices compared to beams without fiber due to the crack localization behavior, and Dancygier and Berkover (2016) also reported that the inclusion of steel fibers resulted in a decrease of flexural ductility of beams with low conventional reinforcement ratios.
Ferrier et al. (2009) also examined the flexural behavior of a new type of hybrid beam, made of glued-laminated wood and UHPFRC planks, including steel and FRP rebar. They mention that structural efficiency was obtained by using the hybrid beams, as a consequence of the increased bending stiffness due to the high elastic modulus of UHPFRC planks. In addition, the inclusion of steel and FRP rebar in the lower UHPFRC plank significantly increased the ultimate load capacity of the hybrid beams, as compared with when only pure wood elements were used. These advantages of using hybrid beams lead to the potential for reducing the beam depth or increasing the span length of the beam, compared with conventional timber structures.
3.3 Shear Resistance of Structural UHPFRC Beams, Girders, and Bridge Decks
In order to replace the open-grid steel decks from moveable bridges, which have several drawbacks, such as poor rideability, high noise levels, susceptibility to fatigue damage, and high maintenance costs, Saleem et al. (2011) examined the structural performance of lightweight UHPFRC bridge decks reinforced with high-strength steel rebar. They properly designed and proposed UHPFRC waffle decks to satisfy the strength, serviceability, and self-weight requirements for moveable bridges. The governing failure mode was shear, and in the multi-unit decks, shear failure was followed by punching shear failure at close to the ultimate state. However, the shear failure was less abrupt and catastrophic as compared with the commonly seen shear failure mode. Thus, Xia et al. (2011) recommended the ductile shear failure with higher post-cracking shear resistance of UHPFRC beams containing high-strength steel rebar as an acceptable failure mode, rather than including transverse reinforcements, because of their economic problems. The use of 180° hooks at both ends of the steel rebar, recommended by ACI 318 (ACI 2014), was also effective in avoiding bond failure, compared with specimens without end anchorage. Based on a thorough analysis of the experimental results, Saleem et al. (2011) noted that although the proposed UHPFRC waffle deck system exhibits shear failure mode, it has great potential to serve as an alternative to open-grid steel decks, which are conventionally used for lightweight or moveable bridges.
3.4 Torsional Behavior of Structural UHPFRC Beams and Girders
3.5 Performance of Structural UHPFRC Beams, Slabs, and Columns Under Extreme Loadings
Fujikake et al. (2006a) and Yoo et al. (2015a, c) examined the impact resistance of reinforced or prestressed UHPFRC beams by testing a number of specimens using a drop-weight impact test machine. In their studies (Fujikake et al. 2006a), an increase in the maximum deflection of UHPFRC beams was observed by increasing the drop height while maintaining the weight of the hammer, owing to the increase of kinetic energy. The initial stiffness in the UHPFRC beams was insignificantly affected by the impact damage because of the excellent fiber bridging capacities after matrix cracking, and the residual load–deflection (or moment–curvature) curves, shifted based on the maximum deflection by impact, exhibited quite similar behaviors with those of the virgin specimens without impact damage. Hence, Fujikake et al. (2006a) mentioned that the maximum deflection response can be used as the most rational index for estimating the overall flexural damage of reinforced UHPFRC beams. Yoo et al. (2015a) reported that better impact resistance, i.e., lower maximum and residual deflections and higher deflection recovery, was obtained by increasing the amount of longitudinal steel rebars, and the maximum and residual deflections of reinforced UHPFRC beams decreased significantly by adding 2 % (by volume) of steel fibers, leading to a change in the damage level from severe to moderate, whereas slight decreases in the maximum and residual deflections were found by increasing the fiber length at identical volume fractions (Yoo et al. 2015c). A higher ultimate load capacity was also obtained for the beams under impact loading, compared to those under quasi-static loading, and the residual load capacity after impact damage improved by including 2 % steel fibers and using the longer steel fibers. Fujikake et al. (2006a) and Yoo et al. (2015a) successfully predicted the mid-span deflection versus the time response of structural UHPFRC beams by using the sectional analysis and single- (or multi-) degree-of-freedom model. Improved mechanical compressive and tensile strengths according to the strain-rate were considered in the analysis by using the equations for the dynamic increase factor (DIF) of the UHPFRC, as suggested by Fujikake et al. (2006b, 2008).
4.1 AFGC-SETRA Recommendations
4.1.1 Material Models
In AFGC-SETRA recommendations, UHPFRC is referred to as a cementitious material with a compressive strength in excess of 150 MPa, possibly obtaining 250 MPa, and including steel (or polymer) fibers to provide a ductile tensile behavior. The parameters of the design strength were suggested based on the mechanical test results of Ductal® (Orange et al. 1999), as follows: f ck = 150–250 MPa, f tj = 8 MPa, and E c = 55 GPa, where f ck is the compressive strength, f tj is the post-cracking direct tensile strength, and E c is the elastic modulus. A partial safety factor γ bf is also introduced, with γ bf = 1.3 in the case of fundamental combinations and γ bf = 1.05 in the case of accident combinations. To consider the fiber orientation effect on the tensile behavior, three different fiber orientation coefficients were suggested, as follows: K = 1 for placement methods validated from test results of a representative model of actual structure, K = 1.25 for all loading other than local effects, and K = 1.75 for local effects.
