1 seismic behaviors of columns in ordinary and intermediate moment resisting concrete frames

Engineering Structures 27 (2005) 951–962
www.elsevier.com/locate/engstruct
Seismic behaviors of columns in ordinary and intermediate moment
resisting concrete frames
Sang Whan Han∗, N.Y. Jee
Department of Architectural Engineering, Hanyang University, Seoul 133-791, Republic of Korea
Received 30 March 2004; received in revised form 31 January 2005; accepted 31 January 2005
Available online 10 March 2005
Abstract
The objective of this study was to investigate the seismic behaviors of columns in Ordinary Moment Resisting Concrete Frames (OMRCF)
and Intermediate Moment Resisting Concrete Frames (IMRCF). For this purpose, two three-story OMRCF and IMRCF were designed
according to the minimum design and reinforcement detailing requirements specified in ACI 318-02. This study assumed that the building
was located in seismic zone 1, as classified by UBC. According to ACI 318-02 the reinforcement detailing requirements for OMRCF are
less stringent than those for either IMRCF or SMRCF (Special Moment Resisting Concrete Frames). Tests were carried out to evaluate the
seismic behaviors of OMRCF and IMRCF columns using 2/3 scale model columns. Each column was considered as consisting of an upper
part and lower part divided at the point of inflection. Quasi-static reversed cyclic loading was applied to the specimens with either constant
or varying axial forces. The test variables of this experimental study were the type of axial force (constant and varying, and low and high),
the existence of lap splices (with or without lap splice) and type of moment resisting concrete frame (OMRCF or IMRCF). It was observed
that all OMRCF and IMRCF column specimens had strength larger than that required by ACI 318, and they had drift capacities greater than
3.0% and 4.5%, respectively. However, the drift capacity of specimens varied with respect to the existence of lap splices and the spacing of
lateral reinforcement at column ends.
© 2005 Elsevier Ltd. All rights reserved.
Keywords: Seismic behavior; Columns; Reinforcement details; Seismic design
1. Introduction
Moment frames have been widely used for seismic
resisting systems due to their superior deformation and
energy dissipation capacities. A moment frame consists
of beams and columns, which are rigidly connected. The
components of a moment frame should resist both gravity
and lateral load. Lateral forces are distributed according to
the flexural rigidity of each component. ACI 318-02 [1]
states the design and reinforcement detailing requirements
for each type of moment frame and each earthquake
risk level. The type of moment frame should be selected
according to levels of seismic risk or seismic design
category. Seismic risk levels can be classified into low,
∗ Corresponding author. Tel.: +82 2 2290 1715; fax: +82 2 2291 1716.
E-mail address: swhan@hanyang.ac.kr (S.W. Han).
0141-0296/$ - see front matter © 2005 Elsevier Ltd. All rights reserved.
doi:10.1016/j.engstruct.2005.01.012
moderate and high according to seismic zones specified in
UBC [2]. Seismic design category is specified in NHERP
[3] and IBC [8].
ACI 318-02 classifies concrete moment frames into
three types: Ordinary Moment Resisting Concrete Frame
(OMRCF); Intermediate Moment Resisting Concrete Frame
(IMRCF); and Special Moment Resisting Concrete Frame
(SMRCF). Fig. 1 shows the minimum column reinforcement
details for the columns of OMRCF, IMRCF, and SMRCF
specified in ACI 318 and notes [1,4].
OMRCF and IMRCF are the most popular types of
moment frames in low to moderate seismic zones. The
design and reinforcement requirements for OMRCF and
IMRCF are less stringent than those for SMRCF. When
a large earthquake occurs, SMRCF is expected to have
superior ductility and provide superior energy dissipation
capacity. In current seismic design procedures, design base

952 S.W. Han, N.Y. Jee / Engineering Structures 27 (2005) 951–962
Fig. 1. Minimum reinforcement details for columns.
