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Engineering Structures 24 (2002) 719–734

www.elsevier.com/locate/engstruct

Seismic performance of a 3-story RC frame in a low-seismicityregion

Han-Seon Lee *, Sung-Woo Woo

Department of Architectural Engineering, Korea University, Seoul 136-701

Received 16 April 2001; received in revised form 7 November 2001; accepted 14 November 2001

Abstract

The objectives of the research stated herein are to investigate the seismic performance of a 3-story reinforced concrete (RC)

ordinary moment-resisting frame, which has not been engineered to resist earthquake excitations, and to evaluate the reliability of

the available static and dynamic inelastic analysis techniques. A 1:5 scale model constructed according to the Korean practice of

nonseismic detailing and the similitude law was subjected to a series of the shaking table motions simulating Taft N21E component

earthquake ground accelerogram. Due to the limitation in the capacity of the used shaking table, a pushover test was performed to

observe the ultimate capacity of the structure after earthquake simulation tests.

The model showed the linear elastic behavior under the Taft N21E motion with the peak ground acceleration of 0.12g, representing

the design earthquake in Korea. The model revealed fairly good resistance to the higher levels of earthquake simulation tests though

it was not designed against earthquakes. The main components of its resistance to the high level of earthquakes appear to be (1)

the high overstrength, (2) the elongation of the fundamental period, (3) the minor energy dissipation by inelastic deformations, and

(4) the increase of the damping ratio. The drifts of the model under these tests were approximately within the allowable limit.

Analysis of the results of the pushover test reveals that the model structure has the overall displacement ductility ratio of 2.4 and

the overstrength coefficient of approximately 8.7. The evaluation of the accuracy of analytical simulation by IDARC-2D leads to

the conclusion that while global and local behaviors can be, in general, simulated with limited accuracy in the dynamic nonlinear

analysis, it is easy to obtain a fairly high level of accuracy in the prediction of global behavior in the static nonlinear analysis.

2002 Elsevier Science Ltd. All rights reserved.

Keywords: Reinforced concrete; Earthquake simulation test; Pushover test; Nonlinear analysis; Overstrength; Nonseismic detailing

1. Introduction

Recently, minor earthquakes have occurred over 20times a year in Korea. The earthquake of December 13,1996 at Yeongweol in Korea, which is known to havea magnitude of 4.5 on the Richter scale imposed signifi-cant nonstructural damages [1]. These recent earth-quakes indicate that the Korean Peninsular is no longersafe from seismic hazard. If a severe earthquake such asthe 1995 Kobe earthquake should occur in Seoul, thedamage would be tremendous.Most building structures in Korea, which are normally

medium- to low-rise reinforced concrete (RC) frames,

* Corresponding author. Tel.: +82-2-3290-3337; fax: +82-2-921-

7947.

E-mail address: [email protected] (H.-S. Lee).

0141-0296/02/$ - see front matter 2002 Elsevier Science Ltd. All rights reserved.

PII: S0141-0296 (01)00135-3

have not been engineered to resist major or moderateearthquakes. Therefore, should any major earthquakeoccur, the damage or collapse of not only general com-mercial buildings, but also public-service buildings suchas police offices, communication centers and hospitals,would implement very large life and economic losses aswell as cause the critical interference with the functionof the nation.Several researchers in the eastern and central United

States have already performed research, including earth-quake simulation tests, on the seismic capacity ofgravity–load-designed RC structures [2–5]. However,attention concentrated mainly on the detail practice inNorth America, particularly according to American Con-crete Institute (ACI) 318 code [6]. The practice inreinforcement detailing and construction in Korea issomewhat different from that of the US, which will beexplained in the section on the design of the model.

720 H.-S. Lee, S.-W. Woo / Engineering Structures 24 (2002) 719–734

The objectives of this research are; (1) to observe theactual response of this kind of low-rise RC ordinarymoment-resisting frame with nonseismic detailing whensubjected to various levels of earthquake groundmotions, (2) to get the information on the ultimatecapacities (strength, deformability and so on) of thestructure, and (3) to provide the calibration to, or tocheck the reliability of, the available static and dynamicinelastic analysis techniques.Considering the capacity of the shaking table to be

used, the reduction scale for the model was determinedas 1:5. Using the techniques for manufacturing themodel according to the similitude requirementsdeveloped through other researches [7,8], a 1:5 scale 2-bay 3-story RC frame model was constructed. Thismodel was then subjected to the shaking table motionssimulating Taft N21E component earthquake groundmotions, whose magnitude of peak ground acceleration(PGA) was modified to approximately 0.12, 0.2, 0.3 and0.4 g. Before and after each earthquake simulation test,free vibration tests were performed to determine thechange in the natural period and the damping ratio ofthe model. The global behavior and damage pattern wereobserved. The lateral accelerations and displacements ateach story and the local deformations at the criticalregions of the structure were measured. The base shearwas measured using self-made load cells. Since the ulti-mate capacity of the structure could not be found due tothe limitation in the capacity of the shaking table, apushover test was performed to observe this capacity ofthe structure after a series of earthquake simulation tests.Based on all the test results, the interpretation on theresponse of the model is carried out.The computer code IDARC-2D [9], one of the codes

widely used in the world for the nonlinear dynamic andstatic analyses of RC frame structures, was adopted forthe analysis to investigate the correlation between theresults of (earthquake simulation and pushover) tests andanalyses. Although the time history analyses correspond-ing to all the earthquake simulation tests were performed,the correlation of experiment and analysis will be inves-tigated only for the case of earthquake simulation testTFTF04 in this paper.

