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Engineering Structures 153 (2017) 290–301
Contents lists available at ScienceDirect
Engineering Structures
journal homepage: www.elsevier.com/locate/engstruct
Experimental study of cyclic behavior of high-strength reinforced
concrete columns with different transverse reinforcement detailing
configurations
Wen-Cheng Liao a, Wisena Perceka a,b,⇑, Michael Wang b
a
b
Department of Civil Engineering, National Taiwan University, No. 1, Sec. 4, Roosevelt Road, Taipei 10617, Taiwan
Center For Earthquake Engineering Research, National Taiwan University, No. 1, Sec. 4, Roosevelt Road, Taipei 10617, Taiwan
a r t i c l e
i n f o
Article history:
Received 29 May 2017
Revised 4 October 2017
Accepted 4 October 2017
Keywords:
Confinement
Conventional closed-hoops
Butt-welded hoops
Single closed-hoop
Cross-ties
High-strength concrete
High-strength steel
High axial compression loading
a b s t r a c t
Detailing of transverse reinforcement is essential to ensure ductile behavior of reinforced concrete (RC)
members, particularly for columns with large displacement demanding under high axial loading level.
Requirements of reinforcing details are even important for high strength concrete due to its brittleness
nature. This paper presents experimental study regarding cyclic behavior of high strength RC columns
with different transverse reinforcement detailing layouts. The three high strength RC columns consisted
of different transverse reinforcement detailing configurations, which were conventional closed-hoops,
butt-welded hoops and single closed-hoop with cross-ties, respectively. The columns were made of
high-strength concrete with compressive strength of 70 MPa and high-strength steel with yield stresses
of 685 MPa and 785 MPa for longitudinal and transverse reinforcement, respectively. Cyclic displacement
static tests subjected to high axial compression loading of 0.3Agf0 c were conducted to verify the adequacy
of transverse detailing. The performance of the column with conventional closed-hoops and that of the
column with butt-welded hoops, which were compliant of ACI 318-14, were almost identical and met
the performance criteria required by ACI 374. In contrast, the column with a single closed-hoop and
cross-tie that was designed based on ACI 318-11 did not perform well as a ductile RC column. These
results prove that the ACI 318-14 minimum requirements for confinement should be followed. In addition, butt-welded hoops are acceptable as a form of transverse reinforcement of a column, since no
fracture was observed at the welded location. The current flexural and shear design equations for the
RC columns are discussed and compared with the test results.
Ó 2017 Elsevier Ltd. All rights reserved.
1. Introduction
The Taiwan New Reinforced Concrete (New RC) research project
was launched in 2008. One of the main objectives of the New RC
research project is to apply high-strength steel and concrete to
RC columns in high-rise buildings [1]. The use of high-strength
concrete and steel has gained attention recently due to the requirement for more available floor area, which can be achieved by limiting the size of lower-story columns in high-rise buildings [2,3]. By
using high-strength concrete and steel, the column section size and
reinforcement congestion can be reduced. In Taiwan, advanced
technology has led to the development of high-strength concrete
with a compressive stress of 70 MPa or greater. Meanwhile,
⇑ Corresponding author at: Department of Civil Engineering, National Taiwan
University, No. 1, Sec. 4, Roosevelt Road, Taipei 10617, Taiwan.
E-mail address: d02521014@ntu.edu.tw (W. Perceka).
https://doi.org/10.1016/j.engstruct.2017.10.011
0141-0296/Ó 2017 Elsevier Ltd. All rights reserved.
deformed steel bars SD685 and SD785 with specified yield stresses
of 685 MPa and 785 MPa for longitudinal and transverse reinforcement, respectively, are commercially available [2,3]. Increasing the
compressive strength, however, renders the concrete more brittle,
and thus greater transverse reinforcement ratio, particularly for
confinement in a high-strength concrete column is necessary
[4,5]. Detailing of transverse reinforcement is essential to ensure
ductile behavior of RC members, particularly for columns with
large displacement demanding under high axial loading level.
Requirements of reinforcing details are even important for high
strength concrete due to its brittleness nature.
The ACI 318-11 requirements for confinement reinforcement,
however, do not specify the limit of concrete strength and axial
load [6]. The results of a review of column test data conducted
by Elwood et al. confirmed that the previous confinement provisions (ACI 318-11) should be changed to include axial load for
designing confinement reinforcement [7]. ACI 318-14 then
W.-C. Liao et al. / Engineering Structures 153 (2017) 290–301
291
Notations
Ag
As
Ash
bc, bw
Cc, C00 s
d
d0
d00
f0 c
Fs
F0 s
fy
fsh
fyt
h
k
Lc
kf
kn
Mm
area of the column section
total area of longitudinal reinforcement
total area of transverse reinforcement parallel to lateral
loading direction
side dimension of the column core section and side
dimension of the column section, respectively, perpendicular to lateral loading direction
compressive force in concrete and compressive force in
longitudinal steel reinforcement, respectively, in cross
section of the column
effective depth of the column section (measured from
extreme-compression fiber to centroid of extremelayer of tensile steel reinforcement)
distance from extreme-compression fiber to centroid of
steel reinforcement in cross section of the column
distance from extreme-compression fiber to centroid of
steel reinforcement closest to extreme-compression
fiber
specified compressive strength of the concrete material
tension force in the extreme-layer of tensile longitudinal steel reinforcement in cross section of the column
compressive or tension force in the longitudinal steel
reinforcement in cross section of the column
yield stress of longitudinal reinforcement
stress of transverse reinforcement used to design shear
strength (limited to not exceed 600 MPa)
stress of transverse reinforcement used to design confinement (limited to not exceed 700 MPa)
height of the column section
factor relating the concrete or transverse reinforcement
capacity to the displacement ductility [Eqs. 12 and 12(a)]
clear height of column specimen
concrete strength factor for designing confinement
reinforcement [Eqs. 8 and 8(a)]
confinement effectiveness factor for designing confinement reinforcement [Eqs. 8 and 8(b)]
factored moment modified to account for effect of axial
compression load [Eqs. 10 and 10(a)]
adopted the results of the study conducted by Elwood et al. to
develop the new confinement provision, in which the axial load
and high-strength concrete factors are directly accounted for [8].
