A Review on Structural Behavior, Design, and Application of Ultra-High-Performance Fiber-Reinforced Concrete
International Journal of Concrete Structures and Materials
A Review on Structural Behavior, Design, and Application of Ultra-High-Performance Fiber-Reinforced Concrete
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; application
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
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
a state-of-the-art report on UHPFRC covering all material
and design aspects was published in Germany in 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,
Lastly, in recent years, the Korea Concrete Institute (KCI)
also developed a design code for UHPFRC
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.
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.
2. Historical Development of UHPFRC
Roy et al. (1972)
Yudenfreund et al. (1972)
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.
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
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)
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
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,
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
Richard and Cheyrezy (1995)
compressive strengths of 200–800 MPa and fracture energies up
to 40 kJ/m2.
3. Performance of Structural UHPFRC
3.1 Bond Behavior Between UHPFRC and Reinforcements
In order to practically apply a newly developed UHPFRC
in the structures, bond performance with reinforcements
should be examined. Many researchers
Muttoni 2004; Ahmad Firas et al. 2011; Yoo et al. 2014a, b,
have investigated the bond behavior of internal steel
and fiber-reinforced polymer (FRP) reinforcements with
Jungwirth and Muttoni (2004)
carried out pullout
test of deformed steel reinforcing bar using a 160 mm cube.
Various bond lengths ranging from 20 to 50 mm and two
different bar diameters of 12 and 20 mm were adopted. In
their study, the average bond strength of steel bars embedded
in UHPFRC was found to be 59 MPa, approximately 10
times higher than the bond strength of steel bars embedded
in ordinary concrete, and the theoretical development length
of deformed steel bars in UHPFRC was suggested by
lb = fydb/4smax, where fy is the yield strength of steel bar, db
is the nominal diameter of steel bar, and smax is the bond
strength. Yoo et al. (2014c) examined the effects of fiber
content and embedment length on the bond behavior of
deformed steel bars embedded in UHPFRC. For this, Yoo
et al. (2014c) performed a number of pullout tests by
modifying the test method, proposed by RILEM
; the 150 mm cubic specimens
with a single bar embedded vertically along the central axis
were fabricated and used for testing. The embedment lengths
were determined by 1 and 2 times the bar diameter, instead
of using 5db, as suggested by the RILEM recommendations.
The bond strength was insignificantly affected by the fiber
content and embedment length, but it clearly correlated with
the compressive strength. The CEB-FIP Model Code
, which defined smax as 2.0fc00.5,
substantially underestimated the bond strength of steel bars
in UHPFRC because the parameters were suggested based
on test data from previous concretes. Thus, Yoo et al.
(2014c) proposed modified coefficients for the bond strength
of steel bars in UHPFRC, based on a number of test data, as
follows (Fig. 1):
In addition, CMR model
(Cosenza et al. 1995)
, which sets
s = smax 9 (1 - e-s/sr)b, 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 smax = 5.0fc00.5, sr = 0.07, and b = 0.8, where
sr and b are coefficients based on the curve fitting of test data.
Ahmad Firas et al. (2011) experimentally investigated the
bond performance between carbon-fiber-reinforced polymer
(CFRP) bars and UHPFRC according to the surface
treatment, embedment length, bar diameter, and concrete age.
Based on the test data, it was noted that the bond strength
was insignificantly affected by the surface treatment of the
glass-fiber-reinforced polymer (GFRP) bar; similar bond
strengths for smooth bars and sand-coated bars were
obtained. On the other hand, a decrease in bond strength was
obtained by increasing both bar diameter and embedment
length. The ultimate bond strength of CFRP bars in
UHPFRC was insignificantly changed by age after 3 days,
because it was primarily affected by the shear strength of the
connection between the core and the outer layer of the CFRP
bars. Ahmad Firas et al. (2011) suggested a development
length for sand-coated bars of approximately 40db, and a
development length for a smooth bar of longer than 40db.
Yoo et al. (2015b) also examined the local bond behavior of
GFRP bars embedded in UHPFRC. The average bond
strengths of GFRP bars in UHPFRC were found to be from
16.7 to 22.8 MPa for a db of 12.7 mm, and from 19.3 to
27.5 MPa for a db of 15.9 mm, which are approximately 73
and 66 % less, respectively, than the bond strengths of
deformed steel bars. Similar to the case of CFRP bars in
(Ahmad Firas et al. 2011)
, bond failure was
generated by the delamination of the resin and fiber in the
bar. Based on a database of 68 pullout test results for GFRP
bars in UHPFRC, Yoo et al. (2015b) suggested an equation
for the relationship between normalized bond strength and
development length by using regression analysis and by
assuming no influence of the normalized cover parameter on
bond strength, as follows (Fig. 2b):
where u is the bond strength (=smax), db is the bar diameter,
and Le is the embedment length.