Notation
Ag gross area of column
fc specified compressive strength of concrete
fy specified yield strength of nonprestressed
reinforcement
Mmax maximum moment resistance of the specimen
MACI moment capacity calculated using ACI 318-99
procedures
VACI nominal shear strength according to ACI 318-
99
∆max maximum displacement
θmax maximum drift angle
shear force shall be calculated by elastic strength demand
divided by R factor (≥1). R factor accounts for ductility,
overstrength, redundancy, and damping inherent in a
structural system. Current design provisions assigned the
highest R factor to SMRCF, and the lowest R factor to
OMRCF because of its less stringent design requirement. It
should be noted that design base shear becomes larger with
decreasing R factor.
This study focused on the behaviors of first story columns
in moment resisting frames. For this purpose, three-story
OMRCF and IMRCF buildings were designed according
to ACI 318-02. Experimental test of columns in the first
story of those buildings was carried out to investigate the
seismic behavior of the columns. This study considered that
a column consisted of an upper part and lower part divided
at the point of inflection.
Eight 2/3 scale column specimens (2 × 2 × 2 = 8)
representing the lower and upper part (2) of the exterior and
interior columns (2) in the OMRCF and IMRCF (2) were
made. Lower column specimens had lap splices whereas
upper column specimens had continuous longitudinal
reinforcement. The most significant difference in OMRCF
and IMRCF columns is the spacing of the transverse
reinforcement at the end of the column. Quasi-static reversed
cyclic loading was applied to the specimens with either fixed
or varying axial forces.
The test variables in this study were type of axial force
(constant and varying, and low and high), existence of lap
splices (with or without lap splice) and type of moment
resisting concrete frames (OMRCF or IMRCF).
2. Previous studies
Han et al. [6,7] have evaluated the seismic performance
of a three-story OMRCF. In their studies maximum lateral
force was measured during the test and compared with the
design base shear force specified in current codes. Moreover,
this study carried out seismic performance evaluation of the
three-story OMRCF using a capacity spectrum method.
Lee and Woo [9] investigated the seismic performance
of low-rise reinforced concrete (RC) frames with non-
seismic detailing. They considered a bare frame and a
masonry infilled frame. An experiment was conducted

S.W. Han, N.Y. Jee / Engineering Structures 27 (2005) 951–962 953
using 1/5 scaled specimens having two bays and three
stories. They found that RC moment frames had a certain
seismic resistance even though they were designed without
considering seismic loads. Furthermore, masonry infill was
beneficial to the seismic performance of the frame.
Aycardi et al. [10] investigated the behavior of gravity
load-designed reinforced concrete column members under
simulated seismic loading. They tested four column
specimens governed by flexure under reversed cyclic loading
at increasing drift amplitudes. It was observed that all
column specimens were capable of sustaining at least two
cycles of loading at a 4% drift angle.
Lynn et al. [11] tested eight reinforced concrete column
specimens having typical details of those built before the
mid-1970s. This study observed that column specimens that
were governed by shear experienced gravity load failure
soon after loss of lateral force resistance. When flexure
was dominant in specimens, gravity load capacity was
maintained to large displacement.
3. Experimental program
3.1. Design of building frames
Typical three-story office buildings were designed.
Seismic resistance for these buildings was provided by
OMRCF and IMRCF. The dimension and design loads were
adopted from Han [6]. Fig. 2 shows the dimensions of the
building.
Specified compressive strength ( fc) of concrete was
assumed to be 23.5 MPa. Longitudinal reinforcement and
reinforcement for hoop and stirrup were assumed to have a
yield strength ( fy) of 392.3 MPa. Design loads for a typical
office building were used (5.2 kPa for dead load and 2.4 kPa
for live load). The section of all columns and beams in the
model frames were assumed to be 330 mm × 330 mm and
250 mm × 500 mm, respectively.
The buildings were assumed to be located in seismic
zone 1 as classified in UBC [2]. Member forces were
obtained using SAP 2000 (2000). Gravity loads governed the
member design. Since the seismic load was small, OMRCF
and IMRCF had the same member sizes as well as the
same amount of reinforcement except for the transverse
reinforcement in columns and beams (see Fig. 3). This
design was desirable for this study since this study attempted
to investigate the effect of different reinforcement details
of OMRCF and IMRCF columns on seismic behaviors of
those columns without interference of other factors such as
amount of longitudinal reinforcement and different member
sizes.