The new confinement provision also considers the number of longitudinal bars around the perimeter of the column core that are
laterally supported by the corner of closed-hoops or cross-ties with
a 135-degree hook [9]. A cross-tie is a transverse reinforcement bar
with a 135-degree hook at one of its edges and a 90-degree hook at
another edge. As is commonly known, the transverse reinforcement detailing may be a conventional closed-hoops model with a
seismic hook at one corner or a single closed-hoop model with
cross-ties [6]. Both ACI 318-11 and ACI 318-14 limit the yield stress
of transverse reinforcement as 420 MPa for designing shear
strength, while the use of yield stress of 700 MPa is permitted for
designing confinement reinforcement. Those codes also specify
not to use butt-welded hoops for ductile RC columns, since it can
lead to local embrittlement of steel [6,8,10]. However, there are
several advantages of using butt-welded hoops, such as the elimination of the possibility of anchor pull-out failure, and the reduction of reinforcement congestion [11].
It is worth mentioning that the amount and detailing of transverse reinforcement is critical to assure ductile behavior of RC columns under cyclic and high axial compression loading. Therefore,
the cyclic behavior of high-strength RC columns using different
Mn
Mpr
Mu
nl
Pn
Po
Pu
Vc
Vn
Vs
Vn,f, Vpr
a
a1
b1
k
qw
X1
nominal moment about an axis perpendicular to lateral
loading direction (computed about the centroid of
rectangular section) [Eq. 4]
maximum moment probable about an axis perpendicular to lateral loading direction
moment corresponding to applied axial compression
load
number of longitudinal bars around the perimeter of the
column core that are laterally supported by the corner
of closed-hoops or conventional cross-ties with a 135degree hook
nominal axial compression force in column section at
given eccentricity [Eq. 3]
nominal strength of reinforced concrete column at zero
eccentricity [Eq. 5]
applied axial load
shear strength provided by concrete [Eqs. 9,10,11,12,
and 13]
nominal shear strength [Eqs. 9,10,11,12, and 13]
shear strength provided by transverse reinforcement
[Eqs. 9,10,11,12, and 13]
shear force corresponding to the flexural strength of the
column section and shear force corresponding to the
maximum moment probable, respectively
reduction factor [Eqs. 13 and 13(a)]
factor relating the magnitude of uniform stress in an
equivalent rectangular compressive stress block to the
specified compressive strength of the concrete [Eq. 1]
ratio of the depth of the equivalent rectangular compressive stress block to the neutral axis [Eq. 2]
factor reflecting the reduced mechanical properties of
lightweight concrete (1 for normal concrete)
ratio of total area of tension steel reinforcement to bw.d
ratio of the mean concrete compressive stress corresponding to the maximum axial load resisted by a concentrically loaded column to the specified compressive
strength of concrete
transverse reinforcement detailing models is investigated in this
study. The high-strength RC columns with high-strength steel were
prepared and tested. All specimens had the same longitudinal bar
configuration, while the transverse reinforcement spacing of all
columns satisfied the minimum requirements of ACI 318-14 for
transverse reinforcement spacing.
2. Research significance
This study investigates the lateral strength and behavior of New
RC columns that use different transverse reinforcement detailing
configurations under double-curvature cyclic and high axial compression loading. The three high strength RC column specimens
consisted of different transverse reinforcement detailing configurations, which were conventional closed-hoops, butt-welded
hoops and single closed-hoop with cross-ties, respectively. The
applied axial compression ratio was 0.3. The specified concrete
compressive strength, yield stress of the longitudinal reinforcing
bar, and yield stress of the transverse reinforcing bar were
70 MPa, 685 MPa, and 785 MPa, respectively. This study is
significant because the tests also simulated the behavior of columns in the lower stories of high-rise buildings, in which a column
is subjected to high axial compression loading. In addition,
research on high-strength RC columns using different transverse
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W.-C. Liao et al. / Engineering Structures 153 (2017) 290–301
reinforcement detailing layouts under cyclic and high axial compression loading is still limited. The designed strengths (flexural
and shear) of three specimens, which were determined using the
current flexural and shear design equations, are compared with
the test results, and further analysis and application of design
equations for New RC columns are clearly discussed accordingly.
3. Experimental program
Three large-scale columns with a square cross section of 60 cm
60 cm and a clear height of 180 cm were prepared and tested
under double-curvature cyclic and high axial compression loading.
Detailed information about the experimental procedure is presented in the next five sub-chapters.
3.1. Material
The cementitious materials used in this study were ASTM Type I
Portland Cement and ground granulated blast-furnace slag (GGBS).