The American Concrete Institute (ACI) 440.1R model
was inappropriate for UHPFRC; it significantly
overestimated the test data (normalized bond strength), as
shown in Fig. 2a.
Yoo et al. (2015b) also pointed out that the previous model
for development length of FRP bar in concrete, suggested by
Wambeke and Shield (2006)
, was not appropriate for
UHPFRC; thus, they proposed an expression for the
development length of GFRP bars in UHPFRC, which is only
valid for the case of pullout failure, as follows:
Ld;pullout ¼ 3:4 f 0
where Ld,pullout is the development length and ffu is the
ultimate strength of rebar.
Scha¨fers 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
Scha¨fers and Seim (2011)
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
(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)
to calculate an appropriate neutral axis depth and moment
capacity at a certain curvature level.
Yoo and Yoon (2015)
first reported test results of a number
of reinforced UHPFRC beams to investigate the effects of
steel fiber aspect ratio and type on flexural performance.
Since a portion of the tensile stress after cracking was
resisted by the steel fibers, low reinforcement ratios (percent)
of 0.94 and 1.50 % were selected. In order to prevent brittle
shear failure, stirrups were conservatively designed based on
the specimens made of ultra-high-performance concrete
(UHPC) without fibers. From the test results (Fig. 3), the
beams made by UHPFRC with 2 % by volume of steel fibers
exhibited much higher post-cracking stiffness and ultimate
load capacity, compared to those made by UHPC without
fiber, called ‘NF’. In addition, the use of long straight or
twisted steel fibers (S19.5, S30, and T30) led to a higher
ductility than the use of short straight steel fibers (S13),
which are applied for commercial UHPFRC available in
, at the identical fiber
volume fraction. However, it is very interesting to note that
much lower ductility indices were obtained by including
steel fibers. This is caused by the fact that due to the very
high bond strength between UHPFRC and steel rebar and its
crack localization behavior, the steel rebar ruptured at a
relatively smaller mid-span deflection, as compared with
UHPC beams without fiber. Thus,
Yoo and Yoon (2015)
concluded that the strain-hardening behavior of UHPFRC
was unfavorable to the ductility of reinforced beams.
In order to establish reasonable design codes for
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)
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.
Yang et al. (2011)
examined the flexural behavior of
largescale prestressed UHPFRC I-beams. They indicated that the
high volume content of steel fibers in UHPFRC effectively
controlled the increase in crack widths, and led to multiple
micro-cracks due to the fiber bridging at crack surfaces. The
flexural strength of prestressed UHPFRC I-beams was
insignificantly affected by the presence of stirrups.
also investigated the flexural behavior of a full-scale
prestressed UHPFRC I-girder (AASHTO Type II girder)
containing 26 prestressing strands. Based on the
experimentally observed behavior, he reported that a UHPFRC
Igirder shows larger flexural capacities than that of a
conventional concrete girder with similar cross-sectional
geometry. In addition, an inversely proportional relationship
between crack spacing and maximum tensile strain was
experimentally observed, as shown in Fig. 4, and the
following equation was suggested:
25ffi2ffiffi0ffiffi þ 2p5; 8ffiffi0ffiffiffi0
e ¼ 450 þ pscr
where e is the tensile strain and scr is the crack spacing.
In recent years, several studies
(Ferrier et al. 2015; Yoo
et al. 2016)
have been carried out to develop a new type of
Fig. 3 Load-deflection curves of steel bar-reinforced UHPFRC/UHPC beams; a q = 0.94 %, b q = 1.50 % [NF = UHPC w/o
fiber, S13 = UHPFRC w/ straight steel fibers (Lf/df = 13/0.2 mm/mm), S19.5 = UHPFRC w/ straight steel fibers (Lf/
df = 19.5/0.2 mm/mm), S30 = UHPFRC w/ straight steel fibers (Lf/df = 30/0.3 mm/mm), T30 = UHPFRC w/ twisted steel
fibers (Lf/df = 30/0.3 mm/mm)]
(Yoo and Yoon 2015)
high-performance lightweight beams by applying UHPFRC
and FRP rebar.
Ferrier et al. (2015)
structural behavior of I-shaped UHPFRC beams reinforced with
CFRP and GFRP rebar, according to the rebar axial stiffness
ranging from 9 MN to 30 MN. Experimental results
indicated that the CFRP rebar was effective in increasing the
bending stiffness, which results in a lower mid-span
deflection, as compared with the case of the GFRP rebar due
to the higher elastic modulus of the former. Thus, they
concluded that the axial stiffness of the FRP reinforcement is
the most influential parameter of bending stiffness of beams.