3.2. Column specimens
As shown in Fig. 2, only the columns in the first story
were evaluated since these columns should resist the largest
axial and lateral forces during an earthquake. The exterior
(a) Plan view.
(b) Elevation.
Fig. 2. Building layouts.
columns in the first story of the two model frames were
designed for an axial force of 644.3 kN and a bending
moment of 37 kN m, and the interior columns were designed
for an axial load of 1234.7 kN and a bending moment of
47.1 kN m. It is noted that those design forces contain load
factors.
All columns had a section of 330 mm × 330 mm
and contained four longitudinal reinforcements (D19
(19.1 mm diameter), fy = 392.3 MPa). Column
longitudinal reinforcement ratio (ρ) was 1.01%, which
slightly exceeded a minimum longitudinal reinforcing steel
of 1.0% (Section 10.9.1 in ACI 318-02).
Maximum shear force in the first story columns induced
by factored gravity loads was 31.4 kN. According to the
formula in Section 11.3.1.2 of the ACI 318-02, the concrete
shear strength of the first story columns (Vc) was calculated
as 73.5 kN. Minimum tie reinforcement (D10 with a
diameter of 9.53 mm) with spacing of 300 mm was placed
throughout the column in the OMRCF (Section 7.10.5.2
in ACI 318-02) and the first tie reinforcement was placed
150 mm from the slab or footing surface according

Fig. 3. Column specimens.
to the requirement of the first tie location specified in
Section 7.10.5.4 in ACI 318-02.
Fig. 3 shows that the first tie was placed 100 mm from the
slab in 2/3 scale specimens. The longitudinal reinforcement
details in the IMRCF columns and the OMRCF columns
were exactly the same. Tie spacing in the plastic hinging
region of the IMRCF columns, however, was half of the tie
spacing in the OMRCF columns. The yield strength of D10
was 392.3 MPa. A lap splice length of 525 mm was required
for tension according to Section 12.15 in ACI 318-02. Lap
splices in a column were placed just above the slabs (see
Figs. 1 and 3).
It is worthwhile to note that as shown in Fig. 1 the upper
part of columns has continuous longitudinal bars whereas
the lower part of columns has lap splices. Thus this study
made specimens representing the upper part and lower part
of columns (see Fig. 2). It was assumed that the inflection
point was located at the mid-height of the column during
earthquakes.
Fig. 3 shows the reinforcement details and dimensions
of the 2/3 scale column test specimens. Table 1 describes
the characteristics of the specimens. All reinforcements
were scaled by 2/3 for test specimens. A total of eight
specimens were made (upper and lower part (2) of interior
and exterior columns in OMRCF and IMRCF (2)). Lap
splice length (350 mm) was also simply scaled by 2/3 of
the original length (525 mm). No attempt was made to re-
calculate lap splice length for scaled reinforcement. Rebar
D13 (12.7 mm diameter) and D6 (6.35 mm diameter) were
used for longitudinal and transverse tie reinforcement to
represent 2/3 of D19 and D10 in the columns of model
frames.
The dimension of column base is shown in Fig. 3.
The longitudinal reinforcement of column specimens was
anchored to the column base satisfying development length
required by ACI 318-02 (chapter 12). In order to fix the
specimen to the strong floor, four high strength bolts having
a diameter of 70 mm were installed (see Fig. 4(a)). All bolts
are fully tightened to provide specimens with fixed base
conditions. In column base, sufficient closed type stirrups
using reinforcement of D13 (13 mm diameter) were placed
to resist the forces transferred from the specimen. No cracks
were detected in the column base after test (see Fig. 4(a)).
3.3. Material tests
Based on concrete trial mixes from various recipes for
attaining the 28-day target strength ( fc) of 23.5 MPa, a
design mix was determined. The maximum size of aggregate
for 2/3 scale model specimens was 25 mm. Cylinder tests
were performed. Each concrete cylinder was 200 mm in
height and 100 mm in diameter. Cylinder test results are
given in Table 2.
Longitudinal and transverse reinforcement in the column
specimens (2/3 scale) were D13 (13 mm diameter) and D6
(6 mm diameter), respectively. The results of material tests
are given in Table 3.