2. Experiment

2.1. Design of the model

The prototype of this test model was adopted from abuilding structure for the police office, actually built andin use in Korea. The plan and elevation of the 1:5 scalemodel are shown in Fig. 1(a,b). The compressivestrength of concrete, f9c, in the prototype structure isassumed to be 20.6 MPa and the nominal yield strengthof reinforcement, 294.2 MPa. The typical sections of

members and the details regarding transverse steel,anchorage and splice are shown in Fig. 1(c–h).The important characteristics in the Korean detailing

practice are as follows: (1) the splice is located at thebottom of the column, (2) the spacing of hoops is rela-tively large, (3) seismic hooks are not used, (4) con-finement reinforcements are not used in beam–columnjoints, and (5) the special style of anchorage in the joints.That is, the length of tension and compression anchorageare usually 40 and 25 db respectively, from the criticalsection, where db means the nominal diameter ofreinforcement. Moreover, the length of the tail in thehook is included in this anchorage length and the tailsof the anchorage of the bottom bars in beams usuallydirect downward into the exterior columns as shown inFig. 1(h).

2.2. Model reinforcement and model concrete

It is essential to maintain the similitude in the materialproperties between prototype and model reinforcement.However, it was difficult to make the cross sections ofthe model reinforcement conform exactly to the simili-tude law. So, the yield forces rather than yield stresseswere selected as the target to be achieved in annealingthe model reinforcement. An electric furnace with a 3-zone vacuum tube was designed and used. The defor-mation on the surface of the model reinforcements wasmade using a deforming device. Reinforcing bars D22and D10 in the full-scale structure match D3 and D2 inthe model, respectively. The target yield forces derivedfrom similitude requirements are shown in Table 1. Theachieved average yield force of 5.0 kN is approximately10% less than the target yield force in the case of D3.The model concrete was made using type I Portlandcement. The average strength of the model concrete atthe time of testing was about 21.6 MPa.

2.3. Instrumentation and experimental setup

The used shaking table is 3×5 m with one degree offreedom. The data were acquired simultaneously at therate of 300 Hz in 32 channels. Displacement transducers,accelerometers and load cells were used to measure thelateral displacement and the angular rotations in someends of beams and columns, acceleration at each story,and shear forces on the columns of the first story. Fig.2(a) and Plate 1 show the view of the experimental setupand instrumentations for the shaking table test. Theinstrumentation was confined to only one frame due tothe limitation in the number of available channels exceptthe load cells at the first-story columns. To measure storydrifts, a reference frame was used as shown in Fig. 2.The white-noise (random-vibration) test, which was per-formed before the earthquake simulation tests, indicatedthat the natural frequency of the reference frame is

721H.-S. Lee, S.-W. Woo / Engineering Structures 24 (2002) 719–734

Fig. 1. Plan, elevation and details (unit: mm). (a) Plan; (b) Elevation; (c) Section C1–C19; (d) Section C2–C29; (e) Section C3–C39; (f) Section B1–

B19; (g) Section B2–B29; and (h) Anchorage detail of beam bars in exterior joint.

approximately 40 Hz. Therefore, the reference framewas considered to have the sufficient rigidity to measureaccurately the drifts of the model. The load cells weredesigned and manufactured following the reports of El-Attar et al. [4] and Bracci et al. [2] and calibrated byusing a universal testing machine (UTM).The volume of the model is reduced to 1/53 of the

prototype while the similitude law, premising the agree-ment in the stress–strain relation of the material betweenthe prototype and the model, requires the mass to be

reduced to 1/52 [10]. Therefore, compensation for thedifference in the mass was artificially made by addingconcrete blocks as shown in Fig. 2(a). The effectiveweight of the model with concrete blocks is estimatedto be 100.6 kN while the weight ideally required by thesimilitude law is 96.1 kN. The error is approximately5%. The hinges between the model and the concreteblocks were designed to transmit only the vertical andhorizontal forces, excluding the moments, due to themass of the concrete block.


Table 1

Similitude requirements for reinforcement

Member and use Prototype (stress) 1/5 Model (force)

Beam and column Bar Yield strength Tensile strength Bar Yield strength Tensile strength

Stress Force Stress Force Stress (1)×1/25* Stress (2)×1/25

(MPa) (kN)(1) (MPa) (kN)(2) (MPa) (kN) (MPa) (kN)

Stirrup and hoop D10 358 25.5 530 37.8 D2 325 1.02 481 1.51

Main reinforcement D22 358 138.5 530 205.3 D3 784 5.54 1161 8.21

* Target yield force in annealing.

Fig. 2. Instrumentation and experimental setup. (a) Earthquake simulation test; and (b) Pushover test.

The experimental setup for the pushover test is shownin Fig. 2(b). Steel plates, each of which has the dimen-sion W×D×L=9×4×35 cm (0.097 kN) were used as theartificial mass. The effective weight of the model withsteel plates is estimated to be 100.7 kN. The roof driftwas obtained by averaging the two measured values inboth Frame A and Frame B. Two different data acqui-sition systems were used due to the limitations in thenumber of available channels in the data acquisition sys-

tems, but the measured data were synchronized witheach other’s system by comparing the displacements oftransducers T5 and T6 installed at the same location inFig. 2(b) but belonging to the different data acquisitionsystem. The experimental results were interpreted,assuming that the behavior of Frame A represents thatof the whole model structure in both earthquake simul-ation test and pushover test. The displacements at thesecond floor (T2) and third floor (T3) were those meas-


Plate 1. Experimental setup for shaking table tests.

ured in the middle of both frames, and the lateral forcedistribution was maintained in the shape of an invertedtriangle by using the whiffle tree.