The maximum size of the coarse aggregate was 19 mm, and the
fine aggregate used was local natural sand. A superplasticizer
(SP) was also used to improve the workability of high-strength
concrete in the fresh state. The design concrete compressive
strengths were 70 MPa for all specimens. The concrete compressive strength of each column specimen was obtained by taking
the average compressive strength of six 100 mm 200 mm cylinders; the compressive strength of the concrete cylinders was measured by the compression tests based on the ASTM C-39 procedure
[12]. Two different sizes of reinforcing bars were used: D29 (6.29
cm2) and D13 (1.27 cm2) with specified yield stresses of 685 MPa
and 785 MPa, respectively. At least three reinforcing bars from
each specified diameter were subjected to a direct tensile test,
and a baseline tensile test determined the yield strength, ultimate
strength and bar strain at the yield stress for the control reinforcing
bar according to ASTM E8/E8M and ASTM A370 [13,14].
The concrete mix proportion is summarized in Table 1. The
average slump flow diameter of fresh concrete was 60 cm. The
label of each specimen begins with CF-C, denoting the concrete column having flexural failure, since the columns were designed to
have flexural failure. The symbols PT, W, and FT refer to the types
of configuration of transverse reinforcement. PT is the specimen
with conventional closed-hoops with a 135-degree hook at one
Table 1
Concrete mix proportion (kg/m3).
Cement
Slag
Coarse aggregate
Fine aggregate
Water
SP
292
238
859
816
170
5.57
Fig. 2. Specimen elevation.
of the corners of each hoop, W is the specimen consisting of
butt-welded hoops, and FT is the specimen with a single closedhoop and cross-ties. The value of 0.3 denotes the axial compression
ratio.
3.2. Specimen design
The details of the specimen design are shown in Figs. 1 and 2.
Information regarding the specified and actual concrete strength
and steel reinforcement strength is summarized in Table 2.
Detailed information about the column design parameters is listed
in Table 3. The concrete cylinders and their corresponding specimens were tested on the same day. Fig. 3 shows the compression
test setup and typical failure of a concrete cylinder, while Fig. 4
shows the failure of welded steel bars due to tensile testing.
Fig. 4 shows that the fracture location on the three steel bars
was far away from the welded location. The current design equations for RC column were used to predict the strengths of three
specimens.
Fig. 1. Square column with different detailing configurations of transverse reinforcement: (a) conventional closed hoops; (b) butt welded-hoops; (c) closed hoop with
cross-ties.
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W.-C. Liao et al. / Engineering Structures 153 (2017) 290–301
Table 2
Specified and actual values of concrete compressive strength, and yield stress of longitudinal and transverse reinforcements.
Specimen ID
Comp. Strength test
(f0 c,specified = 70 MPa)
CF-C-PT-0.3
CF-C-W-0.3
CF-C-FT-0.3
73
72
67
Longitudinal reinforcement SD685
(fy,spec = 685 MPa, fu,spec = 860 MPa)
Size
D29 (#9)
fy,test
698
Transverse reinforcement SD785
(fy,spec = 785 MPa, fu,spec = 930 MPa)
fu,test
920
Size
fy,test
fu,test
D13 (#4)
886
868
886
1095
1104
1095
Table 3
Column parameters.
Specimen ID
CF-C-PT-0.3
CF-C-W-0.3
CF-C-FT-0.3
Longitudinal reinforcement SD685
Transverse reinforcement SD785
Axial load ratio
n-size
ql (%)
Size, spacing (mm)
qt (%)
P/Agf0 c
16D29
2.87
D13(#4), 100
0.84
0.30
Fig. 3. Compressive test setup and failure of a concrete cylinder.
Fig. 4. Fracture locations of welded steel bars after testing.
3.2.1. Nominal flexural and axial strength
A stress–strain model and a equivalent rectangular compressive
stress block have been proposed for flexural analysis and design of
a RC section [15]. The strain distribution, the rectangular stress
block, and the internal forces in the analysis of nominal flexural
strength for the high-strength RC column are presented in Fig. 5.
The parameters a1 and b1 influence the width and height of the
stress block, respectively. The ACI ITG-4.3R-07 code [16] has recommended both parameters a1 and b1 , as shown in Eqs. (1) and
(2), respectively. The ultimate compressive strain at the extreme
fiber is defined as 0.003 [15,16]. The internal axial force and
moment on the column section can be calculated using Eqs. (3)
and (4). Eq. (5) gives the nominal axial strength for high-strength
RC columns as recommended by ACI ITG-4.3R-07. The ratio of
Fig. 5. Strain distribution, equivalent rectangular compressive stress block, and internal forces in analysis of flexural strength for a RC column section.
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W.-C. Liao et al. / Engineering Structures 153 (2017) 290–301
Axial force ratio, P /A .f'
g c
1.2
Interaction diagram
Design point (0.138; 0.3)
0.9
0.6
0.3
0.138
0
0
-0.3
0.04
0.08
0.12
Moment ratio, M/b .h 2
c c .f'c
0.16
-0.6
P/A .f'
g c
Fig. 6. Column interaction diagram with axial compression ratio of 0.3.
(a)
the mean concrete compressive stress corresponding to the maximum axial load resisted by a concentrically loaded column, to the
specified compressive strength of concrete, X1, shall be taken as
0.85 for concrete strengths up to 55 MPa, and shall be reduced continuously at a rate of 0.002 for each MPa of strength for concrete
strength beyond 55 MPa. The factor X1, however, cannot be less
than 0.7 [16]. Fig. 6 shows the dimensionless column interaction
curve. The flexural strength of a column is defined by plotting
the designed axial load on the column interaction curve.