The tensile stress–crack opening displacement (σ–w) model is recommended to be first derived based on an inverse analysis. To apply the σ–w model into the tensile stress block in the cross-section, it needs to be transformed to the tensile stress–strain (σ–ε) model by using the characteristic length, l c , which is l c = 2/3 × h for the case of rectangular or T-beams, where h is the beam height.
The ultimate tensile strain is expressed by ε lim = L f /4l c , where ε lim is the ultimate tensile strain and L f is the fiber length.
The completed material models under compression and tension are given in Fig. 14. Based on the capacity of tensile resistance with crack opening displacement, the AFGC-SETRA recommendations classify the tensile response by two different laws: (1) the strain-softening law (f tj > f bt ) and (2) the strain-hardening law (f tj < f bt ), as shown in Fig. 14.
Poisson’s ratio: 0.2.
Thermal expansion coefficient (×10−6/°C): 11.0.
Long-term creep coefficient: 0.2.
Total (autogenous) shrinkage: 550 × 10−6.
4.1.2 Flexural Design
4.1.3 Shear Design
4.2 JSCE Recommendations
4.2.1 Material Models
Poisson’s ratio: 0.2.
Coefficient of thermal expansion (×10−6/°C): 13.5.
Thermal conductivity (kJ/mh °C): 8.3.
Thermal diffusivity (×10−3m2/h): 3.53.
Specific heat (kJ/kg °C): 0.92.
Total shrinkage: 550 × 10−6.
Creep coefficient: 0.4.
Density (kN/m3) for calculating dead load: 25.5.
4.2.2 Flexural Design
In the JSCE recommendations, for the structural design of UHPFRC elements under bending, two simple assumptions are required to be satisfied: (1) the linear strain distribution and (2) the use of the proposed material models, as given in Fig. 16. The steel fiber contribution in the tensile zone after cracking needs to be considered in the structural design, and the compressive and tensile stress blocks in the cross-section should be considered based on the proposed material stress–strain models (Fig. 16) by considering the equivalent specific length for tension. Although a detailed procedure for calculating the ultimate moment capacity is not introduced in the JSCE recommendations, sectional analysis can be adopted for calculating the moment–curvature behavior of the UHPFRC elements (Yang et al. 2011; Yoo et al. 2016), similar to the case of the AFGC-SETRA recommendations.
4.2.3 Shear Design
As was reported by Yang et al. (2012), the shear design of UHPFRC elements can be carried out by using both AFGC-SETRA and JSCE recommendations, which provide good estimates with test data of I-shaped UHPFRC beams, as shown in Fig. 9.
The use of 2 % steel fibers resulted in a higher post-cracking stiffness and ultimate load capacity of steel bar-reinforced UHPFRC beams, but it decreased their ductility because of the superb bond strength with steel bars and crack localization characteristics. Importantly, the use of UHPFRC was effective in overcoming the major drawbacks of conventional FRP-reinforced concrete structures (large service deflection) due to its strain-hardening response, and the use of CFRP bars was efficient in improving the flexural stiffness of reinforced UHPFRC beams, compared to that of GFRP bars. The hybrid reinforcements (steel + FRP bars), which have been adopted to reduce the service deflection of conventional FRP-reinforced beams, were ineffective in UHPFRC beams, and thus, the use of single FRP bars, instead of hybrid reinforcements, was recommended for the case of UHPFRC.
With the inclusion of 2.5 % steel fibers, approximately 250 % higher shear strength was obtained, compared to that without fibers. The shear strength also increased with an increase in the fiber contents and a decrease in the shear span-to-depth ratio. Due to the excellent fiber bridging at crack surfaces, the shear failure of UHPFRC beams was less abrupt than the commonly seen shear failure mode in conventional concrete. In addition, both AFGC-SETRA and JSCE recommendations provided accurate predictions of shear strength of UHPFRC beams.
The inclusion of steel fibers provided a better cracking performance, higher ultimate and cracking torque, and improved torsional stiffness. The higher ultimate torque was also found by increasing the ratios of stirrups and longitudinal rebars. The diagonal crack angle was influenced by the amount of stirrups and longitudinal rebars, whereas it was not affected by the amount of steel fibers.
Better impact resistance of reinforced UHPFRC beams was obtained by including 2 % steel fibers and a larger number of longitudinal steel bars. The residual capacity after impact damage was also improved by adding 2 % steel fibers and using the longer steel fibers. UHPFRC columns exhibited significantly higher blast resistance such as lower maximum deflection, improved damage tolerance, and higher resistance, compared with the reinforced SCC and NC columns. Thus, it was noted that the application of UHPFRC in the structures subjected to blast loads is effective.
Finally, the international design recommendations on UHPFRC were discussed minutely, and examples of practical applications of UHPFRC in architectural and civil structures were investigated.
This research was supported by a grant from a Construction Technology Research Project 13SCIPS02 (Development of impact/blast resistant HPFRCC and evaluation technique thereof) funded by the Ministry of Land, Infrastructure and Transport.
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