Yoo et al. (2016)
also examined the flexural behavior of
UHPFRC beams reinforced with GFRP rebar and hybrid
reinforcements (steel ? GFRP rebar), according to the axial
stiffness ranging from 13 to 95.5 MN. Hybrid
reinforcements were considered in their study because it has been
considered as one of the most promising methods to
overcome the large service deflection problems of conventional
FRP-reinforced concrete beams reported by several
(Lau and Pam 2010; Yoon et al. 2011)
. Due to
the strain-hardening characteristics of UHPFRC, all tested
beams provided very stiff load versus deflection response
even after the formation of cracks (Fig. 5), which is
distinctive response with conventional FRP-reinforced concrete
beams, and satisfied the service crack width criteria of the
Canadian Standards Association (CAN/CSA) S806 (CAN
2002). Furthermore, the deformability factors suggested by
Jaeger et al. (1995)
were higher than the lower limit (Df = 4)
for all test beams. Therefore,
it was noted that the use of UHPFRC could be a new
solution for solving the major drawbacks limiting the
practical application of FRP rebar instead of steel rebar. An
increase in the GFRP reinforcement ratio led to an
improvement in the flexural performance, such as higher
post-cracking stiffness, load carrying capacity, and ductility.
However, the application of hybrid reinforcements to
UHPFRC nullified the main advantage of using FRP to solve
the corrosion problem and showed insignificant
improvement in the structural performance. Synthetically,
Yoo et al.
recommended the use of GFRP rebar with UHPFRC,
rather than the use of hybrid reinforcements.
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
First cracking ( 54.5 kN)
Number of cracks
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.
To practically apply UHPFRC in real architectural and
civil structures, appropriate design technique should be
suggested based on the material models. Several
(AFGC-SETRA 2002; JSCE 2004;
thus provide material models for designing
flexural members made of UHPFRC. Based these
recommendations, many researchers have already precisely
predicted the flexural behaviors of reinforced UHPFRC beams
(Yang et al. 2010; Yang et al. 2011; Ferrier et al. 2015; Yoo
and Yoon 2015; Yoo et al. 2016)
. In particular, the UHPFRC
beams without stirrups were well predicted by
AFGCSETRA recommendations without consideration of fiber
orientation coefficient (K = 1)
(Yang et al. 2011; Yoo et al.
because the fiber alignment in the direction of beam
length was insignificantly disturbed by the internal rebars.
Yoo and Yoon (2015)
recently reported that the
fiber orientation coefficient that is proposed by the
AFGCSETRA recommendations (i.e., K = 1.25), should be
considered for simulating the flexural behavior of reinforced
UHPFRC beams with stirrups, since the fiber orientation was
clearly disturbed by the stirrups (Fig. 6). In the case of
FRPreinforced concrete elements, it is well known that the
service deflection prediction is the most important parameter
for designing such structures, because of the larger service
deflection than that of beams reinforced with steel rebar.
Ferrier et al. (2015)
Yoo et al. (2016)
predicted the load versus deflection curves of
FRP-bar-reinforced UHPFRC beams by sectional analysis, in which they
considered compressive and tensile stress blocks in the
cross-section, similar to the method used for the above
Yoo and Banthia (2015)
accurately predicted the service deflection of UHPFRC beams
reinforced with GFRP rebar and hybrid reinforcements
(steel ? GFRP rebar), based on a micromechanics-based
finite element (FE) analysis; the average ratios of the
serviceability deflections from predictions and experiments
were found to be 0.91 with a standard deviation of 0.07.
Num. analysis (K=1)
Num. analysis (K=1.25)
3.3 Shear Resistance of Structural UHPFRC
Beams, Girders, and Bridge Decks
Baby et al. (2013b) carried out shear tests of eleven 3-m
long UHPFRC I-shaped girders with various shear
reinforcements (stirrups and/or steel fibers, or neither) combined
with longitudinal prestressing or passive steel bars. To
examine the actual fiber orientation effect on the shear
performance, the three-point flexural tests were performed by
using notched prism specimens extracted from both of the
undamaged ends of I-girders at different inclination angles.
