shake table

EARTHQUAKE ENGINEERING AND STRUCTURAL DYNAMICSEarthquake Engng Struct. Dyn. 2006; 35:1827–1852Published online 2 August 2006 in Wiley InterScience (www.interscience.wiley.com). DOI: 10.1002/eqe.612

Shake-table experiment on reinforced concrete structurecontaining masonry infill wall

Alidad Hashemi‡ and Khalid M. Mosalam∗,†,§

Department of Civil and Environmental Engineering, University of California,Berkeley, CA 94720-1710, U.S.A.

SUMMARY

A hypothetical 5-storey prototype structure with reinforced concrete (RC) frame and unreinforced masonry(URM) wall is considered. The paper focuses on a shake-table experiment conducted on a substructureof this prototype consisting of the middle bays of its first storey. A test structure is constructed torepresent the selected substructure and the relationship between demand parameters of the test structureand those of the prototype structure is established using computational modelling. The dynamic propertiesof the test structure are determined using a number of preliminary tests before performing the shake-tableexperiments. Based on these tests and results obtained from computational modelling of the test structure,the test ground motions and the sequence of shakings are determined. The results of the shake-tabletests in terms of the global and local responses and the effects of the URM infill wall on the structuralbehaviour and the dynamic properties of the RC test structure are presented. Finally, the test results arecompared to analytical ones obtained from further computational modelling of the test structure subjectedto the measured shake-table accelerations. Copyright q 2006 John Wiley & Sons, Ltd.

Received 22 April 2006; Revised 13 June 2006; Accepted 13 June 2006

KEY WORDS: earthquakes; infilled frame; modelling; reinforced concrete; shake-table; URM wall

INTRODUCTION

Complex structures with multiple dissimilar components (hybrid systems) are frequently builtin seismically active regions. Examples include reinforced concrete (RC) building frames withunreinforced masonry (URM) infill walls or steel bridge decks supported on RC piers. In order

∗Correspondence to: Khalid M. Mosalam, 721 Davis Hall, University of California, Berkeley, CA 94720-1710, U.S.A.†E-mail: [email protected]‡Ph.D. Candidate.§Associate Professor.

Contract/grant sponsor: National Science Foundation; contract/grant number: CMS0116005

Copyright q 2006 John Wiley & Sons, Ltd.

1828 A. HASHEMI AND K. M. MOSALAM

to develop new modelling techniques and study the behaviour of RC buildings with URM infillwalls, a two-phase experimental and analytical study is conducted.

Masonry infilled frames have been experimentally investigated for both in-plane and out-of-planeforces by many researchers. Most of these studies are focused on the behaviour of single-framesingle-bay URM infilled frames under monotonic or cyclic lateral loading. Earlier studies can befound in Reference [1] among others. Some exceptions where multiple-bays and multiple-stories aretested can also be found in literature [2, 3]. These studies provide evaluations of the importance ofinfill confinement from bounding frames, the types of failure that can be observed in the infill or thebounding frame, the stiffness and strength of the infilled frames, and the degradation of strengthupon load reversals. Since these tests are performed using monotonic, quasi-static, or pseudo-dynamic loading, it is not clear how well they represent the dynamic properties, e.g. dampingcharacteristics, of a structure with masonry infilled frames subjected to earthquake loading.

Limited data are available on dynamic properties of masonry infilled frames since very fewshake-table experiments are performed on such structures. Fardis et al. [4] report on a shake-table test performed on single-bay two-storey RC frames with eccentric (non-symmetric in plan)masonry infills subjected to bi-directional ground acceleration. The study focused on the effectsof the eccentricity on displacement demands on the corner columns. Zarnic et al. [5] report ontwo shake-table tests performed on 1

4 -scale one and two-storey RC frames with strong-blockweak-mortar masonry infill walls subjected to one-directional sinusoidal motion at the base of thestructure. Dolce et al. [6] report on shake-table tests performed on two-dimensional 1

3.3 -scale three-storey two-bay RC frames designed for low seismicity regions without infill, with masonry infilland with two different types of energy dissipating and re-centring braces. The study comparedthe overall response and the dynamic properties of the three frames subjected to a sequenceof artificially generated accelerograms with increasing intensity. The mentioned experiments andothers are generally performed on small-scale models due to the size limitations of the shake-tablesand are focused on different aspects of the problem, e.g. torsional effects due to eccentric infillwalls. The current experiment is conducted to study the dynamic performance of a seismicallydesigned symmetric large-scale URM-infilled RC frame subjected to real ground motions. Thestudy focuses on evaluating the effects of the URM infill wall on the surrounding structuralelements, namely the RC slab and the RC columns. Moreover, since the structure consists of twoframes without infills and an infilled frame, comparing the frame responses and quantifying thedistribution of lateral forces between the tested frames before and after infill damage is an importantand novel objective of the study. The current study is unique in the aspect of testing a structurewith dissimilar frames. Accordingly, the presented shake-table experiment can be viewed as abenchmark for experimentation using mixed-variables (force and displacement) pseudo-dynamictechnique with substructuring [7] and for analytical modelling of URM infilled frames.

The shake-table experiment is carried out on a reduced-scale one-storey RC moment-resistingframe structure with URM infill wall on the seismic simulator test facility of the Universityof California, Berkeley. The 3

4 -scale test structure represents the first storey middle bays of a5-storey RC prototype structure designed based on the requirements of ACI318-02 [8] and NEHRPrecommendations [9] in seismic regions. URM walls are assumed in the interior frames, Figure 1.