2.4. Experimental program

Since there was no recorded strong-motion accelerog-

ram as of the time of this experiment in Korea, the inputmotion to the shaking table was derived from the

recorded Taft N21E component by adjusting the peak

ground acceleration (PGA) and compressing the timescale by the factor of Î5 according to the similitude law[10]. The most important reason for this selection was

that the test results in this study can be compared easilywith those of several previous earthquake simulation

tests [2–4,11] which used the same accelerogram, and

the other one is that damage potential of Taft N21Ecomponent accelerogram seems to be relatively high.

The design earthquake defined in Korean seismic code

[12] is the event expected in the recurrence period of475 years (the probability of exceedance of 10% in 50

yr). The elastic response spectra of the input Taft N21E(PGA = 0.12 g) motion and the output table motion are

shown in Fig. 3. From this figure, it can be found that

Fig. 3. Response spectrum for input and output table motions and

design spectra compressed by the scale of 1:Î5 (soil factor S 5 1.2).

the fidelity of the shaking table is satisfactory. The com-

parison of the design spectrum according to the Koreanseismic code [12] with the response spectrum of the out-

put motion indicates that the dynamic amplification

implicit in the design spectrum of this code is underesti-mated in the range of short periods whereas it is overesti-

mated in the range of long periods. The final design

spectrum by UBC [13] is quite similar to the Koreandesign spectrum though the used R factors are different.

The program of tests is shown in Table 2. Each testhas the significance stated in the column containing

remarks. For earthquake simulation tests, the model

structure did not show serious damage even after testTFTF04. However, due to the limitation in the capacity

of the shaking table, it was impossible to implement a

higher level of earthquake simulation test. Therefore, inorder to get the information on the capacities (strength,

deformability and so on) of the model structure, a push-

over test (or monotonically-increasing lateral-load statictest) was conducted.

3. Results of earthquake simulation tests and

interpretation

3.1. Global responses

Free vibration tests were performed by enforcing aninitial lateral displacement of approximately 1–2 mm at

the roof of the model, and releasing. From these tests,the natural periods and damping ratios of the model were

obtained by using the Fourier transform and logarithmic

decrement method. Table 3 shows that natural periodsand damping ratios tend to increase as the model experi-

ences higher levels of ground motions.

The drifts at the roof were measured at two locations.The test results indicate that almost negligible torsional

behaviors occurred in this model for both earthquake

simulation and pushover tests. Therefore, the measureddata in one of the two plane frames are assumed to rep-

resent the behaviors of both frames. Table 4 summarizes

the measured maximum response quantities. Time his-tories of the floor accelerations for TFTF04 are given in

Fig. 4, which indicates that the response has the pre-

dominant period approximately equal to the period ofthe first mode of the model.

Table 4 shows the maximum interstory drift indices

(I.D.I.) at the time when the roof undergoes themaximum drift. For the design earthquake of Korea

(TFTF012), the model has a maximum I.D.I. of 0.26%which is much less than the maximum allowable value

of 1.5%. However, the maximum I.D.I. for the test of

TFTF04 appears to be 1.68%. Therefore, regarding driftcontrol, the behavior of the model appears to be satisfac-

tory. From the profiles of measured drift and the

interstory drift indices in Fig. 5, it can be noted that the


Table 2

Test program

Identification of test PGA (g) Remarks (return period)

Earthquake simulation test TFTF012 0.12 Design earthquake in Korea (500 years)

TFTF02 0.2 Max. earthquake in Korea (1000 years)

TFTF03 0.3 Max. considered earthquake in Korea (2000 years)

TFTF04 0.4 Severe earthquake in high seismic regions of the world

Pushover static test PUSH Ultimate capacity of the structure

Table 3

Natural period and damping ratio by free-vibration tests

Identification of test Natural period (sec) Damping ratio (%)

Before TFTF012 0.226 4.1

After TFTF012 0.229 4.6

After TFTF02 0.265 4.4



displaced shapes of the frame remained almost

unchanged during the course of the tests and the defor-

mations tend to be concentrated at the second story forthe higher levels of motions.

The model behaves linear-elastically under TFTF012,

which represents the design earthquake in Korea. Thedeveloped maximum base shear is 17.64 kN. The

maximum base shears for other table motions are given

in Table 4. The model has yielded under TFTF04. Herethe yielding base shear appears to be 37.14 kN.

Though the building structures built before 1988 or of

Table 4

Summary of measured maximum response amplitudes

Test Table acceleration Roof drift (mm) I.D.I. (%) Roof acceleration Base shear (load V/W

(g) (g) cell) (kN)

TFTF012 0.138 4.5 0.26 0.28 17.64 0.18

TFTF02 0.21 14.06 0.78 0.53 30.77 0.32

TFTF03 0.31 17.87 1.08 0.61 35.08 0.37

TFTF04 0.4 29.88 1.68 0.69 37.14 0.39

Fig. 4. Time histories of floor accelerations for TFTF04.

less than six stories need not be designed against theearthquake in Korea, it is worthwhile to evaluate the test

results with regards to the current Korean seismic code

[12]. According to this code, the seismic coefficient, Cw,for the RC ordinary moment-resisting frame is determ-

ined as follows:

Cw 5 SVWD 5

AISC

R5

(0.12)(1.0)(1.39)