3.4. Loading and measurements
All eight-column specimens were laterally loaded at the
level of 1000 mm above the base (mid-height of a column)
(see Fig. 4). It was assumed that the inflection point of a
column is located at mid-height of the column.

Fig. 4. Test setting.
Table 1
Characteristics of the column specimens
Location Specimen name Loading plan Lap splice
OMRCF (IMRCF)
Interior
Lower part OIL (IIL)
Constant
column axial load
Upper part OIN (IIN)
(P = 0.28)
×
Exterior
Lower part OEL (IEL)
Varying
column axial load
Upper part OEN (IEN)
(P = 1.83V + 17.1)
×
O I L
(1): OMRCF (O), IMRCF (I) : having lap splices
(1) (2) (3)
(2): Interior (I), Exterior (E) ×: not having lap splices
(3): with lap splice (L), without lap splice (N)
Quasi-static reversed cyclic loading (displacement con-
trolled) was applied. Two cycles were applied for each load-
ing amplitude. The loading history is shown in Fig. 5. Since
ﬂuctuation of axial loads in the exterior columns was more
signiﬁcant than in the interior columns, varying axial load-
ing was applied to the exterior column specimens, whereas a
constant axial load was applied to the interior column spec-
imens. Larger axial loads were applied to interior columns
due to a larger tributary area.
Axial loads were calculated by frame analysis using
SAP 2000 commercial software [5] without considering load
factors. Since specimens were scaled by 2/3, calculated

Fig. 5. Loading history.
Table 2
Concrete properties of the specimens
Design strength 28 day strength Strain at ultimate Young’s modulus
(MPa) (MPa) strength, εco (MPa)
23.5 24.6 0.003 23,438
Table 3
Reinforcing steel properties
Bar Yielding strength Yielding strain Ultimate strength Young’s modulus
(MPa) (×10−6) (MPa) (MPa)
D6 374 2206 598 176,520
D13 397 2035 594 194,956
force for original frame members should be scaled by 4/9
for the scaled model. Interior column specimens (OIL, OIN,
IIL, IIN) were tested with a constant axial load of 333.4 kN
(0.28 Ag fc; Ag is the gross sectional area of a column). For
exterior column specimens, varying axial loads (P (axial
force) = 1.83V (lateral force) + Pg) were applied. Pg is
gravity axial load, which is 167.6 kN, and V is lateral
shear force acting on the column. This relationship was
also obtained using SAP 2000 under first mode lateral force
profile. The range of varying axial loads for exterior column
specimens is 0.07 ∼ 0.20Ag fc.
As shown in Fig. 4, one actuator was installed to control
lateral forces and two actuators were used to control
axial forces. Three pairs of linear transducers were placed
at the column faces to capture column curvatures (see
Fig. 4(a)). However, curvature results were not included in
this paper. One LVDT was also placed to measure a lateral
displacement of the stub of column specimens. Since it
was observed that slip between the stub and strong floor is
negligible during the test, recorded displacement of column
specimens was not modified. This was due to tightened bolts
passing through stubs of specimens. Curvature of stubs was
not measured since it was assumed that tightened bolts made
a rigid base condition. Four linear transducers were installed
to measure the lateral displacements of specimens.
4. Test results and observation
4.1. Observations
Within a drift ratio of ±1%, the first flexural crack was
observed in all specimens. The lateral force causing the first
crack in all specimens was about 25 kN. At a drift ratio of
±3%, the concrete cover started spalling at the column ends.
It is noted that drift ratio was defined as a measured lateral
displacement divided by the height of specimens.
In the exterior column specimens having lap splices
(OEL, IEL), spalling and cracks (vertical and horizontal)
were more prominent when applied lateral loading was in
the positive direction (the direction in which axial load
increases). As mentioned earlier, varying axial loads (P =
1.83V + 167.6 (kN)) were applied to exterior column
specimens. Moreover, many vertical cracks in the region of
lap splices were observed. Flexural cracks were relatively
uniformly distributed in the plastic hinging region of the
specimens.
In the final stage of the test, all column specimens failed
due to buckling of the longitudinal bars and crushing of the
concrete.