The experimental study serves the purpose of calibrating analytical models being developedusing Open System for Earthquake Engineering Simulation (OpenSees) [10]. The objectives ofthe modelling effort are to enable accurate representation of the in-plane behaviour of URM infillwalls, and to refine the modelling techniques of hysteretic strength and stiffness degradation inRC elements and joints of RC moment frames interacting with URM infill walls.

Copyright q 2006 John Wiley & Sons, Ltd. Earthquake Engng Struct. Dyn. 2006; 35:1827–1852DOI: 10.1002/eqe

SHAKE-TABLE EXPERIMENT ON REINFORCED CONCRETE STRUCTURE 1829

C2 C3

A3A0 A2A1

B0 B1 B3B2

(a) (b)

N

Figure 1. Development of the shake-table test structure: (a) prototype structure;and (b) test structure on the shake-table.

PROTOTYPE STRUCTURE

The 5-storey prototype moment-resisting frame structure is designed with its exterior columns(A0, A3, B0, B3, C0, and C3), as the primary lateral load resisting system. Although it is commonto have masonry walls on the perimeter of buildings, the URM walls are assumed in the middleframes for two practical reasons. First, having two walls on the perimeter for the test structurewould have required about twice as much shear force to damage the walls, thus exceeding thecapacity of the shake-table. Second, the failure of the two URM walls would not have beensimultaneous due to inherent material and construction variability and that would have caused asignificant shift in the centre of rigidity of the test structure compared to its centre of mass andproduced large torsional demands with loading patterns out of scope of the study. Due to thelimited size of the available shake-table, the prototype structure is scaled to 75% of its originaldesign size. A typical floor plan of the 3

4 -scaled prototype building is shown in Figure 2. Uniformlydistributed superimposed dead and live loads on the scaled prototype structure (DL and LL) areindependent of the length scale factor. For each floor, the DL and LL are 160 Pa (110 psf) and 74 Pa(50 psf), respectively, and for the roof, they are 130 Pa (90 psf) and 15 Pa (10 psf), respectively. Inthe following text, the scaled prototype structure is referred to as the prototype structure.

The prototype substructure is selected as the middle bays of the first storey of the prototypebuilding as shown in Figure 2. Although these bays are typically not a part of the primary lateral loadresisting system, they are detailed as such because they would have to endure the same displacementdemands in an earthquake event and maintain their axial load carrying capacity. The test structureis designed to represent the prototype substructure as described before with the exception that thetransverse span is reduced to 1

2 its prototype length due to the size limitation of the shake-table.In order to determine the boundary conditions and necessary adjustments for this representation,an analytical model of the prototype building is constructed using OpenSees, Figure 3. Beamsand columns are modelled using nonlinearBeamColumn element in OpenSees, which is based onforce formulation, and considers the spread of plasticity along the length of the element. Sectionsare defined using fibre discretization with distinct fibres for longitudinal reinforcement. Concretematerial is modelled using Concrete01, which is a uniaxial concrete material object with degradedlinear unloading/reloading stiffness in compression and no tensile strength. Confining effect due



0 1 2 3

A

B

C

230x305

C530x530

C305x305

URM

INFILL230x305

340x265

TY

P.

TY

P.

340x265 340x265

TYP. TYP. TYP.

411041104110

Experiment

substructure

3660

3660

Transverse

direction

Longitudinal direction

Figure 2. Floor plan of the 34 -scaled prototype structure (dimensions in mm).

Figure 3. Computational model.

Table I. Concrete model properties.

Column ColumnProperty Foundation Beam cover core

Peak compressive 34.4 (4.98) 37.2 (5.39) 38.4 (5.56) 45.3 (6.57)stress (MPa (ksi))Strain at peak 0.002 0.002 0.002 0.004compressive stressUltimate strain 0.006 0.006 0.006 0.020Stress at ultimate 0 0 0 6.90 (1.00)strain (MPa (ksi))

to the prescribed transverse reinforcement is accounted for using confined concrete properties forcolumn core concrete material [11, 12]. Steel reinforcing bars are modelled using Steel01, whichis a uniaxial bilinear material object with kinematic hardening. The material properties used forthe elements of the model are defined in Tables I and II. The column–footing joints are modelledusing the recommendations of FEMA 356 [13] by a tri-linear moment-rotation relationship as



Table II. Steel model properties.

Property Parameter

Yield stress (MPa (ksi)) 458 (66.5)Yield strain (rad) 0.0023Modulus of elasticity (GPa (ksi)) 200 (29 000)Kinematic hardening ratio 0.01

Y Pcr

MY

Mcr

MP

M

Figure 4. Column–footing joint model.

Table III. Column–footing joint model properties (refer to Figure 4).