3.55 0.048 (1)

where V: base shear, W: 96.1 kN (effective weight of

the structure), A: 0.12 (zone factor), I: 1.0 (importance

factor), T: 0.23×Î5 (scale factor) =0.514 s (natural

period), S: 1.2 (soil factor), C:1

1.2ÎT5 1.162#1.5

(dynamic factor), SC: 1.39 (#1.75), R: 3.5 (response

modification factor; ordinary moment frame).Table 4 also gives the coefficient derived from the

experiment. For the table motion (TFTF012) simulating

the design earthquake in Korea, the actual developedmaximum base shear is about 3.8 times the design base

shear. It can be noted from Table 4 that this model


Fig. 5. Profiles of drift envelopes and interstory drift indices in earthquake simulation test. (a) Drift envelopes; and (b) Interstory drift indices

(I.D.I.).

resisted the table motions of higher PGA’s with a veryhigh lateral strength of up to 0.39 W. In addition, the

degradation of stiffness has elongated the natural period

of the model as shown in Table 3 and hence caused thedecreased dynamic amplification factor. The test of

TFTF04 has shown that the model yielded and that the

story displacement ductility at the first story was about1.3. From these findings, it can be concluded that the

model behaved linear-elastically under the design earth-

quake (PGA>0.12 g) while the overstrength up to 0.39W, the degradation of stiffness, the energy dissipation

due to inelastic behavior in the model, and the increaseof damping ratio were the additional major contributors

to its resistance against the higher table motions

(PGA=0.2, 0.3 and 0.4 g).The time histories of the total absorbed energy at each

story for the cases of TFTF012 and TFTF04 are depicted

in Fig. 6. From this figure we can find that the totalamount of the absorbed energy in the case of TFTF04

is about 13 times that in the case of TFTF012, and that

most of the energy absorption took place in the lowertwo stories.

The model did not show serious damage even after

the test of TFTF04 simulating a severe earthquake in ahigh-seismicity zone of the world. There was no appar-

ent crack after tests TFTF012 and TFTF02. After test

Fig. 6. Time histories of absorbed energy in earthquake simulation

test.

TFTF03, a minor flexural crack could be noticed in beamend at the exterior joint of the second floor. Several

cracks occurred after test TFTF04. The beam ends at the

exterior joint of the second floor revealed new cracks,which imply a yield in these regions. In particular, the

exterior columns at the first story have revealed both

flexural and shear cracks. However, it is interesting tonote that the interior columns, which have experienced

larger rotations, did not show any apparent cracks.

3.2. Local responses

The distance over which the rotational angle was mea-

sured is the full depth for beams, and the cross-sectional

dimension parallel to the shaking direction for columns.Fig. 7 shows the example of some time histories of angu-

lar rotations in the ends of members for earthquake

simulation test. The bottom end of the interior columnat the second story has shown the largest angular rotation

of all that were measured. Relatively large rotations

occurred for the interior columns (R4, R3, and R8 inFig. 2), when compared with those for the exterior col-

umns (R6, R7, and R9 in Fig. 2). Since the model has

smaller cross-sections in the interior columns than in theexterior columns, these larger rotations at the interior

column could be expected. For the exterior joint at the

second floor, the beam (R5) has larger rotation than the

Fig. 7. Angular rotations under TFTF04 (exterior joint).


connected columns (R6 and R7). On the contrary, the

angular rotations at the beams (R1 and R2) are generallymuch less than those at the columns (R3 and R4) in the

case of the interior joint.

Fig. 8 shows the distribution of shear forces in thecolumns. It is interesting to note that column (1) and

column (3) have a directional bias in shear forces due

to the influence of the axial compressive force on theshear resistance.

4. Correlation of analytical and experimental

dynamic responses

4.1. Analytical model

The computer code IDARC-2D [9], one of the codes

widely used in the world for the nonlinear dynamic and

static analyses of RC framed structures, was adopted.This code utilizes a global Takeda-like model. The

objective of this analytical study is to find the mostappropriate values of parameters in the analysis byIDARC-2D to simulate the responses given by experi-ments, and then to evaluate the degree of accuracy inthe obtained simulations. Eventually, this evaluation ofreliability for IDARC-2D will lead to the more carefulinterpretation of analysis results for other types of RCframe structures such as schools, hospitals, and the like.The material models used to derive the relation

between moment and curvature at critical sections areshown in Fig. 9. The program RESPONSE, developedby Felber and Andreas [14], was used to obtain theenvelope curve for the moment–curvature relations, asdepicted in Fig. 10. The effective width (410 mm) ofACI code 318-95 [6] was used to model the relationbetween the curvature and the moment for the time his-tory analyses.The hysteretic model incorporates stiffness degra-

dation (HC, a), strength deterioration (HBD, HBE, b),non-symmetric response, slip-lock (HS, g), and a trilin-ear monotonic envelope. The model traces the hysteretic

Fig. 8. Time histories of column shears under TFTF04.

behavior of an element as it changes from one linearstage to another, depending on the history of defor-mations. For a complete description of the hystereticmodel, refer to Park et al. [15].Aycardi, Mander, and Reinhorn [5] used a 5 0.5,b 5 0.04, and g 5 0.7 based on the experimental testresults for elements. Stiffness degradation is severer inthe model than in the prototype, and an α value between0.5 and 1.0 has been used in the analytical model, whichis scaled as 1/4 or below [9]. In this study, the samevalue of hysteretic parameters as adopted by Aycardi etal. are used to simulate the test results.The artificial weight (concrete blocks in this case)

loaded to compensate the mass according to the simili-tude law acts as concentrated loads on the girders.Hence, the girders are divided into three elements whichalso properly represent the change in the reinforcementof the sections.