4.2. Hysteretic performance
Hysteretic curves for the OMRCF and IMRCF columns
are presented in Figs. 6 and 7, respectively. Fig. 6(a), (b) and
(c), (d) shows the hysteretic curves for the interior specimens
OIL and OIN, and for the exterior column specimens OEL
and OEN of the OMRCF. Fig. 7(a), (b) and (c), (d) show
the curves for the interior specimens IIL and IIN, and for
the exterior column specimens IEL and IEN of the IMRCF.
All column specimens behaved almost elastically until a drift
ratio of ±0.3%.
According to Fig. 6(a) and (b), strength of specimens
OIN and OIL was similar irrespective of the existence
of lap splices. However, specimen OIN exhibited greater
deformation capacity than the OIL specimen having lap
splices. In the IMRCF interior column specimens, IIN and

(a) OIL.
(b) OIN.
(c) OEL.
(d) OEN.
Fig. 6. Hysteretic curves of the OMRCF columns.
(a) IIL.
(b) IIN.
(c) IEL.
(d) IEN.
Fig. 7. Hysteretic curves of the IMRCF columns.

IIL, (Fig. 7(a) and (b)), no discernible difference could be
found in either strength or deformation capacity.
The strength of all the OMRCF and IMRCF interior
column specimens (OIN, OIL, IIN, IIL) was similar,
whereas OMRCF interior columns had inferior deformation
capacities to the IMRCF interior columns. Moreover,
pinching was detected more prominently in OMRCF
specimens. Strength drops in the second cycle were,
however, similar between the OMRCF interior column
specimens and their corresponding IMRCF specimens.
In interior column specimens, IMRCF has a more stable
and fuller hysteretic curve than OMRCF. In particular, the
improvement was more prominent when columns had lap
splices (compare Figs. 6(a) and 7(a)). The better hysteretic
behavior in interior columns of IMRCF was attributed to
narrower spacing of transverse reinforcement in the plastic
hinging region of the column.
Fig. 8(a) shows the envelope curves extracted from
the hysteretic curves of the interior column specimens.
According to this figure, the IMRCF specimens had fuller
envelope curves than the OMRCF specimens. Thus, it
is expected that interior column specimens of IMRCF
exhibited better seismic behavior than those of OMRCF.
Since larger axial force was applied to the interior specimens
due to a larger tributary area of gravity loads, narrower
spacing of transverse reinforcement in the plastic hinge
region improved the seismic behavior of the columns
significantly.
Figs. 6(c) and (d), and 7(c) and (d) show hysteretic
behaviors of the exterior column specimens of the OMRCF
(OEL, OEN) and IMRCF (IEL, IEN), respectively. In these
figures, the positive loading direction was the direction in
which the axial load increases.
The hysteretic behaviors of the exterior OMRCF
and IMRCF column specimens were relatively similar
between the OMRCF specimens and corresponding IMRCF
specimens (OEN versus IEN, and OEL versus IEL).
However, the strength drop of specimen OEL in the
second cycle was greater than that of specimen IEL. The
specimens having lap splices showed inferior deformation
capacities compared to the specimens without lap splices
(OEL versus OEN, and IEL versus IEN). Unlike interior
columns, hysteretic behavior was not significantly improved
by narrower spacing of lateral reinforcement (OEL versus
IEL, and OEN versus IEN). Fig. 8(b) shows the envelope
curves. From this figure, the specimens having lap splices
also had narrower envelope curve. Also, this figure shows
that the envelope curves of the exterior column specimens
of OMRCF and IMRCF are quite similar (OEN versus
IEN, and OEL versus IEL). However, more strength drop in
specimen OEL was observed than that in IEL in the negative
loading direction. In exterior column specimens, lateral
reinforcement spacing does not affect seismic behavior as
much as in the interior column specimens. However, it was
observed that seismic behavior in exterior column specimens
became poor when they had lap splices.
4.3. Maximum strength
Table 4 shows the ratio of actual maximum strength Mmax
obtained from experimental tests to the calculated strength
(MACI). Actual maximum strength Mmax is calculated by
maximum shear force times the height of the column
specimen (Mmax = Vmax × 1m(h)). Fig. 9 shows P–M
interaction curves of column specimens. This figure also
includes the actual moment strength (Mmax) of column
specimens. All nominal strength (MACI) in this study was
calculated using material strength obtained from material
tests. According to Fig. 9 and Table 4(7), all OMRCF
and IMRCF column specimens have strength larger than
calculated nominal strength.