Test results, as shown in Fig. 7, clearly indicated that the
fiber orientation significantly influenced the mechanical
(flexural) performance; thus, they noted that the actual fiber
orientation needs to be taken into account for shear design,
as recommended by AFGC-SETRA recommendations
. By including 2.5 % steel fibers, an
almost 250 % increase in shear strength was observed
et al. 2013c)
. The stirrups yielded first, while localization of
the shear crack took place significantly later, as shown in
Fig. 8. Thus, crack localization is primarily influenced by
the strain capacity of the UHPFRC, and the contributions of
the fiber bridging and the stirrups up to their yield strength
seem to be effective only when the tensile strain capacity of
the UHPFRC is much higher than the yield strain of the
stirrups. In their study
(Baby et al. 2013c)
AFGCSETRA recommendations were conservative for the
shearcracking strength, but reasonable for the ultimate shear
strength prediction of UHPFRC I-girders. Baby et al.
(2013a) also examined the feasibility of applying the
modified compression field theory for the shear capacity of
reinforced or prestressed UHPFRC beams. Based on their
analytical results, the modified compression field theory was
determined to be applicable for predicting the shear behavior
with an effective estimation of the reorientation of the
compressive struts with an increase in the load.
Voo et al. (2010)
investigated the shear strength of
prestressed UHPFRC I-beams without stirrups, according to the
shear span-to-depth ratio (a/d) and the type of steel fibers.
They indicated that a higher shear strength was obtained by
using a higher fiber volume content and a lower a/d. The
theory of the plastic shear variable engagement model
presented a good basis for their shear design and a good
relationship to the experimental results; the ratio of shear
strengths obtained from experiment and theory was found to
be 0.92, with a coefficient of variation of 0.12. In addition,
Bertram and Hegger (2012)
Yang et al. (2012)
Tadepalli et al. (2015)
mentioned that the shear strength increased
with an increase in the fiber content and a decrease in the a/d
ratio. For instance, the inclusion of 2.5 % steel fibers led to a
177 % higher ultimate load than that without fiber, and by
changing the a/d ratio from 3.5 to 4.4, the shear capacity was
reduced by 10 %
(Bertram and Hegger 2012)
also noted that the size effect on shear
strength was more substantially affected by the beam height
as compared with the web thickness, and that about
12–14 % higher shear capacity was obtained when the
effective prestressing force increased by 20 %. By
comparing the test results with computed values,
Yang et al. (2012)
noticed that the predictions using the AFGC-SETRA and
JSCE recommendations provided accurate estimates of the
shear strength of UHPFRC I-beams (Fig. 9).
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
selfweight 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)
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
, was also effective in
avoiding bond failure, compared with specimens without
end anchorage. Based on a thorough analysis of the
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
opengrid steel decks, which are conventionally used for
lightweight or moveable bridges.
3.4 Torsional Behavior of Structural UHPFRC
Beams and Girders
Empelmann and Oettel (2012)
examined the effect of
adding steel fibers (vf of 1.5 and 2.5 %) on the torsional
behavior of UHPFRC box girders. They experimentally
observed that the inclusion of steel fibers led to a better
cracking performance such as smaller crack widths and
multitudinous cracks, higher ultimate and cracking torque,
and improved torsional stiffness. Interestingly, the angle of
the diagonal cracks was found to be approximately 45 for
all test series, regardless of the steel fiber contents.
et al. (2013)
also investigated the torsional behavior of
UHPFRC beams reinforced with mild steel rebars. In order
to estimate the effects of steel fiber content and transverse
and longitudinal rebar ratios, thirteen UHPFRC beams were
fabricated and tested. Based on their test results
(Yang et al.
, an improvement in the initial cracking and ultimate
torque were obtained by increasing the fiber volume
fractions (Fig. 10a), which is consistent with the findings from
Empelmann and Oettel (2012)
. Moreover, higher ultimate
torque was found with increases in the ratio of stirrups with
longitudinal rebar (Figs. 10b and 11c). In addition, the
torsional stiffness after initial cracking was also improved by
increasing the ratio of stirrups, as shown in Fig. 10b. In
contrast to Empelmann and Oettel’s findings,
Yang et al.
reported that the angle of the diagonal compressive
stress ranged from 27 to 53 , and was affected by the
number of stirrups and longitudinal rebar. For example, the
angle of localized diagonal cracks increased with an increase
in the number of stirrups, as illustrated in Fig. 11.
and Ismail (2012)
also reported similar test results for the
torsional behavior of UHPFRC elements. They specifically
said that the inclusion of steel fibers was effective in
improving the torsional performance, such as cracking and
ultimate torsional capacities, torsional ductility,
post-cracking stiffness, and toughness. The use of longitudinal
reinforcements and stirrups also obviously improved the
Fig. 10 Torque-twist curves of UHPFRC beams (vf = volume fraction of steel fiber, qs = transverse reinforcement ratio,
ql = longitudinal reinforcement ratio); a effect of steel fiber content, b effect of transverse reinforcement ratio, c effect of
longitudinal reinforcement ratio
(Yang et al. 2013)
Fig. 11 Angle of localized diagonal cracks (vf = volume fraction of steel fiber, qs = transverse reinforcement ratio, ql =
longitudinal reinforcement ratio)
(Yang et al. 2013)
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.