0
For a1 : 0:70 6 a1 ¼ 0:85 0:0022ðf c 55Þ 6 0:85
0
ð2Þ
P ¼ 0 ) Pn Cc C00s þ F0s þ Fs ¼ 0
ð3Þ
h
h
h 1
00
C00s
d CC
b1 c
Mh ¼ 0 ) Mn Fs d 2
2
2
2 2
ð4Þ
For b1 : 0:65 6 b1 ¼ 0:85 0:0073ðf c 27:5Þ 6 0:85
X
X
ð1Þ
0
Po ¼ X1 f C ðAg As Þ þ f y As
ð5Þ
3.2.2. Confinement
In accordance with ACI 318-11, the amount of confinement
shall be the greater of Eqs. (6) and (7). As mentioned, ACI 318-11
does not specify the limit of concrete strength and axial load in
designing confinement reinforcement [6]. The new confinement
provision given in ACI 318-14 directly considers the axial compression, high-strength concrete, and the number of longitudinal bars
around the perimeter of the column core that are laterally supported by the corner of closed-hoops or cross-ties with seismic
hooks [8]. Therefore, the greatest of Eqs. (6)–(8) shall be taken in
determining the amount of confinement for a column with an axial
compression load ratio greater than 0.3, or with a concrete compressive strength of 70 MPa or greater. Fig. 7 shows the required
confining reinforcement ratio using ACI 318-11 and ACI 318-14
for the square column with a sectional dimension of 60 cm 60
cm and a longitudinal reinforcement ratio of 2.87% (16D29). It shall
be noted that there were 4 longitudinal bars that were not
supported by crossties (Fig. 1). By observing Fig. 7, it can be recognized that there is a significant difference between ACI 318-11 and
ACI 318-14 in designing confinement reinforcement for high
strength RC column under high axial compression loading.
Moreover, an examination of confinement provisions for specimen
CF-C-FT-0.3 (Fig. 7(b)) shows that the ACI 318-11 minimum requirements for confinement are questionable, particularly for highstrength concrete and steel in a column under high axial compression loading.
P/A .f'
g c
(b)
Fig. 7. Confining reinforcement ratio requirement of ACI 318-11 and ACI 318-14 for
specimens: (a) CF-C-PT-0.3 and CF-C-W-0.3, (b) CF-C-FT-0.3.
0
Ash
Ag
f
¼ 0:3
1 c
s bc
Ach
f yt
ð6Þ
0
Ash
f
¼ 0:09 c
s bc
f yt
ð7Þ
Ash
Pu
¼ 0:2kf kn
s bc
f yt Ach
ð8Þ
0
with : kf ¼
kn ¼
fc
þ 0:6 P 1:0
175
nl
ðnl 2Þ
ð8aÞ
ð8bÞ
3.2.3. Nominal shear strength
The shear strength equations provided in ACI 318-11 [6], proposed by Sezen [17], and proposed by Ou and Kurniawan [3] were
used to calculate the shear strength of the column specimens. The
use of a yield stress limit of 420 MPa in calculating shear strength
provided by transverse reinforcement is clearly specified in ACI
318-11 [6]. As reported by Alrasyid, a stress limit of 600 MPa is
acceptable to calculate the shear strength provided by highstrength transverse reinforcement (steel bar with a specified yield
stress of 785 MPa) [18]. The following equations present the nominal shear strength, Vn, used in this study.
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W.-C. Liao et al. / Engineering Structures 153 (2017) 290–301
ACI 318-11 [6]:
Simple equation:
with : a ¼
qffiffiffiffi
Pu
Av f sh d
0
k f c bw d þ
Vn ¼ Vc þ Vs ¼ 0:17 1 þ
s
14Ag
ðNÞ
ð9Þ
The detailed equation is taken as the smaller of Eqs. (10)
and (11):
Vn ¼ Vc þ Vs ¼
qffiffiffiffi
Vu d
Av f sh d
0
bw d þ
0:16k f c þ 17qw
Mm
s
4h d
P0
with : Mm ¼ Mu Pu
8
sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi
qffiffiffiffi
0:29Pu Av f sh d
0
þ
Vn ¼ Vc þ Vs ¼ 0:29k f c bw d 1 þ
s
Ag
ðNÞ
ð10Þ
ð10aÞ
ðNÞ
ð11Þ
sffiffiffiffiffiffiffiffiffi!
Pu
1 0:85
0 ;
Ag f c
06
Pu
0 6 0:6
Ag f c
ð13aÞ
To avoid shear failure, the designed shear strength must be
greater than the shear force corresponding to the maximum
moment probable, Mpr, at both ends of the column, Vpr. The total
Mpr at both ends of a column divided by the clear height of the column, Lc, is equal to the shear force corresponding to the maximum
moment probable, Vpr, (Vpr = 2Mpr/Lc). The Mpr occurs due to the
strain-hardening behavior of longitudinal reinforcement. In
accordance with the current concrete code [6,8], the yield stress
of tensile reinforcement in Mpr calculation shall be multiplied by
1.25; for the New RC columns, the Mpr can be approximated as
1.3Mn [19]. The designed strengths of the specimens are listed in
Table 4, in which the specified material strengths are used for all
calculations.