Mcr (kNm (kip in)) 29.9 (265)�cr (rad) 0.002MY (kNm (kip in)) 130 (1150)�Y (rad) 0.015MP (kNm (kip in)) 158 (1400)�P (rad) 0.030

mo mu

mof

muf

Strut

Tension Compression

ParabolaStraight line

Figure 5. URM infill strut model.

shown in Figure 4 and Table III. The masonry infill wall is modelled using equivalent diagonalcompression-only struts as shown in Figure 5 and Table IV. For the 102mm (4′′) thick URMwall, the cross-sectional area of the equivalent strut is estimated using FEMA 356 guidelines as462 cm2 (71.6 in2). This value proved inconsistent with the test results and later modified basedon the estimated stiffness and strength of the infill panel during snap-back and shake-table tests



Table IV. Masonry strut properties (refer to Figure 5).

fmo (MPa (ksi)) 17.0 (2.46)�mo 0.0028

fmu (MPa (ksi)) 1.99 (0.29)�mu 0.0041

0 500 1000 1500-300

-150

0

150

300

450

600

750

900

Moment [kip-in]

Axi

al lo

ad [

kip]

0 50 100 150 200

-1000

0

1000

2000

3000

4000

[kN

]

[kN-m]

Prototype substructure

Teststructure

Figure 6. Range of change in axial load of the middle columns.

to 223 cm2 (34.6 in2). The concrete slab is modelled using horizontal elastic truss members. Thematerial properties and the area of these truss members are determined such that they have thesame stiffness as the concrete slab, which is expected to remain elastic during the test.

Using the OpenSees model, a non-linear time history analysis of the prototype structure sub-jected to different levels of selected ground motion is performed. These levels are discussed insubsequent sections. Analysing the results, the axial load and base shear affecting the prototypesubstructure are determined for each level of input ground motion. With the same assumptionsas the prototype structure, an OpenSees model of the test structure is analysed and the requiredamount of additional mass is determined for the test structure. This determination is based onmatching the computationally determined base shear of the test structure to that of the prototypesubstructure when subjected to the design-level ground motion.

In order to explore the effects of the column axial loads due to the weight of the upper stories inthe prototype structure, the range of change in the axial force during each level of ground motionis superposed on the moment–axial interaction diagram for the middle columns. An exampleof such results for the design-level ground motion is shown in Figure 6. The static axial forceon the column due to upper storey dead and live loads is 386 kN (86.8 kips) corresponding to160 kNm (1420 k in) moment capacity for the column section. During the design level groundmotion, the axial load ranges from 220 (49.5 kips) to 429 kN (96.5 kips) and the correspondingmoment capacity of the section ranges from 144 (1270 k in) to 164 kNm (1450 k in). Comparingthis variation with the moment capacity of the section at zero axial force, 120 kNm (1060 k in),



0 0.2 0.4 0.6 0.80

50

100

150

Corresponding displacement [in]

Max

imum

bas

e sh

ear

[kip

s]

Prototype substructureTest structure

0 5 10 15 20

0

100

200

300

400

500

600

[kN

]

[mm]

0

50

100

150

200

250

Time [sec]

Hys

tere

tic e

nerg

y [k

ip-i

n]

Prototype substructureTest structure

0

5

10

15

20

25

[kN

-m]

(a) (b)

120100806040200

Figure 7. Comparison between prototype substructure and test structure: (a) max. base shear versuscorresponding displacement; and (b) cumulative hysteretic energy.

it is concluded that the axial load on the column has significant effect on its flexural capacity(up to 37% increase) of the section. To accommodate this observation for the test structure, thestatic column axial load due to weight of the upper stories is applied to the columns in the form ofunbonded concentric prestressing. The resulting range of change in the axial load in the columnscorresponding to the design level ground motion for the test structure with the axial prestressingis also obtained using OpenSees model and shown to be very close to that of the prototype asshown in Figure 6.

Comparison between the response of the prototype substructure and the test structure when sub-jected to different levels of selected input ground motion is performed using non-linear time historyanalyses. The plots for maximum base shear versus its corresponding first storey displacementfor both the prototype substructure and the test structure for different levels of the input groundmotion are shown in Figure 7(a). The cumulative hysteretic energy plots for these consecutive runsare compared for the two structures in Figure 7(b). These results show reasonable agreements,e.g. less than 13% mismatch for the cumulative hysteretic energy at the end of all runs, betweenthe response of the prototype substructure and the test structure.

As mentioned earlier, by providing additional mass and prestressing in the test structure, anacceptable match between the responses of the two structures is achieved. However, the shortcom-ings are due to the effects of the overturning moment and higher modes in the prototype structure.In general, the overturning moments have considerable impact on the axial force of the columnson the perimeter of the structure. These effects are less significant when one considers the middlecolumns. Preliminary analysis of the prototype structure shows that the variation of axial loadin the first storey columns due to the overturning moments is about 175% for perimeter column(e.g. column A0) and only about 15% for middle columns (e.g. column A1). To explore the effectsof higher modes on the prototype structure, a modal response spectrum analysis [14] for the pro-totype structure subjected to the NEHRP design spectra is performed. The resulting modal baseshears and first storey displacements are compared and it is concluded that the effects of highermodes is limited to about 10% of the first mode effects. Accordingly, the effects of overturningmoments and higher modes are neglected when relating the test structure to the prototype.



CONFIGURATION AND INSTRUMENTATION OF THE TEST STRUCTURE

The overall dimensions of the test structure are 4.88m× 4.42m (16′-0′′ × 14′-6′′) in plan and3.43m (11′-3′′) in height, Figure 2. Member sizes and reinforcement details for different ele-ments of the test structure are summarized in Table V. The reinforcement is specified as ASTMA615 [15] Grade 60. The yield stress of the 19mm diameter (#6) bars is 458MPa (66.5 ksi) permill certification. The specified 28-day compressive strength of the standard concrete cylinder perASTM C 837-99 [16] is 31MPa (4.5 ksi). The masonry wall is made of clay bricks with modularsize of 102mm× 203mm× 68mm (4′′ × 8′′ × 22

3′′) and ASTM C270 [17] Type N mortar. The

measured average 28-day compressive strength of the standard masonry prism according to ASTMC 1314 [18] is 17MPa (2.46 ksi). The average shear strength for 102mm (4′′) thick masonry panelsconstructed and measured in accordance with ASTM E519 [19] is 1.81MPa (263 psi).