Aoude et al. (2015)
investigated the blast resistance of
full-scale self-consolidating concrete (SCC) and UHPFRC
columns under various blast-impulse combinations based on
a shock-tube instrumentation. They verified that the steel
bar-reinforced UHPFRC columns showed substantially
higher blast resistance than the reinforced SCC columns in
terms of reducing the maximum and residual deflections,
enhancing damage tolerance, and eliminating secondary blast
fragments. Based on the single-degree-of-freedom (SDOF)
model and lumped inelasticity approach,
Aoude et al. (2015)
predicted the inelastic deflection-time histories. From the
numerical results, several important findings were obtained as
follows; (1) since the numerical predictions are sensitive to
the choice of DIF, as given in Fig. 12a, further study needs to
be done to develop the strain-rate models for using in the blast
analysis of UHPFRC columns, and (2) the plastic hinge
length (Lp) seems to be reduced in UHPFRC columns from
Lp = d (column effective depth), which has been used for the
analysis in conventional reinforced concrete columns, as
shown in Fig. 12b.
Astarlioglu and Krauthammer (2014)
numerically simulated the response of normal-strength
concrete (NC) and UHPFRC columns subjected to blast loadings
based on SDOF models using the dynamic structural analysis
suite (DSAS) and reported that the UHPFRC columns
presented lower mid-span displacement and sustained more than
four times the impulse as compared with the NC columns.
Wu et al. (2009)
carried out a series of blast tests of NC
and UHPFRC slabs w/ and w/o reinforcements to examine
their blast resistance. When the similar blast loads were
applied, the UHPFRC slabs without reinforcement exhibited
less damage than the NC slabs with reinforcements, and
thus, they noticed that the application of UHPFRC is
effective in blast design. The UHPFRC slab with passive
reinforcements was superior to all other slab specimens, and
the strengthening of NC slabs with external FRP plates in the
compressive zone was efficient in improving the blast
Yi et al. (2012)
examined the blast resistance of
the reinforced slabs made of NC, ultra-high-strength
concrete (UHSC), and RPC, which is identical to UHPFRC. By
analyzing the crack patterns and maximum and residual
deflections, they indicated that RPC has the best
blast-resistant capacity, followed by UHSC and then NC. For
example, the maximum deflections of NC, UHSC, and RPC
slabs from 15.88 ANFO charge were found to be 18.57,
15.14, and 13.09 mm, respectively.
Mao et al. (2014)
investigated the capability of modeling the impact behavior
of UHPFRC slabs using the commercial explicit FE
program, LS-DYNA (2007). Through FE analysis, they also
studied the effects of steel fibers and rebar on the blast
resistance of UHPFRC slabs. Importantly, they observed that
the K&C model (mostly used for simulating the blast
behavior of concrete structures) with automatically
generated parameters provided a much better ductile response than
the actual behavior, and thus, a modified parameter b2 from
1.35 to -2 should be applied for UHPFRC. After verifying
the numerical modeling with test data, a parametric study
was carried out, and some useful results were obtained: (1)
the additional use of steel fibers and rebar provide similar
influence in the form of extra resistance to the UHPFRC
panel under far field blast loading, and (2) under near field
blast loading, the resistance of the UHPFRC panels
increased substantially with steel rebar, as shown in Fig. 13.
4. Structural Design of UHPFRC Based
on AFGC-SETRA and JSCE
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: fck = 150–250 MPa,
ftj = 8 MPa, and Ec = 55 GPa, where fck is the compressive
strength, ftj is the post-cracking direct tensile strength, and Ec
is the elastic modulus. A partial safety factor cbf is also
introduced, with cbf = 1.3 in the case of fundamental
combinations and cbf = 1.05 in the case of accident
combinations. To consider the fiber orientation effect on the tensile
behavior, three different fiber orientation coefficients were
Fig. 12 Displacement predictions from SDOF analysis for UHPFRC columns (Note CRC means UHPFRC); a DIF = 1.14 vs. 1.4,
b hinge length, Lp = d vs. 0.5d
(Aoude et al. 2015)
þ cbf Ec
where ee is the elastic strain, w0.3 is the crack width of
0.3 mm, e0.3 is the strain at crack width of 0.3 mm, w1% is
the crack width corresponding to 0.01H (H is the specimen
height), and e1% is the strain at crack width corresponding to
The ultimate tensile strain is expressed by elim = Lf/4lc,
where elim is the ultimate tensile strain and Lf is the fiber
The stresses at two different characteristic points (at w0.3
and w1%) are expressed as follows:
where fbt is the stress at a crack width of 0.3 mm and f1% is
the stress at a crack width corresponding to 0.01H.