In exterior column specimens, moment capacities (Mmax)
in the positive loading direction were greater than those in
the negative loading direction. Since axial loads in the test
are not large, strength of column specimens is below the
balanced failure point in P–M interaction curve. Thus, in
exterior column specimens, moment strength increases as
axial force is larger (positive lateral loading direction).
In the P–M interaction curve (Fig. 9), calculated moment
strength (MACI) of exterior column specimens in the positive
loading direction is larger than that in the negative loading
direction as well. In those specimens with relatively low
axial loads, the ratio of maximum strength (Mmax) to
calculated strength (MACI) was almost the same as that of
the interior column specimens.
It is worthwhile to note that all column specimens
were designed according to the minimum design and detail
requirements in ACI 318-02. Since all specimens in this
study were governed by flexure rather than shear (see
Table 4(10)), different results can be obtained for columns
governed by shear.
4.4. Deformation capacity
In this study the maximum drift ratio (θmax = ∆max/h)
was used to determine the drift capacities, where ∆max is
the maximum drift and h is the height of column specimens.
The maximum drift was obtained when the lateral strength
of the specimen was reduced by 20% of maximum strength.
Table 4 and Fig. 10 show the maximum drift ratios of the
specimens.
Table 4 shows that the drift capacities (θmax) of
OMRCF and IMRCF exceeded 3.0% and 4.5%, respectively.
Specimen IEN had the largest drift capacity (5.98%) in the
negative loading direction, whereas specimen OEL had the
least drift capacity (3.00%) in the negative loading direction.
Among the interior column specimens that resisted a
larger constant axial load throughout the test, the IMRCF
columns had greater drift capacity. This is due to the
narrower spacing of transverse reinforcement in the plastic
hinging region. It should be noted that the drift capacity of
the specimen IIL having lap splices was even greater than
that of OIN without lap splices. Moreover, the difference

(a) Interior columns.
(b) Exterior columns.
Fig. 8. Envelope curves.
Fig. 9. Maximum strength.

Table 4
Test result of specimens
Specimen Pu
Ag fc
Vmax ∆max µ∆ θu MACI
Mmax
MACI
VACI VP
VP
VACI
(1) (2) (3) (4) (5) (6) (7) (8) (9) (10)
OMRCF
OIL
+
0.28
49.1 34.9 4.26 3.49 36.8 1.33 74.3 37.9 0.51
− 49.5 35.3 4.46 3.53 36.8 1.33 74.3 37.9 0.51
OIN
+ 46.9 44.5 4.41 4.45 36.8 1.28 74.3 37.9 0.51
− 48.8 44.5 4.39 4.35 36.8 1.33 74.3 37.9 0.51
OEL
+ 0.2 36.9 42.6 4.34 4.26 33.3 1.25 69.6 32.6 0.47
− 0.07 37.2 28.5 3.02 3.01 25.5 1.25 63.4 26.0 0.41
OEN
+ 0.2 48.5 45.3 4.59 4.53 33.4 1.45 69.6 32.6 0.47
− 0.07 28.3 58.1 6.05 5.81 25.5 1.11 63.4 26.0 0.41
IMRCF
IIL
+
0.28
48.7 53.8 6.02 5.38 38.8 1.28 9.83 3.88 0.39
− 45.9 53.5 5.57 5.35 38.8 1.21 9.83 3.88 0.39
IIN
+ 46.0 49.5 6.10 4.95 38.8 1.34 9.83 3.88 0.39
− 47.6 50.5 6.01 5.05 38.8 1.22 9.83 3.88 0.39
IEL
+ 0.2 45.5 45.2 5.82 4.52 33.0 1.55 9.36 3.30 0.35
− 0.07 33.2 47.0 5.90 4.70 26.2 1.10 8.72 2.62 0.30
IEN
+ 0.2 49.4 48.0 5.9 4.80 33.0 1.37 9.36 3.30 0.35
− 0.07 34.2 59.8 6.27 5.98 26.2 1.31 8.72 2.62 0.30
(2) = maximum shear force (kN), (3) = maximum displacement (mm), (4) = displacement ductility (= ∆max/∆y), where ∆y is yield displacemet, (5)
= maximum drift angle (%), (6) = moment capacity calculated using ACI 318-02 procedures (kN m), (7) = ratio of the maximum moment capacity,
Mmax(Vmax × h) to MACI (h: height of specimen (m)), (8) = nominal shear strength according to ACI 318-02 (kN), (9) = shear force corresponding to
MACI or 2MACI/l, where l = the column clear height (kN), (10) = ratio of VP to VACI.