Uniformly distributed mass is added to the slab in the form of stacked lead ingots bolted tothe slab using 10mm (38

′′) diameter high strength rods. Static tests confirmed that the friction

forces between the slab and the lead ingots are large enough to accommodate up to 4.0g lateralacceleration at the slab level.

To measure the floor acceleration in three directions, 11 accelerometers are installed on the floorlevel (Figure 8) as follows: in the longitudinal direction, one at each of the six columns and oneat the middle of the beam in frame B; in the transverse direction, three on the diagonal of theslab on opposite corners and at the centre; and in the vertical direction, one at the middle of the

Table V. Member sized and reinforcement details for the test structure.

Structural Main Transverseelement Dimensions reinforcement reinforcement

Concrete slab 95mm (334′′) thick M10 (#3) top and

bottom @ 305mm(12′′) o.c. each way

None

Columns 305mm× 305mm(12′′ × 12′′)

8–19mm diam.(#6), 32mm diam.(114

′′) prestressing

rod

M10@95mm (#3@334′′) over 610mm

(24′′) from the face of the joints andM10@152mm (#3@6′′) elsewhere

Longitudinal beams(single span)

267mm× 343mm(1012

′′ × 1312 )3–19mm diameter(#6) top and bottom

M10@70mm (#3@234′′) over the

711mm (28′′) from the face of thebeam-column joint and M10@203mm(#3@8′′) elsewhere

Short direction beams(double span)

305mm× 229mm(12′′ × 9′′)

2–19mm diameter(#6) top and bottom

M10@305mm (#3@12′′)

Footing 356mm× 457mm(14′′ × 18′′)

4–22mm diameter(#7) top and bottom

M10@102mm (#3@4′′)

Masonry wall 102mm (4′′) thick None None



13'-6

" [4

115

mm

]

2

1

6'-0" [1829mm]A B C

6'-0" [1829 mm]

FLOOR PLAN

D04 D05 D06

D07

D08

A06A04

A05

A07A08

A13

A09

A10

A14

A11

A12

Accelerometer

Displacementtransducer

A B C

1

2

20'-0

" [6

096

mm

] Sq

uare

Shake-tableoutline

19'-0

" [5

791

mm

]

A01

A02

A03

D01

D03

D02

FOUNDATION PLAN

Figure 8. Global instrumentation plan for the test structure.

beam in frame B. Moreover, three accelerometers at the base of the URM wall on the footing (onein each direction) and eight accelerometers built-in the actuators beneath the shake-table (one perhorizontal and vertical actuator) are monitored.

To measure global displacements of the shake-table and the test structure with respect to thestationary ground, eight displacement transducers (Figure 8) are used. Five of them measure thedisplacement of the first floor (three in the longitudinal direction and two in the transverse direction)and the remaining three measure the displacements of the shake-table itself (two in the longitudinaldirection and one in the transverse direction).

To measure local displacements and rotations, 75 displacement transducers are used: nine areinstalled in the plane of the slab, 14 are installed in the plane of the URM wall measuring thediagonal deformations, wall sliding and opening with respect to the bounding frame, four aremeasuring the diagonal displacements of the frames and the rest are installed on four of thecolumns (12 per column) to measure the rotations and average curvatures along the length of thecolumn. Finally, 78 strain gauges out of more than 150 strain gauges installed on the reinforcingbars throughout the test structure are used during each run of the shake-table. As some of the straingauges are damaged during the different test runs, alternative gauges are selected and monitored.The prestressing rods are also gauged and monitored during all test runs. The layouts of theselocal measurements, e.g. Figures 10, 17, and 20, are presented with results in the relevant sections.

SYSTEM PROPERTIES

Pull (snap-back) tests are performed on the test structure before and after the wall construc-tion to determine the stiffness, natural frequency and damping ratio of the structural systembefore starting the shake-table experiment. These tests are conducted for both in-plane and



Table VI. Snap-back test results (refer to Figure 1(c) for orientation of the North (N) direction).

In-plane (North–South direction) Out-of-plane (East–West direction)

Conditions of the test Stiffness Stiffnessstructure at time of the Natural Damping (kN/mm Natural Damping (kN/mmpull (snap-back) test period (s) ratio (%) (kips/in)) period (s) ratio (%) (kips/in))

Before building the wallcolumns not prestressedno additional mass

0.135 4.30 19.8 (113.3) 0.134 4.40 23.5 (134.0)

After building the wallcolumns prestressed noadditional mass

0.055 5.70 74.5 (425.5) 0.122 4.30 29.3 (167.1)

After building the wallcolumns prestressedwith additional mass

0.134 6.85 75.5 (431.0) 0.232 4.25 29.4 (168.0)

Table VII. Ground motion specifications.

Ground motion Station Direction PGA (g) PGV (mm/s (in/s)) PGD (mm (in))

Northridge, CA, 1994 Tarzana 090 1.570 920 (36.23) 130 (5.13)Duzce, Turkey, 1999 Lamont N 0.762 329 (12.97) 19 (0.75)

Table VIII. Scale factors for different levels of input ground motions.