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
AFGCSETRA recommendations classify the tensile response by
two different laws: (1) the strain-softening law (ftj [ fbt) and
(2) the strain-hardening law (ftj \ fbt), as shown in Fig. 14.
Other important material properties of UHPFRC with heat
treatment are given by AFGC-SETRA recommendations, as
Poisson’s ratio: 0.2.
Thermal expansion coefficient (910-6/ C): 11.0.
Long-term creep coefficient: 0.2.
Total (autogenous) shrinkage: 550 9 10-6.
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.
In the case of compressive model, a bilinear stress–strain
model can be used, as shown in Fig. 14, with the design
strength and strain parameters—rbcu = 0.85fck/hcb and
eu = 0.003.
The tensile stress–crack opening displacement (r–
w) model is recommended to be first derived based on an
inverse analysis. To apply the r–w model into the tensile
stress block in the cross-section, it needs to be transformed
to the tensile stress–strain (r–e) model by using the
characteristic length, lc, which is lc = 2/3 9 h for the case of
rectangular or T-beams, where h is the beam height.
To obtain the tensile r–e model, the elastic tensile strain and
strains at crack widths of 0.3 mm and 1 % of beam height
need to be calculated based on the following equations:
ee ¼ Ec
4.1.2 Flexural Design
A chapter for the flexural design of UHPFRC structures is
not included in the AFGC-SETRA recommendations.
However, based on the suggested material models in
Fig. 14, the design of a UHPFRC element subjected to
bending can be performed by using the sectional analysis
(Fujikake et al. 2006a; Yang et al. 2011; Ferrier et al. 2015;
Yoo and Yoon 2015; Yoo et al. 2016)
. A schematic
description of a multi-layered cross-section with strain and
stress distributions and an algorithm for sectional analysis
are shown in Fig. 15
(Yoo and Yoon 2015)
crosssection of the elements is first divided into a number of
layers along the height. After that, the compressive and
tensile stresses at each layer can be calculated by assuming
that the plane section remains a plane at a given curvature.
Then, the neutral axis depth can be calculated from the
force equilibrium condition. Lastly, the moment is
calculated. The calculation was repeated until the ultimate strain
of the steel rebar was reached.
4.1.3 Shear Design
Since the steel fibers can resist a portion of the stress at
shear cracks, they mentioned that stirrups may be used, but
the shear strength given by the fibers may make it possible to
dispense with the stirrups. The ultimate shear strength Vu is
Vu ¼ VRb þ Va þ Vf
where VRb is the term for the participation of the concrete, Va
is the term for the participation of the reinforcement, and Vf
is the term for the participation of the fibers.
The shear strength by reinforcement Va is given by:
Va ¼ 0:9d Astt cfes ðsin a þ cos aÞ
VRb ¼ cE cb
where d is the effective depth, At is the area of the stirrups, st
is the spacing of the stirrups, fe is the stress in the stirrups, cs
is the partial safety factor, and a is the inclination angle of
VRb is expressed by two equations for reinforced and
prestressed concrete, as follow:
1 0:21 pffiffiffiffiffi
k fck b0d ðfor reinforced concreteÞ ð12Þ
VRb ¼ cE cb
1 0:24 pffiffiffiffiffi
ðfor prestressed concreteÞ
where cE is the safety coefficient such that: cE 9 cb = 1.5,
b0 is the element width, rm is the mean stress in the total
section of concrete under the normal design force,
k = 1 ? 3rcm/ftj for compression, k = 1 - 0.7rtm/ftj for
tension, and z is the distance from the top fiber to the center
of prestressing strand.
Lastly, the shear strength of the fibers is calculated by
using the following equation:
where S is the area of the fiber effect (S = 0.9b0d or
b0d for rectangular or T-sections, and S = 0.8(0.9d)2 or
0.8z2 for circular sections), bu is the inclination angle
between a diagonal crack and the longitudinal direction of
the beam, K is the orientation coefficient for general
effects, r(w) is the experimental characteristic
postcracking stress for a crack width of w, and wu is the
ultimate crack width.