Fig. 10. Drift capacity (%).
between drift capacities of specimens OIL and IIL was
greater than that of specimens OIN and IIN. Thus, spacing of
transverse reinforcement in the plastic hinging region affects
the drift capacity of the interior columns significantly. This
was more prominent when the specimen had lap splices.
It is noted that larger gravity load (0.28Ag fc) was applied
to interior column specimens. Also, in interior column
specimens of OMRCF, the existence of lap splices in the
column specimen prominently influences the drift capacity.
In contrast, the existence of lap splices did not affect drift
capacities of IMRCF column specimens.
In the exterior column specimens that resisted varying
axial loads during the test, the existence of lap splices also
affected drift capacities. Specimens IEL and OEL having lap
splices had smaller drift capacities. The difference between
drift capacities of the exterior columns of the OMRCF and
IMRCF was small compared to that between the interior
columns of the OMRCF and IMRCF. Thus, in the exterior
columns, the spacing of transverse reinforcement was not
influential. It is noted that this study considered a three-story
building, in which axial forces in columns are small
compared with those in high-rise buildings. When axial
forces in columns become large, the spacing of transverse
reinforcement may be influential even on the behavior of
exterior columns.
This study also calculated the displacement ductility
of each specimen, which is shown in Table 4. Yield drift
(∆y or θy) was obtained from a bilinear representation of
force (V) versus drift ratio (∆ or θ). Secant stiffness was
used to connect the origin to the point of 0.75Vmax. The

Fig. 11. Energy dissipation.
maximum drift (∆max, θmax) was approximated as the drift
ratio corresponding to the strength deteriorated by 20% of
Vmax (0.8 times Vmax). As shown in Table 4, all specimens
have a ductility capacity exceeding 3.0. The IMRCF exterior
column specimen without lap splices (IEN) has the largest
ductility capacity (6.27) in the negative loading direction
which has the least axial loading condition ( Pu
Ag fc
= 0.07).
In contrast, the OMRCF exterior column specimen with lap
splices (OEL) has the smallest ductility capacity (3.02) in
the negative loading direction. Specimen IEL has a ductility
capacity of 5.90 in the negative loading direction. Thus the
effect of lateral reinforcement spacing on displacement and
ductility capacity becomes important to all interior column
specimens, and exterior specimens having lap splices.
4.5. Energy dissipation
Fig. 11 shows the amount of energy dissipated in each
loading cycle. According to this figure, all specimens
dissipated a similar amount of energy until cycle 6 (1%
drift ratio). Beyond that cycle, every specimen dissipated a
different amount of energy.
Specimens IIN, IIL, and IEN dissipated similar amounts
of energy until failure. They have larger energy dissipation
capacity than other specimens. Specimens OEN, OIN, and
IEL dissipated similar amounts of energy until cycle 13
(5%). Since specimen OIL failed in an earlier loading stage
(cycle 11), the amount of energy of specimen OIL is smaller
than specimens OEN, OIN and IEL. Specimen OEL has
the least amount of energy dissipation capacity among the
specimens.
5. Conclusions
This study investigates the seismic behaviors of columns
in OMRCF and IMRCF. For this purpose eight specimens
were modeled and tested. The specimens were designed
and reinforced in compliance with the ACI 318-02 for
OMRCF and IMRCF. It should be noted that all specimens
were governed by flexure rather than shear. Thus, different
results may be obtained for columns governed by shear. The
conclusions of this study are as follows:
(1) The strength of all OMRCF and IMRCF column
specimens exceeded the strength calculated with the
code formula (ACI 318-02). Thus, the OMRCF and
IMRCF columns had satisfactory strength irrespective
of the existence of the lap splices and the transverse
reinforcement spacing in the plastic hinging region.