Sezen [17]:
0 qffiffiffiffi
1
0 vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi
u
0:5
f
2P
Av f sh d
cu
B
u C
Vn ¼ Vc þ Vs ¼ k@
t1 þ qffiffiffiffi A0:8Ag þ k
0
a=d
s
f c Ag
3.3. Test setup
ðNÞ
ð12Þ
k ¼ 1 for displacement ductility;
with :
k ¼ 0:7 for displacement ductility;
lD 6 2
lD P 6
ð12aÞ
Ou and Kurniawan [3]:
vffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi
qffiffiffiffi
u
2Pu
Av f sh d
u
0
þ
Vn ¼ Vc þ Vs ¼ 0:29ka f c bw dt1 þ qffiffiffiffi
0
s
a f c bc d
ðNÞ
ð13Þ
The Multi-Axial Testing System (MATS) at the National Center
for Research on Earthquake Engineering (NCREE) Taiwan was used
to test the column specimens. The maximum axial and lateral loads
that can be applied by MATS are 60 MN and 7 MN, respectively.
The lateral load was applied by hydraulic actuators placed at the
bottom of the MATS (Fig. 8(a)) using displacement control loading
history (Fig. 8(b)). The test was terminated when a significant loss
in axial loading capacity of the column was observed. Significant
loss in axial loading capacity was determined to have occurred
when there was a rapid reduction in axial loading capacity or the
reported displacement from the actuators was suddenly greater
than or equal to 20 mm. The tests were conducted under high axial
compression, resulting in a noticeable P–D effect, and consequently, correction was required. Correction methodology is
Table 4
Designed strengths.
Specimen ID
Nominal Shear Strength, Vn (kN)
Flexural strength
CF-C-PT-0.3
CF-C-W-0.3
CF-C-FT-0.3
Shear force due to Mn
Shear force due to Mpr
ACI 318-11
Mn (kNm)
Mpr (kNm)
Vnf (kN)
Vpr (kN)
Simple
Detail
Sezen
Ou and Kurniawan
2131.33
2770.73
2368.14
3078.60
2272.10
3201.30
3372.02
3032
Fig. 8. (a) Test setup on MATS and (b) loading protocol.
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W.-C. Liao et al. / Engineering Structures 153 (2017) 290–301
Fig. 9. Methodology of P–D effect correction.
Fig. 10. Hysteretic behavior and envelope response for specimens: CF-C-PT-0.3 ((a) and (b)), CF-C-W-0.3 ((c) and (d)), CF-C-FT-0.3 ((e) and (f)).
W.-C. Liao et al. / Engineering Structures 153 (2017) 290–301
detailed in Fig. 9, and calibrated complete hysteresis loops for the
column specimens can be modified accordingly.
4. Test results
Test results were examined based on visual observations and
the data that were recorded by the MATS and strain gauges. As
reported by Elwood et al. [7], a sufficient column should satisfy
the seismic performance criteria, namely, the ultimate drift shall
be at least 3% with lateral strength corresponding to ultimate drift
of 80% maximum lateral strength. Meanwhile, ACI 374-1-0.5 specifies that the acceptance performance criteria for a RC column is
when a column reaches the ultimate drift ratio of 0.035 and the
lateral strength corresponding to ultimate drift of 75% maximum
lateral strength [20]. The following sections present the test results
and discussions.
4.1. Lateral strength–displacement relationships
The lateral strength–drift ratio relationship and the envelope
curve for each column are shown in Fig. 10, in which the P–D effect
in every case has been eliminated. The important points for observation are the following: the yield point of longitudinal and
transverse reinforcements, the idealized yield point, the peak lateral strength, the ultimate lateral strength and corresponding drift
ratio, and ductility. The appearance of the columns post-testing is
presented in Fig. 11. The following is method to obtain the idealized yield point: a straight line from zero was draw to intersect
the lateral strength-drift curve at 70% of the maximum lateral
strength. That straight line was then extended to the intersection
(a) CF-C-PT-0.3
(b) CF-C-W-0.3
(c) CF-C-FT-0.3
297
with a horizontal line corresponding to the maximum lateral load,
and then projected onto the horizontal axis to obtain the yield
displacement (idealized Dy) [17].
Specimen CF-C-PT-0.3 had conventional closed-hoops with a
135-degree hook at one of the corners of each hoop and had longitudinal and transverse reinforcement ratios of 2.87% and 0.84%,
respectively. Four longitudinal bars around the perimeter of column core were not laterally supported by cross-ties with seismic
hooks. The applied axial compression load was 7861 kN (axial load
ratio of 0.3). The specimen behaved elastic until the drift ratio
reached 0.83%. The lateral stiffness degraded once drift ratios were
greater than 0.83%, and the peak lateral strength of 2888 kN was
reached at a drift ratio of 1.82%. The longitudinal and transverse
reinforcement yielded at a drift ratio of 2% and 1.5%, respectively.
The ultimate lateral strength was 2311 kN and occurred at a drift
ratio of 3.17%. The specimen CF-C-PT-0.3 failed at a drift ratio of
4.0% when it lost its axial load-carrying capacity. Both longitudinal
reinforcement buckling and anchor pullout failure of the transverse reinforcement were noticeable at the end of the test.