4.2 JSCE Recommendations
4.2.1 Material Models
In JSCE recommendations, UHPFRC is defined as the
concrete with fc0 C 150 MPa, fcrk C 4 MPa, and ftk C 5
MPa, where fc0 is the compressive strength, fcrk is the
cracking strength, and ftk is the tensile stress at crack width
of 0.5 mm. JSCE recommendations were proposed based on
Ductal (Orange et al. 1999), which is commercially
available UHPFRC with heat treatment and 2 vol.% of steel fibers
having df = 0.2 mm and Lf = 15 mm, and provided the
strength properties used for structural design as follows;
fc0 = 180 MPa, fcrk = 8 MPa, ftk = 8.8 MPa, and Ec = 50
GPa. Importantly, they suggested a bilinear stress–strain
curve for compression (Fig. 16a) and a bilinear
tensionsoftening curve (TSC) with ftk = 8.8 MPa, w1k = 0.5 mm,
and w2k = 4.3 mm for tension (Fig. 16b). In order to obtain
the structural safety, cc = 1.3 was proposed for the partial
safety factor. To take into account the suggested TSC for
tensile stress block in the cross-section, the crack opening
displacement should be transformed to a strain by using the
equivalent specific length, Leq, as follows
Leq ¼ 0:8h41
1:05 þ 6h=lch
where h is the overall depth of beam, lch is the characteristic
length (=GFEc/ft2k), and GF is the fracture energy.
By using the equivalent specific length, the tensile stress–
strain model is obtained by the following equations, based
on the TSC:
Determination of material properties
· Compressive stress-strain model
· Tensile stress-strain model
Assuming curvature at initial stage
Assuming neutral axis
Determining strain distribution by assuming
that plane section remains plane
Calculating stress distribution using stress-strain
models under compression and tension
Calculating sectional compressive and tensile forces
Checking force equilibrium condition
Checking εs < εu ?
Determining moment-curvature curve
Fig. 15 Sectional analysis; a schematic description of stress and strain distributions in cross-section, b algorithm for sectional
(Yoo and Yoon 2015)
where ecr is the factored elastic strain, e1 is the strain at the
end of the initial plateau, w1k is the crack opening
displacement for which a certain stress level is retained after the
first crack, e2 is the strain at zero tensile stress, and w2k is the
crack opening displacement at zero tensile stress.
Other material properties of the UHPFRC, proposed by
the JSCE recommendations, are as follows:
Poisson’s ratio: 0.2.
Coefficient of thermal expansion (910-6/ C): 13.5.
Thermal conductivity (kJ/mh C): 8.3.
Thermal diffusivity (910-3m2/h): 3.53.
Specific heat (kJ/kg C): 0.92.
Total shrinkage: 550 9 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
(Yang et al. 2011; Yoo et al. 2016)
, similar to the
case of the AFGC-SETRA recommendations.
4.2.3 Shear Design
For the shear design of the UHPFRC elements, the shear
resistance from the matrix and the included steel fibers is
required to be calculated. In accordance with the JSCE
recommendations, the total shear resistance can be
Vyd ¼ Vrpcd þ Vfd þ Vped
where Vyd is the total shear resistance of the reinforced
UHPFRC beams, Vrpcd is the shear resistance of the matrix
without fiber, Vfd is the shear resistance of the steel fibers,
and Vped is the shear resistance of the stirrups.
Herein, Vrpcd and Vped are obtained by using Eqs. (21) and
Vrpcd ¼ 0:18 fc0bwd cb ð21Þ
Vfd ¼ ðfvd=tan buÞbwz=cb
where bw is the web width of the beam, d is the effective
depth of the beam, cb is the strength reduction factor
(cb = 1.3), fvd is the design tensile strength perpendicular to
the diagonal crack (= ftk/cb), z is the distance between the
point-acting compressive force and the center tensile
reinforcement (=d/1.15), and bu is the inclination angle between
the diagonal crack and the longitudinal direction of the
beam. The inclination angle bu should be larger than 30 and
is calculated by bu = 1/2tan-1[2s/(rxu - ryu)] - b0, where
s is the average shear stress, rxu and ryu are the average
compressive stress in longitudinal and transverse directions,
and b0 is the inclination angle without axial force.
In addition, Vped can be calculated as follows
Vped ¼ Ped sin ap cb
where Ped is the effective tensile force of the prestressing
strand, ap is the inclination angle between the prestressing
strand and longitudinal axis of beam, and cb is the strength
reduction factor (cb = 1.1).
As was reported by
Yang et al. (2012)
, the shear design of
UHPFRC elements can be carried out by using both
AFGCSETRA and JSCE recommendations, which provide good
estimates with test data of I-shaped UHPFRC beams, as
shown in Fig. 9.