The parameters determining hysteretic behavior in theM–f relation have actually been adjusted to simulatemost closely the response, particularly the roof drift.Though the roof drift history shown in Fig. 11 is thebest simulation among all the trials, there is still a dis-crepancy of 4.11 mm (about 14%) in maximum values,and in phase in the latter part. Though the peak drift atthe roof in the analysis is smaller than in the experiment,the peak response acceleration at the roof of the analysisappears to be a little larger than that of the experiment.Nevertheless, the analysis reveals a smaller maximumbase shear than the experiment in Fig. 12. This discrep-ancy in the maximum base shear between analysis andexperiment can be attributed to the stronger effects ofthe second or higher mode in the analysis. In general,the histories of story drifts are similar in shape in bothcases of analysis and experiment, indicating that the firstmode governs, though the peak values differ.

4.3. Local responses

The time histories of the column shear in Fig. 13imply that the analysis could not simulate the bias in theshear force caused by the increase of shear stiffness dueto the increase of the axial compressive forces in col-umns. Fig. 14 compares the time histories of angularrotations at the bottom of the second-story columnobtained from both the experiment and analysis. Theangular rotations were calculated by multiplying the cur-vature at the end of members in analysis with the samelengths over which the rotations were measured in theexperiment. Generally, the magnitude of the angularrotation in the analysis is much smaller than that in theexperiment. The discrepancy in magnitude is considered


Fig. 9. Material model. (a) Concrete; and (b) Reinforcement.

Fig. 10. Moment–curvature envelope for T beams by RESPONSE [14].

Fig. 11. Comparison of roof drift history.

to be due to the limitation in the member modeling of

spread plasticity in IDARC-2D.

5. Interpretation on results of pushover test and

analyses

5.1. Analytical model

The model for the pushover analysis is the same as

that for the time history analysis except that the artificial

Fig. 12. Comparison of base shears derived from column shears.

weight needed to compensate the gravity weight was dis-

tributed over the slabs using steel plates rather than beingconcentrated weights as in the case of earthquake simul-

ation tests. To take into consideration the increase in the

contribution of slabs at the ultimate strength of the struc-ture, two types of sections were considered to model T

beams (the effective width of T beam, = 410 mm

(PUSH-I), =840 mm (PUSH-II)) as shown in Fig. 10.


Fig. 13. Time histories of column shear at first story (column (1)).

Fig. 14. Time histories of angular rotations in interior column of

second story for TFTF04.


The relations between the lateral load and the roof

drift obtained through the pushover test and analyses(PUSH-I and PUSH-II) are shown in Fig. 15. The find-

ings regarding this figure are as follows: (1) IDARC-2D

could not predict or simulate the brittle failure at thepoint of ultimate strength. (2) The additional contri-

bution of the enlarged width of the slab to the ultimate

strength, which is introduced by using PUSH-II modelinstead of PUSH-I is only approximately 4 kN. (3)

Fig. 15. Base shear versus roof drift in tests and analyses.

PUSH-II has indicated a still lower strength than that by

experiment. The difference is about 8 kN (17%). In thepushover test, the model structure yielded gradually and

underwent a series of abrupt strength drops after reach-

ing the ultimate strength of 51.35 kN through brittle fail-ures in the local regions, the locations of which could

not be clearly identified, with the exception of the last

failure by concrete crushing in the upper end of theinterior column at the first story. This clearly indicates

that the model structure has the characteristics of thestructure, which has limited ductility due to non-seismic

details. The solid marker, representing the maximum

base shear and the corresponding roof drift for eachearthquake simulation test, was superposed on the graphs

of the pushover test and analyses in Fig. 15. It can be

seen from this figure that the markers, in general, complywith the curve given by the analysis PUSH-II. The dis-

crepancy in the initial part between these markers (or

the curve of the analysis PUSH-II) and the curveobtained through pushover test implies the stiffness

degradation which had occurred due to the damages

implemented during the course of previous earthquakesimulation tests. The reason for the discrepancy in the

latter part between the curves of the pushover test and

analysis PUSH-II is considered the strain-aging effect[16,17] in the model reinforcement. Erasmus [17]

showed that the effect of strain aging can induce the

increase of more than 75% in the discontinuous yieldingstress. In other words, the model structure showed the

clear yielding phenomenon for the earthquake simulationtest of TFTF04. Additionally, the time that elapsed

between the earthquake simulation tests and the push-

over test was more than one year. These two facts musthave affected the increase of the yielding strength of the

model reinforcement through the effect of strain aging.

However, the hollow markers, which indicate themaximum base shear and the corresponding roof drift

derived from the time history nonlinear analyses of

TFTF012 and TFTF04 by IDARC-2D, deviate signifi-cantly from the results of the earthquake simulation test

and pushover test. This means the dynamic nonlinear

analysis does not necessarily provide a higher level ofaccuracy or reliability than the static nonlinear analysis,

as is commonly expected.

With the assumption that the results of analysisPUSH-II would be the actual behavior of the model

structure if the model structure had not experienced the

previous earthquake simulation tests and the effect ofstrain aging, the displacement ductility ratio of the model

structure turns outs to be approximately 2.4 as shown inFig. 15 when the concept of equal energy and the elasto-

perfect-plastic (EPP) bilinear model are adopted. The

effective yielding strength appears to be 40.0 kN inthis figure.

The maximum interstory drift at the first story at the

collapse state in pushover test was 28.7 mm (3.7%) as


Fig. 16. Story shear versus interstory drift in pushover test and analy-

sis.

shown in Fig. 16. The increase of drift at only the first

story after the strength degradation indicates that a fail-ure mechanism was formed in the first story. The dis-

placement ductility ratio at the first story turns out to be

about 1.6 for the ultimate strength, and 2.3 at the timeof collapse. In case of analysis PUSH-II, these values

should be replaced by 2.4 and 3.4, respectively, if the

deformation capacity of the model structure is assumedto be the same as that obtained from the experiment.