The transverse reinforcement detailing of specimen CF-C-W-0.3
was similar to that of specimen CF-C-PT-0.3. As previously mentioned, the use of butt-welded hoops in specimen CF-C-W-0.3 is
to reduce the reinforcement congestion in the plastic hinge region
and to avoid the possibility of pullout failure of an anchor. The
applied axial compression load was 7805 kN. Specimen CF-C-W0.3 behaved elastic at drift ratios up to 0.73%. The degradation of
lateral stiffness started once the drift ratio exceeded 0.73%, and a
peak lateral strength of 2715 kN was reached at a drift ratio of
1.36%. The yield points of longitudinal and transverse reinforcement were found at a drift ratio of 1.5%. The ultimate lateral
strength of 2172 kN corresponded with a drift ratio of 3.05%. The
test was stopped at a drift ratio of 5.12% when the axial loadcarrying capacity of specimen CF-C-W-0.3 was lost. At the end of
the test, longitudinal reinforcement buckling was found, but no
fracture occurred at the welded locations on the transverse
reinforcement.
The transverse reinforcement for specimen CF-C-FT-0.3 was
based on the provisions given in ACI 318-11. The applied axial
compression load was 7208 kN. As has been observed in specimen
CF-C-PT-0.3 or CF-C-W-0.3, the elastic behavior of specimen CF-CFT-0.3 was found at drift ratios in the range of 0–0.8%. However, a
peak lateral strength of 2654 kN was reached at a drift ratio of
1.17%. It was found that the longitudinal and transverse reinforcements yielded at a drift ratio of 1.5%. Finally, an ultimate lateral
strength of 2123 kN occurred at a drift ratio of 2.31%, and the specimen lost its axial load-carrying capacity at a drift ratio of 3.00%.
Longitudinal reinforcement buckling and anchor pullout failure
of the transverse reinforcement were clearly identified. The idealized yield point, the peak lateral strength, the ultimate lateral
strength and corresponding drift ratio, and ductility for each
column specimen are summarized in Table 5.
4.2. Crack patterns
(d) CF-C-PT-0.3
(e) CF-C-W-0.3
(f) CF-C-FT-0.3
Fig. 11. The final appearance of specimens at the end of the test ((a) to (c)) and the
appearance of concrete cores ((d) to (f)).
Flexural cracking near the top of the column occurred on specimens CF-C-PT-0.3, CF-C-W-0.3, and CF-C-FT-0.3 at drift ratios of
0.25%, 0.25%, and 0.375%, respectively. The crack width increased
as drift increased, and Table 6 shows the crack width occurring
at each drift ratio. In general, the shear cracking appeared as diagonal cracking developed from flexural cracking at the top and bottom of the column. The first shear cracking occurred in specimens
CF-C-PT-0.3 and CF-C-W-0.3 at drift ratios of 0.75% and 0.375%,
respectively, while the first shear cracking in specimen CF-C-FT0.3 appeared at a drift ratio of 0.25% (before flexural cracking
appeared). For specimen CF-C-PT-0.3, the concrete shell started
crushing at the top and bottom of the column at a drift ratio of
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W.-C. Liao et al. / Engineering Structures 153 (2017) 290–301
Table 5
Results of force–displacement relationships.
Specimen ID
Idealized yield drift %
Peak lateral strength (kN)
Ultimate lateral strength (kN)
Ultimate drift %
Ductility
CF-C-PT-0.3
CF-C-W-0.3
CF-C-FT-0.3
0.57
0.57
0.56
2888
2715
2654
2311
2172
2132
3.17
3.05
2.31
5.56
5.35
4.12
Table 6
Maximum crack widths of specimens.
Drift
(%)
CF-C-PT-0.3
Flexural crack width
(mm)
Shear crack width
(mm)
Flexural crack width
(mm)
Shear crack width
(mm)
Flexural crack width
(mm)
Shear crack width
(mm)
0
0.25
0.375
0.50
0.75
1.0
1.5
2.0
N.A
0.05
0.08
0.10
0.30
0.60
1.20
2.50
N.A
N.A
N.A
N.A
0.20
0.30
0.60
2.20
N.A
0.10
0.25
0.25
0.60
0.65
0.90
3.00
N.A
N.A
0.05
0.05
0.30
0.45
0.65
4.00
N.A
N.A
0.10
0.25
0.25
0.40
1.20
2.00
N.A
0.05
0.05
0.10
0.25
0.35
0.70
5.00
Crack angle (Degree)
60
50
CF-C-W-0.3
CF-C-FT-0.3
CF-C-PT-0.3
CF-C-W-0.3
CF-C-FT-0.3
40
30
20
10
0.25
0.5
0.75
1
1.25
1.5
1.75
2
Drift ratio (%)
Fig. 12. Crack angle–drift ratio relationships.
1.5% and the concrete shell gradually spalled off as the drift ratio
increased. In addition, the diagonal cracks at the middle of the column started appearing at a drift ratio of 0.75%, and as the drift ratio
further increased, several diagonal cracks appeared at the top and
bottom of the column. More concrete shell at the bottom of the column spalled off at a drift ratio of 2%.
For specimen CF-C-W-0.3, the concrete shell started crushing at
a drift ratio of 1%. The diagonal cracks were noticeable once the
drift ratio reached 1.5%. More concrete shell crushed and more
diagonal cracks appeared at the middle of the column as the drift
ratio increased. The specimen CF-C-FT-0.3 showed that the concrete shell at the top and bottom of column crushed at a drift ratio
of 1%. Diagonal cracks were seen once the drift ratio reached 2%,
and more diagonal cracks appeared as the drift ratio increased.