5. Field Applications of UHPFRC
Due to its excellent mechanical performance, UHPFRC
can lead to a reduction in the number of sections, the
elimination of passive reinforcements, and the possibility for
the design of structures that is not possible with ordinary
. For this reason, UHPFRC has
attracted much attention from engineers for field applications
from 1995 to 2010
(Voo et al. 2012)
. The representative
application examples in North America, Europe, and Asia
are briefly explained herein. The first structural application
of UHPFRC was the prestressed hybrid pedestrian bridge at
Sherbrooke in Canada, constructed in 1997
(Fig. 17a). This precast and prestressed pedestrian
bridge includes a post-tensioned open-web space UHPFRC
truss with 4 access spans made by conventional
high-performance concrete. A ribbed slab with a 30-mm thickness
was adopted, and a transverse prestressing was applied with
sheathed monostrands. The UHPFRC truss webs were
confined by steel tubes, and the structure was longitudinally
prestressed by both internal and external prestressing
strands. The total span length of the bridge was 60 m, and
the main span was assembled from six 10-m prefabricated
match-cast segments. The Bourg-le`s-Valence bridge was the
first UHPC road bridge in the world, built in France in 2001
(Hajar et al. 2004)
, as shown in Fig. 17b. The bridge was
built from five assembled p-shaped precast UHPFRC beams,
and the joints were made by in situ UHPFRC with internal
reinforcements. The bridge consisted of two isolated spans
with a length of approximately 20 m, and all the p-shaped
beams were prestressed without transverse passive
reinforcement. The Seonyu Footbridge, completed in 2002 in
Seoul, Korea, is currently the longest footbridge made by
UHPFRC (Ductal ) with a single span of 120 m and no
central support (Fig. 17c). It consists of a p-shaped arch
supporting a ribbed UHPFRC slab with a thickness of
30 mm, and transverse prestressing was provided by
sheathed monostrands. With equivalent load carrying
capacity and strength properties, the bridge needed only half
of the amount of materials required for conventional
(Voo et al. 2014)
. The first UHPFRC
highway bridge, built in Iowa, USA, is the Mars Hill bridge,
as shown in Fig. 17d. This is a simple single-span bridge
consisting of three precast and prestressed concrete beams
with a length of 33.5 m (modified 1.14-m-deep Iowa
bulbtee beams), and the cast-in-place concrete bridge deck was
topped. No stirrup was applied, and each beam included 47
low-relaxation prestressing strands with a diameter of
(Russell and Graybeal 2013)
Curved UHPFRC panels were applied to the building
named the Atrium, which was built in Victoria, BC, Canada
in 2013 (Fig. 18a)
. In this project, UHPFRC
was selected as an appropriate material for the curved panel
system, due to its ability to form monolithically tight radial
curves, and consequently, it could improve the energy
efficiency of the building by eliminating unattractive
seams and openings. The Community Center in Sedan,
France was built in 2008 with a double skin facade to
protect the glass fascia behind and to provide privacy
using light blue UHPFRC perforated panels (Fig. 18b).
They designed the UHPFRC panels to be 2 m by 4 m by
45 mm thick, to permit sunlight to stream through to the
interior spaces. The main reason for choosing UHPFRC
panels instead of traditional perforated panels, made of
metal, painted steel, cast aluminum, cast iron, and
stainless steel, was that UHPFRC is durable, and requires less
energy consumption for fabrication and for maintenance
(Henry et al. 2011)
Mega-architectural projects were carried out to build the
Stade Jean Bouin and the MuCEM in France (Fig. 19)
. In the case of the Stade Jean Bouin, the first
application of a precast UHPFRC lattice-style facade
system in the world was made. The 23,000 m2 envelope,
which contains a 12,000 m2 roof, was built from 3600
selfsupporting UHPFRC panels that are 8–9 m long by 2.5 m
wide, and 45 mm thick. The MuCEM consists of several
types of precast UHPFRC structural elements, such as
78-m-long footbridge, main cubic structures with a
15,000 m2 surface area, flooring, and lattice-style envelope
with a series of slender N- and Y-shaped columns. The
architect designed the MuCEM to carry the entire external
load by columns, and to satisfy the maximum requirements
regarding both seismic and fire resistance, as specified by
Fig. 18 Examples of UHPFRC applications in buildings.
This paper reviewed the current state-of-the-art for
structural performance, design recommendations, and
applications of UHPFRC. From the above literature review and
discussions, the following conclusions are drawn:
(1) The use of 2 % steel fibers resulted in a higher
postcracking 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.
(2) 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
(3) 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.
(4) 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.
(5) 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
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