Level 1 2 3 4 6 7 8

Northridge, CA, 1994 (TAR) 0.05 0.17 0.23 0.39 0.59 — —Duzce, Turkey, 1999 (DUZ) — — — — — 1.50 2.00

out-of-plane directions of the test structure, separately. The results of these tests are summarized inTable VI.

GROUND MOTIONS

Two different ground motions as described in Table VII are used in the experiment. In this table,PGA, PGV, and PGD refer to peak ground acceleration, velocity, and displacement, respectively.These ground motions are intended to be unidirectional in the direction parallel to the URMinfill wall of the test structure (longitudinal direction). Each ground motion is scaled to generatedifferent levels of intensity as listed in Table VIII. The scaling is based on the average spectralacceleration of selected ground motions and the NEHRP design spectrum for a site with mappedspectral response acceleration at 1 s, S1 = 0.65g and at short periods, Ss = 1.60g and site classD with 5% damping over the range that the period of structure is expected to vary during theexperiment. For Northridge Tarzana and Duzce ground motions, this range is estimated from the



0 0.2 0.4 0.6 0.8 10

0.5

1

1.5

2

2.5 Begining of the test (TAR 1)

TAR 4

TAR 6

DUZ 7

DUZ 8

After removalof the wall

Spec

tral

acc

eler

atio

n [g

]

Period [sec]

DuzceNorthridgeDesign Spectrum

0 5 10 15

-1

0

1

Northridge, Tarzana

0 5 10 15

-1

0

1

Acc

eler

atio

n [g

]

Time [sec]

Duzce

1.57g

Figure 9. Response spectra (5% damping) for the selected ground motions.

period of the undamaged infilled structure to the period of the structure after removal of the infill(between 0.15 and 0.28 s). The scaled spectra of the two selected ground motions and the designspectrum are shown in Figure 9. In this figure the natural period of the test structure correspondingto important milestones of the experiment are marked. The test structure is subjected to a sequenceof ground motions starting with Northridge Tarzana level 1–6 (denoted by TAR 1 through TAR 6)and Duzce levels 7, 8 and finally 7 again (denoted by DUZ 7, DUZ 8 and DUZ 7-2). Level TAR 1is selected as a small amplitude motion to check the performance of the shake-table and dataacquisition system. Levels TAR 2 and TAR 3 are selected as intermediate intensity levels andlevels TAR 4 and TAR 6¶ correspond to 10/50 (design) and 2/50 (MCE) spectra, respectively.Note that, e.g. 10/50 means 10% probability of being exceeded in 50 years. Levels DUZ 7 and DUZ8 are selected to achieve higher demands on the test structure up to the limits of the shake-table.The ground motion records are compressed in time by

√3/4 factor to account for the reduced

length scale of the test structure. In this way the frequency content of the compressed record atthe natural frequency of the undamaged reduced-scale test structure is the same as that of theun-compressed record at the natural frequency of the test structure without scaling.

SYSTEM IDENTIFICATION

Figure 10 shows the methodology used to find the average damping coefficient and average stiffnessof the test structure during each run and the distribution of forces in different elements of the teststructure. For this purpose, the test structure is idealized as a single degree of freedom (SDOF)system represented by the average floor displacement. The total floor acceleration ut is calculatedby taking the weighted average of the measured accelerations on the floor. Dividing the floor

¶During the last stages of planning for the sequence of ground motion, level TAR 5 with PGA of 0.49g is deemedredundant and is not used for the actual experiment.



System Identification:

ucFF IS ˆ

um um um

uIF m

m

tumtu

6

1icolSwall VFF

topM,top

botbot M,

h

MMV bottop

col

Measurements for column i

top1

top2

bot1

bot2

h

1d

ckgressionucukIF

dtuudtuugutuu

ˆ,ˆRe)(

, 2H

Hh

topcol VV

topV

topM

wallF

2

h

botM

walltop FV

Figure 10. System identification and member shear force calculation in the test structure.

into three strips consistent with the tributary area of each frame and designating the mass andthe measured acceleration associated with the i th strip as mi and uti , respectively, the total flooracceleration ut and the total inertia force acting on the structure FI are determined as follows:

ut =3∑

i=1mi u

ti

/3∑

i=1mi (1)

FI =3∑

i=1mi u

ti (2)

The acceleration of the floor relative to the shake-table u is calculated by subtracting the groundacceleration from ut . The corresponding relative velocity and displacements are determined byintegrating the relative acceleration in time. The dynamic equilibrium equation for the idealizedSDOF system can be written as FI + FD + FS = 0 at each instant of time. With the assumption ofviscous damping, FD = cu is the damping force and FS is the restoring force of the test structure.Assuming a constant damping coefficient c, and a constant average stiffness over the duration ofeach run, the least-square estimates of damping coefficient c and average stiffness k are obtainedusing regression function FI = − (ku+cu) in vector space of (u, u) considering all the data pointsin the duration of each run. Using the estimated value of the damping coefficient, the restoring



force in the structure is calculated from the dynamic equilibrium, i.e. FS = −FI − cu referred toas the total restoring force of the test structure.