The crack width that occurred at each drift ratio is listed in Table 6,
and the crack angle–drift relationships are shown in Fig. 12.
length of longitudinal reinforcement in specimen CF-C-PT-0.3 is
shorter than that of in specimen CF-C-W-0.3. This is because the
number of layers of conventional closed-hoops is slightly greater
than that of butt-welded hoops. A small clear spacing of transverse
reinforcement improves the effective confinement pressure in the
column core. More information regarding confinement mechanisms can be found in existing literature [21]. Regardless of the
performance aspects of specimen CF-C-W-0.3, butt-welded hoops
can be an alternative transverse reinforcement for high-strength
concrete columns in seismic regions. As presented in Fig. 11(e),
no fractures on the welded location were found. The test results
of specimen CF-C-FT-0.3 prove that the confinement provisions
provided in ACI 318-14 should be followed. Although, the clear
spacing of transverse reinforcement and transverse reinforcement
ratio of specimens CF-C-PT-0.3 and CF-FT-FT-0.3 were identical,
the pullout failures of ties with a 90-degree hook were found in
specimen CF-C-FT-0.3. According to ACI 318-2014, the cross-ties
with a 90-degree hook are not as effective as cross-ties with a
135-degree hook [8]. The test results prove that the number of longitudinal bars laterally supported by crossties with seismic hooks
influence the performance of column. Specimens CF-C-PT-03 and
CF-C-W-0.3 having 16 longitudinal bars with 4 longitudinal bars
laterally un-supported still met the performance criteria required
by ACI 374, while specimen CF-C-FT-0.3 did not perform well as
a ductile RC column. Specimen CF-C-FT-0.3 might achieve better
performance if the provided transverse reinforcement ratio was
larger than 0.84% or the provided crossties had a bend of 135°.
The smallest and greatest flexural cracks were found on specimen CF-C-FT-0.3 and CF-C-W-0.3, respectively. By contrast, the
greatest shear crack was found on CF-C-FT-0.3. It can be seen in
Fig. 12 that the trends of the crack angle–drift curves are similar.
The crack angle decreases as drift ratio increases.
5. Comparison between measured and designed strengths
4.3. Discussion
The results show that behavior of specimen CF-C-PT-0.3 was
more desirable than that of specimen CF-C-W-0.3. The peak lateral
strength, ultimate lateral strength and corresponding drift ratio,
and ductility of specimen CF-C-PT-0.3 were greater than those of
specimen CF-C-W-0.3, although the two columns had the same
transverse reinforcement ratio. It can be known that the unbraced
The final step of this study is to evaluate the current strength
design equations for RC column. The peak lateral strength of each
column specimen is verified using the ACI 318-11 simplified shear
strength equation (Vn,1 in Table 7, by Eq. (9)), ACI 318-11 detailed
equations (Vn,2 in Table 7, by Eqs. (10) and (11)), Sezen’s shear
strength equation (Vn,3 and Vn,4 in Table 7, by Eq. 12(a) for lD 2
and lD > 2), shear strength proposed by Ou and Kurniawan
(Vn,5 in Table 7, by Eq. (13)), and shear force corresponding to
299
W.-C. Liao et al. / Engineering Structures 153 (2017) 290–301
Table 7
Comparison between peak lateral and designed strengths.
Specimen
Vtest/Vn,1
Vtest/Vn,2
Vtest/Vn,3
Vtest/Vn,4
Vtest/Vn,5
Mtest/Mn
Vtest/Vn,f
Vtest/Vpr
CF-C-PT-0.3
CF-C-W-0.3
CF-C-FT-0.3
1.27
1.19
1.17
0.90
0.85
0.83
0.86
0.81
0.79
0.94
0.83
0.79
0.95
0.90
0.88
1.22
1.15
1.12
1.22
1.15
1.12
0.94
0.88
0.86
(a)
(b)
(c)
Fig. 13. The envelope measured and calculated shear strengths for specimens: (a) CF-C-PT-0.3, (b) CF-C-W-0.3, (c) CF-C-FT-0.3.
300
W.-C. Liao et al. / Engineering Structures 153 (2017) 290–301
the nominal flexural strength given by ACI ITG-4.3R-07 (Vn,f in
Table 7). Vn,4 was obtained by calculating k corresponding to the
ratio of drift at peak lateral load to idealized yield drift (D,peakload/
Dy). Fig. 13 shows the envelope measured and designed shear
strengths.
The ACI 318-11 simplified shear equation only predicts the
shear strength that is smaller than the peak lateral strength of
specimens obtained from test results. In addition, ACI 318-11 specified that the maximum yield stress of transverse reinforcement for
shear strength design is 420 MPa. Therefore, designing the shear
strength of the New RC columns using the ACI 381-11 simplified
shear equation may be too conservative. On the other hand, the
use of the ACI 318-11 detailed equations tends to overestimate
the nominal shear strength. Table 4 shows that nominal shear
strengths from the ACI 318-11 detailed equations were higher than
both the peak lateral force and the shear force corresponding to the
maximum probable flexural strength, although the yield stress
used in the calculation was 420 MPa. The shear force corresponding to the flexural strength of a column, which is based on ACI
ITG-4.3R-07, was in agreement with the ultimate shear force
(shear force corresponding to ultimate drift ratio). By observing
Mtest/Mn, it can be concluded that the use of the equivalent rectangular compressive stress block recommended in ACI ITG-4.3R-07 is
acceptable for determining the nominal flexural strength of the
New RC column. Furthermore, the measured lateral strengths are
smaller than the calculated nominal shear strengths based on the
shear strength equations proposed by Sezen (Eq. (12)) and Ou
and Kurniawan (Eq. (13)). This result occurs because the type of
failure in the three column specimens was flexural-shear failure.