This clearly represents the characteristics of the RCframe with nonseismic detailing.

Fig. 17 depicts the development of cracks in experi-

ment. The final distribution of plastic hinges observedfrom experiment, and those obtained through two analy-

ses (PUSH-I and PUSH-II), are shown in Fig. 18(a–c),

respectively. The collapse mechanism in the experimentis the soft-story mechanism (Mechanism 1), while that

of analysis is Mechanism 2 as shown in Fig. 19. Simple

plastic analyses were performed for the two collapse

Fig. 17. Development of cracks in pushover test.

mechanisms (Mechanism 1 and Mechanism 2), assuming

the section properties as given PUSH-I and PUSH-II.With the section properties of PUSH-II, the two collapse

mechanisms have a difference of only 0.7 kN (1.8%) in

strength, which is almost negligible when the uncertaintyin the material and section properties are considered.

5.3. Local reponses

Fig. 20(a) shows the maximum angular rotations in

the ends of members at the collapse state. Angular

rotations in the ends of columns are within the range ofapproximately 0.04 rad and the largest value was

obtained at the bottom of the exterior column located at

the longer span of the first story. The failure occurredin the mode of the flexural compressive crushing in the

upper end of the interior column at the first story, and

the experiment was stopped at this point. In the beam,the largest value of 0.026 rad was observed in the end

of the longer-span beam adjacent to the exterior joint of

the second floor. Fig. 21 shows the histories of angularrotations in the ends of beams and columns in the push-

over test. The ends of columns (1 and 2) have rotations

approximately twice larger than those of beams (3 and4) under the same load at the interior joint. The end of

beam (5) exhibits larger rotations than the ends of beams(3 and 4). The member ends (1, 6, and 7) deformed sig-

nificantly more even after the strength drops. This

means, again, that the deformations were concentratedat the first story after the strength drop.

The rotational angles in the member ends at the time

of the ultimate strength (not at the collapse state) inexperiment (roof drift = 47.2 mm) are recorded in Fig.

20(b) to compare with those in analysis PUSH-II. The

general trend is that the value obtained through analysis


Fig. 18. Distribution of plastic hinges at the ultimate collapse state in pushover test and analysis. (a) Experiment; (b) Analysis (PUSH–I); and

(c) Analysis (PUSH-II).

Fig. 19. Collapse mechanism and collapse load in pushover test and analyses.

overestimates the angles in columns while it underesti-

mates those in girders. However, it should be noted thatthe rotational angles in girders obtained by the analysis

are very small when compared with those given by the

experiment in the case of the interior joint of the secondfloor (analysis (rad)/experiment (rad) = 0.0025/0.014

and 0.0048/0.013).

6. Analysis of overstrength

The overstrength factor is defined on the level of the

whole structure as a ratio of the actual structural yieldlevel to the code-prescribed strength demand arising

from the application of prescribed loads and forces

[18,19]. As shown in Fig. 22, the overstrength coef-

ficient, V, based on the global behavior of a structure is

defined as follows:

V 5Cy

Cw(2)

where Cy is the base shear coefficient corresponding to

the idealized yield displacement Dy of the structure using

the concept of equal energy; Cw is the code-prescribedunfactored design base shear coefficient. This factor can

be further decomposed into two factors, Vs and Vy, asfollows:

V 5Cs

Cw×Cy

Cw5 Vs × Vy (3)

where Cs is the base shear coefficient corresponding to

the first significant yield of the structure.


Fig. 20. Angular rotations (unit: 1024 rad). (a) Maximum angular rotations at collapse state in pushover test; and (b) Comparison of angular

rotations at the maximum load.

Fig. 21. Histories of angular rotations in pushover test. (a) Column ends; and (b) Beam ends.

Fig. 22. Typical global structural response idealized as linearly elas-

tic-perfectly plastic curve.

Bertero et al. [20] indicated that it is due to the over-strength that buildings designed according to the code

could resist severe seismic excitations. Shahrooz and

Moehle [21] showed through an experimental study on

a 1/4-scale, 6-story RC structure that a structure

designed for an unfactored base shear coefficient of

0.091 could theoretically resist 7.5 times as much. Mir-

anda and Bertero [22], based on the study of low-rise

buildings in Mexico City, have noted the value of over-

strength in the range of 2–5. Meli [23] further showed

that the available overstrength varies widely depending

on the type of structure and on the characteristics of the

ground motion. In the studies on the variation of over-

strength factor for a typical interior frame of an office

building located in a region of high seismic risk, Uang

[19] showed that low-rise buildings usually have higher

overstrength, while the overstrength is not very sensitive

to the number of bays.

Most experimental and analytical research in earth-

quake engineering is focused on high-risk seismic zones.

While formulating the design codes, debate is normally

focused on the seismic coefficient for higher zones; the

coefficient for lower zones is simply prorated in pro-

portion to the expected ground motion intensity in differ-

ent zones. So, Jain and Navin [24] carried out a study


on the seismic overstrength of four-bay, three-, six-, and

nine-story RC frames designed for seismic zones I to Vas per Indian codes. The overstrength against lateral load

is very significantly affected by the factored gravity

loads used in design. This results in overstrength beingmuch higher for low seismic zones, for low-rise build-

ings, and for higher design live load. They showed that

the overstrength of a three-story interior frame in zoneV is 3.3, while it is as high as 15.0 in zone I.