The portion of the total restoring force carried by each column can be calculated using the datafrom the strain gauges located at the top and bottom of each column. By making the Bernoulliassumption of plane section remains plane after bending at those locations and knowing thegeometric dimensions, section curvature is calculated as in Equation (3), where d1 is the distancebetween the two strain-gauges in the section. Using strains, curvature and constitutive relationshipsfor the reinforcing bars and concrete, the bending moments at each end of the column segmentbetween the two pairs of strain-gauges (top and bottom) are determined by section analysis. Fromequilibrium, column shear is determined from Equation (4):

� = (�1 − �2)/d1 (3)

Vcol = (Mtop + Mbot)/h (4)

where Mtop, Mbot, and h are defined in Figure 10.For the middle frame columns where there is contact between the URM wall and the column,

the equation for column shear force above the contact length is rewritten as in Equation (5), where�′h is the contact length between the URM infill and the RC column segment bounded by thetwo instrumented sections and F ′

Wall is the horizontal component of the portion of the force inthe URM infill wall that is transferred to the column within the portion of contact length �Hnamely �′h as shown in Figure 10. Since the strain-gauges are located at a section well above thecolumn–footing joint, the values of both �′ and F ′

Wall are relatively small and the second term ofEquation (5) is neglected in the calculations of the middle column shear force.

Vcol = (Mtop + Mbot)/h − F ′Wall(�

′/2) (5)

Finally, the shear force in the URM infill wall is calculated as the total restoring force minus thesum of shear forces in all six columns comprising the test structure RC framing, Figure 10.

TEST RESULTS AND DISCUSSION

The intact test structure with the URM wall is subjected to the sequence of ground motions asdescribed before. The global response of the structure in terms of the overall drift, stiffness anddamping ratio as well as the local response of the different elements of the test structure arequantified and discussed.

Global response

The total base shear versus displacement plots for selected test levels are presented in Figure 11.Total base shear is defined in the following as the sum of restoring and damping forces in thestructure, which from dynamic equilibrium is the inertial forces FI as described before. In thediscussion of the results presented in Figure 11, the stiffness is estimated by the tangent stiffnessof the loading branch evaluated at different times during the experiments. During levels TAR 1and TAR 2, there is no considerable change in the stiffness of the structure. The plot for levelTAR 3 (Figure 11(a)) shows slight reduction in stiffness (about 9% from 75.5 kN/mm (431 kips/in)



-3 -2 -1 0 1 2 3-200

-150

-100

-50

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(a) (b)




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-60 -40 -20 0 20 40 60[mm]

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K=365 kips/in K=391 kips/in

Initial stiffnessK=278 kips/in

K=159 kips/in

K=62 kips/in

K=160 kips/in Initial stiffnessK=364 kips/in

Final stiffnessK=289 kips/in

Initial stiffnessK=116 kips/in

Final stiffness K=63 kips/in

Static frictionK=281 kips/in

Static frictionK=177 kips/in

Final stiffness K=51 kips/in

Figure 11. Total base shear versus lateral displacement for different test levels: (a) TAR 3; (b) TAR 4;(c) TAR 6; (d) DUZ 7; (e) DUZ 8; and (f ) DUZ 7-2.



of the pull test results) to 68.4 kN/mm (391 kips/in) but there is no visible sign of damage inthe test structure. The response of the test structure in level TAR 4 (design level) as depicted inFigure 11(b) shows some drop in the stiffness (about 15% reduction), but the overall behaviourremains almost linear. Close observation of the wall after the completion of TAR 4 reveals smallvisible cracks at the wall–column interfaces.

Figure 11(c), corresponding to level TAR 6, shows the first significant signs of damage. Thestiffness of 63.7 kN/mm (364 kips/in) in the initial motion shifts to 50.6 kN/mm (289 kips/in) atthe peak of the ground motion (21% shift). Observations after the test suggests that some cracksare developing especially along the column-wall interface and some small vertical splitting cracksare observed in the mortar head joints at the wall corners. The maximum total base shear reachedin this stage is 605 kN (136 kips) corresponding to 144% of the total gravity load.

The response of the test structure during level DUZ 7, Figure 11(d), shows the most significantchange in the structural behaviour of the system. The stiffness at the beginning of this level is48.7 kN/mm (278 kips/in). Significant wall cracks with clear pattern and load path definition areformed during this test level. The force–displacement behaviour of the assembly at this point canbe described as a bi-linear relationship. For small displacements (less than about 6mm (14

′′)), the

cracks on the wall open and close without engaging the URM infill wall resulting in an observedlateral stiffness of about 10.9 kN/mm (62 kips/in). This stiffness can be interpreted as the stiffnessof the cracked RC frame before full contact with the URM wall is reached. Once the cracks close,the URM infill wall picks up the load causing further damage in the wall and stiffness increaseto 28.0 kN/mm (160 kips/in). The peak total base shear observed during all stages of the test,namely 756 kN (170 kips) corresponding to 180% of total gravity load, at floor displacement of19.1mm (0.75′′) is produced at this stage of the experiment and right before a major horizontalcrack in the URM infill wall is developed.