(2) The IMRCF interior column specimens had superior
drift capacities compared to the OMRCF column
specimens. Existence of lap splices influenced drift
capacities of the OMRCF columns. However, the effect
of lap splice in the IMRCF columns was not as large as
that in the OMRCF specimens. This is attributed to the
narrower transverse reinforcement spacing in the plastic
hinging region of the IMRCF.
(3) In the exterior column specimens, no discernible
differences in the hysteretic behaviors were found
between OMRCF and IMRCF column specimens. The
specimens having lap splices showed inferior hysteretic
behavior to the corresponding specimens without lap
splices. Since smaller axial force was applied to the
exterior column specimens, the effect of transverse
reinforcement spacing seems to be less influential than
interior columns. However, it is noted that the effect of
transverse reinforcement spacing becomes important to
exterior specimen having lap splices, particularly in the
negative loading direction.
(4) IMRCF interior column specimens had more stable and
fuller hysteretic curves than OMRCF interior column
specimens. In OMRCF interior column specimens,
existence of lap splices changed the shape of hysteretic
curve significantly. However, in IMRCF interior column
specimens, the change due to lap splices was not
noticeable.

(5) In exterior column specimens of OMRCF and IMRCF,
hysteretic curves became narrower when they had lap
splices. However, the hysteretic curves of OMRCF
exterior columns are similar to those of corresponding
IMRCF exterior columns.
(6) According to the test results, the OMRCF and IMRCF
column specimens had drift capacities greater than
3.0% and 4.5%, respectively. Ductility capacity of those
specimens exceeded 3.01 and 4.53, respectively.
Acknowledgement
The support of Hanyang University is greatly acknowl-
edged.
References
[1] American Concrete Institute. Building code requirements for
reinforced concrete, ACI 318-95, 99, 02, Detroit, Michigan, 1995,
1999, 2002.
[2] Uniform Building Code (UBC). International conference on building
ofﬁcials. 1997.
[3] Building Seismic Safety Council. NEHRP recommended provisions
for the development of seismic regulation for new buildings, Part 1
and 2, provisions and commentary, Washington, DC: FEMA; 1994,
1997.
[4] Portland Cement Association. Notes on ACI 318-99, 02 building code
requirements for structural concrete, Skokie, Illinois, 1999, 2002.
[5] Computers and Structures Inc. SAP2000, 1997.
[6] Han SW, Kwon OS, Lee L-H, Foutch DA. Seismic performance
evaluation of ordinary moment resisting concrete frames. SRS No.
633, Urbana, Illinois; 2002 November.
[7] Han SW, Kwon OS, Lee L-H. Evaluation of the seismic performance
of a three-story ordinary moment resisting concrete frame. Earthquake
engineering and structural dynamics, vol. 33. Wiely and Sons; 2004.
p. 669–85.
[8] BOCA, ICBO, SBCCI. International Building Code. 2000.
[9] Lee H-S, Woo S-W. Effect of masonry inﬁlls on seismic performance
of a 3-storey RC frame with non-seismic detailing. Earthquake
engineering and structural dynamics, vol. 31. Wiely and Sons; 2002.
p. 353–78.
[10] Aycardi LE, Mander JB, Reinhorn AM. Seismic resistance of
reinforced concrete frame structures designed only for gravity loads:
experimental performance of subassemblages. Structural Journal
American Concrete Institute 1994;91:552–63.
[11] Lynn A, Moehle J, Mahin SA, Holmes WT. Seismic evaluation
of existing RC building columns. Earthquake Spectra 1996;12(4):
715–39.
Sang Whan Han. Associate Professor, Dept. of Architectural Eng.,
Hanyang Univ., Seoul 133-791, Korea. Ph.D. (1994), University of Illinois
at Urbana-Champaign, Professional Engineer (US, Korea).
N.Y. Jee. Professor, Dept. of Architectural Eng., Hanyang Univ., Seoul 133-
791, Korea. Ph.D. (1992), Tokyo National University.

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