The overstrength factor, V, of the model structure canbe demonstrated by calculating both Vs 5 Cs /Cw and

Vy 5 Cy /Cs with respect to the flexural moment

capacity as follows:First, the coefficient, Vs, can be calculated through the

linear elastic analysis of the model structure up to the

occurrence of the first plastic hinge. It was found throughlinear elastic analyses that the model structure can meet

the flexural moment demands under the load cases of

1.4D+1.7L and 0.75(1.4D+1.7L±1.87E), with the mini-mum margin of safety being 25% and 34%, respectively.

However, since the gravity load condition during the

earthquake simulation tests can be described as 1.0D, theratios of the demanded flexural moment to the capacity

for the load case, 1.0D, are recalculated. Then, by com-

paring the demanded flexural moment for the load caseof earthquake (1.0E) to the reserved flexural capacity

(Capacity—1.0D), we can obtain the coefficient, Vs,

which is the least value among the values of SS12

1.0D

CapacityD / ±

1.0E

CapacityD as shown in Fig. 23. This value,

Vs, appears to be 5.06 and this is similar to the ratioVs 5 Cs /Cw derived from pushover analysis,

(24.33 kN) / (4.61 kN) 5 5.28, as shown Fig. 19.

Secondly, Vy 5 Cy /Cs can be calculated in case ofthe experiment (with strain aging) Cy /Cs 5

(51.35 kN) / (24.33 kN) 5 2.11 and, in case of the analy-

Fig. 23. Effective earthquake load factor for first significant yield at

critical member ends.

sis (without strain aging) Cy /Cs 5

(40.00 kN)/ (24.33 kN) 5 1.64 in Fig. 19. Therefore, theoverstrength coefficient, V, can be obtained by multiply-

ing these two coefficients as 11.1 (with strain aging) or

8.7 (without strain aging). These large values of theoverstrength coefficient account for the reason why the

low-rise RC building structures have the large reserved

strength for severe earthquakes even though they weredesigned only for the gravity loads in the lower seis-

mic zones.

7. Conclusions

1. Though the model structure in this study was

designed only for the gravity loads in zones of lowseismicity, the structure could resist not only the

design earthquake, which it would be supposed to

resist if it were to be designed against earthquake, butalso the higher levels of the earthquake excitations.

The main components of its resistance to the high

level of earthquakes appear to be (1) the high over-strength, (2) the elongation of the fundamental period,

(3) the minor energy dissipation by inelastic defor-

mations, and (4) the increase of the damping ratio.The design base shear derived from the linear elastic

base shear of the structure divided by the response

modification factor, R 5 3.5, seems to be completelyfictitious or misleading because the high overstrength

factor, V 5 8.7, implicit in the structure due to thepre-existing overstrength of the materials and section

properties and the reduction in the dead and live loads

with regards to the reactive weight caused the modelstructure behave entirely linear elastically under the

design earthquake. Therefore, as far as this study

alone is concerned, it is more reasonable that the con-cept of the reduction of the design base shear con-

sidering the energy dissipation by the inelastic

response under the design earthquake be waived forthe low-rise building structures in the low-seismicity

regions. However, considering the possibility of

unexpected large earthquakes, the structures in low-to-moderate seismicity regions should retain the duc-

tility to some extent, which can be achieved through

the implementation of the requirements on thedetailing of reinforcement and structural layout of

important lateral-load-resisting elements.

2. The nonlinear static and time history analyses becomemore widely used tools nowadays than in the past due

to the development of more advanced design [25] andevaluation [26] procedures utilizing these techniques.

Here, the reliable prediction of the demand and supply

in the global and local responses of the structure mustbe a crucial factor for the successful design and evalu-

ation. However, the available experimental data to

calibrate the existing nonlinear analysis softwares


seem to be rather scarce and lack to ensure their

reliability. The correlation study between analysis andexperiment performed herein in this context reveals

that special care should be taken, when the IDARC-

2D is used, in the interpretation of the analyticalresults since the time history analyses sometimes

underestimate significantly the global responses and

both time history and static analyses generally tendto underestimate the rotations in some ends of beams

though the static pushover analyses give the reliableglobal responses. Efforts to improve the reliability of

the software should be made from the side of

developing more sophisticated analytical models aswell as from the side of calibrating these developed

models through the correlation study whenever more

experimental data become available.3. The new points in this study, which have not been

addressed in the previous studies on the nonseismic

RC frames [2–5], are as follows:O The detailing practice in the model structure as

shown in Fig. 1(h) is different from that of ACI

318 for the nonseismic regions. The exterior jointsin the model structure revealed no significant bond

slippage and premature strength degradation due

to the failure of anchorage, which were generallyobserved in the American researches [3,5].

O The global ductility ratio and the overstrength fac-

tor of this model structure turn out to be 2.4 and8.7, respectively, though the previous researches

[2–4] did not clarify specifically the capacities ofnonductile RC building structures such as the ulti-

mate strength and deformability of nonductile RC

frames through the experiment.O The previous correlation studies generally concen-

trated on the global responses only [2–4]. How-

ever, the correlation study herein covers the localresponses as well as the global responses. In gen-

eral, the local responses predicted by using

IDARC-2D appear to be not so reliable as the glo-bal responses.