Figure 11(e) shows the gradual disintegration of the URM infill wall as the test structure iscycled back and forth in level DUZ 8. The measured stiffness of the test structure at the beginningof DUZ 8 is 49.2 kN/mm (281 kips/in) for small displacements. Comparison between this stiffnessand that of the previous run suggests that at small force demands, the force transferred throughthe wall is not enough to overcome the static friction between the cracked surfaces. Accordingly,at such small forces, the wall, although cracked, acts as a whole increasing the apparent stiffnessof the structural system. Once the force demands at the crack surfaces exceed the static friction (atabout 110 kN (25 kips) corresponding to 27% of total gravity load), the cracked portions start tomove with respect to each other and the stiffness reduces to that of RC frames including crackedURM infill wall, i.e. 20.3 kN/mm (116 kips/in) as shown in Figure 11(e). Note that the stiffnessof the test structure before building the wall (Table VI) is 19.8 kN/mm (113 kips/in). From thispoint on, the URM infill wall is considered as structurally insignificant. As the test structure goesthrough large displacements, the RC frame starts to accumulate damage mostly concentrated atthe bases of the columns. The stiffness of the test structure at the end of this level is reduced to11.0 kN/mm (63 kips/in) suggesting significant damage in the RC frame in addition to the URMinfill wall collapse.

Finally, Figure 11(f ) shows the results obtained from the repeat run of level DUZ 7, i.e.DUZ 7-2. Beyond the static friction at the beginning of the motion where the stiffness is high at avalue of 31.0 kN/mm (177 kips/in), the stiffness of the test structure is 8.9 kN/mm (51 kips/in),i.e. about 20% reduction from that at the end of DUZ 8.

Figure 12(a) shows the change in the ‘effective’ stiffness of the test structure for all levels oftesting plotted along with the measured initial stiffness of the test structure with and without the



TAR 1 TAR 2 TAR 3 TAR 4 TAR 6 DUZ 7 DUZ 8 DUZ 7-20

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/mm

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Test level

Peri

od [

sec]

Cracking of the wall

DUZ 8

Wall removal

Results using ground motion signalsResults using white-noise signals

(a)(b)

Figure 12. Variation of dynamic properties of the test structure: (a) effectivestiffness; and (b) natural period.

URM infill wall. This effective stiffness denotes the average tangent stiffness of the test structureafter the demand force exceeds the initial static friction. It can be observed that the existence ofthe URM wall considerably increases the stiffness of the structural system. As the wall undergoesdamage, the stiffness reduces with most rapid reduction taking place during levels TAR 6, DUZ 7and DUZ 8 suggesting significant disintegration of the wall in these levels. Note that the stiffnessof the test structure reduces to a level less than the stiffness of the elastic structure without thewall, which is due to the damage at column–footing and beam–column joints in the RC framestructure.

Corresponding to the change in the stiffness of the structure, there is significant change in thenatural period of the test structure. In order to identify this change throughout the test runs, a lowamplitude white noise signal with approximately constant small amplitude of 0.07g over frequencyrange of 1-10Hz is applied before each run and the resulting acceleration of the floor is analysedin frequency domain. The period corresponding to the peak amplitude of the frequency response istaken as the natural period of the test structure before each run. Alternatively, the transfer functionof the test structure obtained from the ratio of the Fourier transform of the measured floor and baseaccelerations is examined. The effective natural frequency of the structure during each level of thetest is defined as the frequency corresponding to the peak in this transfer function. Figure 12(b)demonstrates the variations in natural period of the test structure as the shake-table experimentprogressed, which are determined using both the white noise and ground motion signals. Becausethe white noise tests are performed with very low amplitudes, existing cracks in the structureremain closed and the structure appears stiffer. Thus, the resulting natural periods from white-noise tests represent a lower bound for the natural period of the test structure during the actualground motion. Significant elongation in the natural period of the test structure is observed, i.e.from 0.147 s during TAR 1 to 0.392 s during DUZ 7-2 with 167% elongation.

Another indication of the change in the dynamic properties of the test structure due to the induceddamage during the shake-table experiment is the change in the damping ratio. The average dampingcoefficient c during each shake-table run is obtained with the procedure described earlier, Figure 10.The damping ratio is subsequently obtained as �= c/ccr where ccr = 2m�n is the critical damping




2

4

6

8

10

12

14

Test levels

Dam

ping

rat

io [

%]

RegressionEnergy equivalent

Figure 13. Variation of the equivalent damping ratio.

of the test structure where m and �n are the total mass and natural frequency, respectively, of thetest structure. Alternatively, the equivalent damping ratio can be estimated as 1/4� times the ratioof the dissipated energy to the maximum strain energy in each cycle [14]. The resulting variationof the damping ratio � for different levels of shaking calculated using both methods is shownin Figure 13. It should be emphasized that while the regression method, assuming an averagelinear response, only limits the damping to the velocity related term in the equation of motion,the equivalent energy method includes both the viscous damping and the inelastic deformations assources of dissipating energy. Both methods indicate smaller damping ratios (4–6% on average)for the intact structure (levels TAR 1–TAR 4) and large damping ratios (11–13%) where significantdamage occurs in the test structure (levels TAR 6–DUZ 7-2).

Local response

URM infill. The main failure mode of the URM infill wall takes place in level DUZ 7 and ischaracterized by large cracks at 60◦ from the horizontal axis starting from the top corners ofthe wall and connecting with a long horizontal crack at the lower third of the wall to a seriesof 45◦ cracks propagating into the opposite bottom corners along each of the wall diagonals. Atthe same time early signs of corner crushing are observed at the top corners. The markings inFigure 14(a) show the observed crack pattern in the wall after DUZ 7. The sharp angle of crackingat the top corner is attributed to the weaker bond between the upper most mortar bed joint and theRC beam relative to the bond between the side vertical mortar joints and the RC columns. Partialcollapse of the top corners and sides of the wall follows the formation of the crack pattern in levelDUZ 8. The markings in Figure 14(b) show the final crack pattern on the wall at the end of levelDUZ 8. Finally, in level DUZ 7-2 the loose portions of the wall collapsed leading to the damagestate shown in Figure 14(c).