The shear strength corresponding to the maximum moment
probable was close to the nominal shear strength proposed by
Sezen (Eq. (12)) and Ou and Kurniawan (Eq. (13)). It should be
noted that the shear strength corresponding to the maximum
probable moment tends to be rare occurrence in a high-strength
column under high axial compression loading. It is well known that
the higher axial compression loading can reduce the moment
capacity, as has been presented in the axial-moment interaction
diagram. The axial load increases the compression stress in compression zone so that the distance between the resultant compression force and the resultant tension force from tension steel
reinforcement (level arm) and the height of tension zone decrease.
Therefore, the stress of tension longitudinal reinforcement may be
equal to or smaller than its yield stress once the flexural capacity of
column is reached. In compression zone, the stress of compression
reinforcement tends to be far from yield since compression stress
acting on compression zone is received by concrete and compression reinforcement.
Meanwhile, the test results show that the yielding longitudinal
reinforcement almost corresponded with the peak lateral
strengths. The post yield lateral force might have increased if the
tensile stress in the longitudinal reinforcement increased to its
ultimate stress (1.25 times yield stress). Therefore, at high axial
loading level, this shows that applying factor 1.25 Vpr does not
improve column flexural strength and will overestimate the actual
maximum lateral strength corresponding to flexural strength.
The nominal shear strength equation proposed by Sezen
describes shear strength degradation well. It can be seen that the
curves of the shear strengths calculated using Sezen’s equation
with k larger than 1, and the experimental curves of the specimens
tend to have a similar trend.
6. Conclusions
Based on the experimental results and analysis that are presented in this paper, the following conclusions can be drawn:
1. The performance of specimen with butt-welded hoops (CFC-W-0.3) tended to be similar to that of specimen with conventional closed-hoops (CF-C-PT-0.3). Therefore, the use of
high strength butt-welded hoops (SD785) for confining reinforcement in the New RC columns is acceptable.
2. The ultimate drift ratio of a specimen having conventional
closed-hoop transverse reinforcement and cross-ties with
90-degree and 135-degree hooks (CF-C-FT-0.3) was only
2.31%; therefore, the ultimate drift ratio of CF-C-FT-0.3
did not reach the minimum criteria for a ductile RC
column.
3. The use of cross-ties with 90-degree hook increases the
effectiveness factor (kn). In accordance with new confinement provisions given in ACI 318-14, the improvement of
effectiveness factor (kn) increases the required confinement.
As shown in Fig. 7(b), the amount of confinement in CF-CFT-0.3 only satisfy the ACI 318-11 minimum confinement.
In addition, ACI 318-11 does not specify clearly the limits
of concrete strength and axial compression ratio for designing confinement reinforcement.
4. The examination of confinement equations and test results
prove that the confinement provisions given in ACI 318-14
should be followed.
5. The ACI 318-11 simplified shear strength equation only predicts shear strength before achieving the peak lateral force,
for all specimens. By contrast, the ACI 318-11 detailed equations may overestimate the shear strength of the New RC
columns.
6. Equivalent rectangular compressive stress block recommended in ACI ITG-4.3R-07 is a conservative method for predicting the nominal flexural strength of the New RC
columns.
7. The nominal shear strength proposed by Sezen (Eq. (12)) and
Ou and Kurniawan (Eq. (13)) were close to the shear
strength corresponding to the maximum probable flexural
strength, Vpr. Meanwhile, the peak lateral strengths from
the test the results were smaller than the nominal shear
strengths calculated using Eqs. (12) and (13) due to the three
column specimens experiencing flexural-shear failure.
8. The shear corresponding to the maximum probable
moment tends to be rare occurrence in columns under
high axial compression loading. At high axial loading level,
the axial load increases the compression stress in compression zone. This reduces the distance between the resultant
compression force and the resultant tension force from
tension steel reinforcement (level arm) and the height of
tension zone. The stress of tension longitudinal reinforcement can be either equal to or smaller than its yield stress
once the flexural strength is column is reached. In addition, the possibility of compression reinforcement reaching
its yield stress is low, since compression stress acting on
compression zone is received by concrete and compression
reinforcement.
According to the test results, the yielding longitudinal reinforcement almost corresponded with the peak lateral load.
The post yield lateral force might have increased if the stress
of the longitudinal reinforcement increased to its ultimate
stress (1.25 times yield stress). Therefore, at high axial loading level, applying factor 1.25 does not improve column flexural strength and will overestimate the actual maximum
lateral force corresponding to flexural strength.
9. The nominal shear strength equation proposed by Sezen
(Eq. (12)) describes shear strength degradation well. As
shown, the trend of experimental curves was similar to that
of the curves obtained using shear strengths calculated by
Sezen’s equation with k larger than 1.
W.-C. Liao et al. / Engineering Structures 153 (2017) 290–301
Acknowledgments
The research described in this paper was sponsored by the Ministry of Science and Technology (MOST) of Taiwan under Grant No.
105-2221-E-002-057-MY2 and 106R890803, and the National Centre for Research on Earthquake Engineering (NCREE) Taiwan is
gratefully acknowledged. The opinions expressed in this paper
are those of the authors and do not necessarily reflect the views
of the sponsor. The authors would like to thank Professor ShyhJiann Hwang from Department of Civil Engineering at National Taiwan University. The Doctoral scholarships for two academic years
(2016/2017 and 2017/2018) received by the second author are provided by Centre for Earthquake Engineering Research (CEER) at
National Taiwan University.
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