Acknowledgements

The research stated herein was supported by the Min-istry of Construction and Transportation, the Republic

of Korea, and several private companies including SSan-

gYong Engineering and Construction Corp., DongBuCorp., Hyundai Construction Corp., and DongYang

Structural Safety Consultants Corp. The contributions ofgraduate students at Korea University, Dong-Woo Ko,

Yun-Sup Heo, and Kyi-Yong Kang, were crucial to the

success of this research. The technique which is essentialto the manufacture of the reduced-scale model was

developed by the support of advanced STructure

RESearch Station (STRESS) at Hanyang University. The

authors express their deepest gratitude to all of these

supports.

References

[1] Jo BG, Kim SK et al. Evaluation of intensity of 13 December

1996 Yeongweol earthquake and attenuation properties of Korean

Peninsula. In: Proceedings of the EESK Conference—Spring

1997. Seoul: Earthquake Engineering Society of Korea; 1997. p.

21–6 (in Korean).

[2] Bracci JM, Reinhorn AM, Mander JB. Seismic resistance of

reinforced concrete frame structures designed only for gravity

loads: part 1—design and properties of a one-third scale model

structure. Technical report NCEER-92-0027, State University of

New York, Buffalo, 1992.

[3] Bracci JM, Reinhorn AM, Mander JB. Seismic resistance of

reinforced concrete frame structures designed for gravity loads:

performance of structural system. ACI Structural Journal

1995;92(5):597–609.

[4] El-Attar AG, White RN, Gergely P. Behavior of gravity load

designed reinforced concrete buildings subjected to earthquake.

ACI Structural Journal 1997;94(2):133–45.

[5] Aycardi LE, Mander JB, Reinhorn AM. Seismic resistance of

reinforced concrete frame structures designed only for gravity

loads: experimental performance of subassemblage. ACI Struc-

tural Journal 1994;91(5):552–63.

[6] Building code for structural concrete (ACI 318-95) and commen-

tary (ACI 318R-95). Detroit (MI): American Concrete Institute,

1995.

[7] Lee HS, Woo SW et al. Manufacturing technique and material

properties for 1/5-scale reinforced concrete frame model. In: Pro-

ceedings of Autumn Conference. Korea Concrete Institute; 1997.

p. 575–80 (in Korean).

[8] Lee HS, Woo SW. An experimental study on the similitude in

structural behaviors for small-scale modeling of reinforced con-

crete structures. In: Proceedings of the 6th US National Confer-

ence on Earthquake Engineering. Oakland, CA: EERI; 1998.

[9] Valles RE, Reinhorn AM, Kunnath SK, Li C, Madan A. IDARC

2D version 4.0: a program for the inelastic damage analysis of

buildings. Technical report NCEER-96-0010, State University of

New York, Buffalo, 1996.

[10] Harris HG, Sabnis GM. Structural modeling and experimental

techniques. Boca Raton, FL: CRC Press, 1999.

[11] Hidalgo P, Clough RW. Earthquake simulator study of a

reinforced concrete frame. Report no. UCB/EERC-74/13, Earth-

quake Engineering Research Center, University of California,

1974.

[12] Regulations concerning the structure standards of building, etc.

Republic of Korea: Ministry of Construction and Transpor-

tation, 1996.

[13] UBC, Uniform building code 1994. International Conference of

Building Officials, 1994.

[14] Felber, AJ. RESPONSE: a program to determine the load–defor-

mation response of reinforced concrete sections. Department of

Civil Engineering, University of Toronto, 1990.

[15] Park YJ, Reinhorn AM, Kunnath SK. IDARC: inelastic damage

analysis of reinforced concrete frame–shear–wall structures.

Technical report NCEER-87-0008, State University of New York

at Buffalo, 1987.

[16] Booth E, editor. Concrete structures in earthquake regions: design

and analysis. New York: John Wiley and Sons; 1994.

[17] Erasmus LA. Cold straightening of partially embedded reinforc-

ing bars—a different view. Concete International: Design and

Construction 1981;3(6):47–52.


[18] Bertero VV. Ductility based structural design, state-of-the-art

report. In: Proceedings of 9th World Conference on Earthquake

Engineering, Tokyo, Kyoto, vol. VIII, 1989, pp. 673–86.

[19] Uang CM. Establishing R (or Rw) and Cd factors for building

seismic provisions. Journal of Structural Engineering ASCE

1991;117(1):19–28.

[20] Bertero VV, Anderson JC, Krawinkler H, Miranda H. Design

guidelines for ductility and drift limits: review of the state-of-

the-practice and state-of-the-art in ductility and drift-based earth-

quake-resistant design of buildings. Report no. UCB/EERC-

91/15, Earthquake Engineering Research Center, University of

California, Berkeley, CA, 1991.

[21] Shahrooz BM, Moehle JP. Evaluation of seismic performance of

reinforced concrete frames. Journal of Structural Engineering

ASCE 1990;116(5):1403–22.

[22] Miranda E, Bertero VV. The Mexico earthquake of September

19, 1985—performance of low-rise buildings in Mexico City.

Earthquake Spectra 1989;5(3):571–90.

[23] Meli R. Code-prescribed seismic actions and performance of

buildings. In: Proceedings of the 10th World Conference on

Earthquake Engineering, X. Rotterdam: A.A. Balkema; 1992. p.

5783–8.

[24] Jain SK, Navin R. Seismic overstrength in reinforced concrete

frames. Journal of Structural Engineering ASCE

1995;121(3):580–5.

[25] Kappos AJ, Manafpour A. Seismic design of R/C buildings with

the aid of advanced analytical techniques. Engineering Structures

2001;23(4):319–32.

[26] FEMA (Federal Emergency Management Agency, US). NEHRP

recommended provisions for seismic regulations for new build-

ings and other structures (1997 ed.): part 1—provisions; part 2—

commentary. Washington, DC, BSSC, 1998.

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