The shear forces carried by the URM infill wall and by each RC column are calculated asdescribed earlier, Figure 10. The shear force carried by the RC frame is the sum of all the shearforces of the six columns. Figure 15 shows the portion of the total base shear which is carried



(a) (b) (c)

Figure 14. Observed damage of the test structure: (a) cracking after Duzce 7; (b) partial collapse afterDuzce 8; and (c) final state after Duzce 7-2.


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(a) (b)

Figure 15. Effect of progression of damage on load sharing between URM infill and RC frames:(a) at peak base shear; and (b) at peak floor displacement.

by the URM infill wall compared to that carried by the three RC frames: at the peak of thetotal base shear (Figure 15(a)) and at the peak of the lateral floor displacement for all testinglevels (Figure 15(b)). The plots confirm that (before level TAR6) the undamaged URM infill wallgoverns the behaviour of the test structure. As the wall experiences damage, the RC frames pickup a larger portion of the load. At level DUZ 7-2, the wall is completely disintegrated and can beconsidered structurally insignificant as the load is carried almost entirely by the three RC frames.The distinction between the force distributions for peak base shear and peak floor displacement ismade to emphasize the different state of the test structure at these different peak points particularlyfor levels DUZ 7 and DUZ 8. In these levels, the point of peak base shear takes place at the timewhere the wall is still resisting large portion of the load at its incipient failure. However, the peakfloor displacement takes place after the damage in the wall has occurred. As an example, the timehistories of the shear forces in the wall and that carried by the RC frames and the correspondingfloor displacements are presented in Figure 16. From this figure, the maximum base shear takesplace at 8.8 s while the maximum floor displacement occurs at 17.3 s following wall crackingat 15.4 s.

An important aspect of the response of the URM infill wall is the shear force versus sheardeformation in the plane of the wall. The shear deformation is determined using the diagonal



7 8 9 10 11 12-130-100

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[mm

]

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displacementMaximum

base shear

Wall cracking

Figure 16. Time histories of shear forces and displacements for level DUZ 7: (a) partial time history from7 to 12 s; and (b) partial time history from 14 to 19 s.

H

L

L H

L2 + H2

2

∆2 − ∆1∆1 ∆2

Figure 17. Shear deformation of the URM infill wall.

measurements in the plane of the wall as illustrated in Figure 17. The shear force in the wall, FW ,is estimated as discussed earlier, Figure 10. The shear force versus shear deformation plot for theURM infill wall for levels TAR 6 and DUZ 7 are shown in Figure 18. Before cracking, i.e. levelTAR 6, a linear shear force versus shear deformation relationship is obtained. After cracking, i.e.level DUZ 7, there is a rapid degradation of the shear stiffness of the URM infill wall due tosignificant increase in the shear deformation.

RC slab. The 95mm (334′′) thick RC slab is supported on boundary beams from all sides and spans

between the bare and infilled frames. In its plane, the RC slab acts as a diaphragm distributing theinertia force to the lateral resisting elements of the test structure by both deforming in shear andin-plane bending. The inertia force is generated by the acceleration of the mass at the slab level inthe test structure consisting of the tributary mass of the test structure and the added mass of thelead ingots.

In the out-of-plane direction, the slab spans between the three frames. Since the aspect ratio ofthe slab is 2.25, the out-of-plane behaviour of the slab can be characterized as one-way action.This is confirmed by observing the gravity induced crack patterns on the slab after loading thelead ingots. These cracks run more or less parallel to the long edges of the slab.

While the vertical forces on the slab hardly change during the shake-table test, the lateral inertiaforces acting on the slab and the resisting reactions from the frames beneath change dramatically



-8.5 -7 -5 -3 -1 0 1 3 5 7 8.5-8.5 -7 -5 -3 -1 0 1 3 5 7 8.5

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-500-400-300-200-1000100200300400500

[kN

]

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[kN

]

(a) (b)

Figure 18. Shear force versus shear deformation of the URM infill wall (thickness= 102mm (4′′)):(a) level TAR 6; and (b) level DUZ 7.

as the test structure is subjected to different levels of shaking. Assuming uniform distributions ofthe inertia forces along the slab width, i.e. normal to the shaking direction, the maximum shearin the slab, VS,max, can be estimated as stated in Equation (6), where Vb denotes the total inertiaforce at each instant of time and VA and VC are the total shear forces resisted by the two bareframes on axes A and C, Figure 1, respectively.

VS,max = max{VA, (Vb/2) − VA, VC, (Vb/2) − VC} (6)

In order to investigate the change in the demand force in the slab, we consider two extremecases. (1) Assuming that the infilled frame (frame B) is infinitely stiff compared to the two bareframes (frames A and C), then frame B attracts all the inertia forces in the slab. It is implicit inthis assumption that the slab is rigid in its plane relative to the frames. In this case, VB = Vb andVA = VC = 0, and the maximum shear in the slab is Vb/2. (2) Assuming that the URM infill wallis completely disintegrated and the three frames, A, B and C, have the same lateral stiffness. Inthis case it is reasonable to assume that Vb is evenly distributed between the three frames and theshear force in each frame would be Vb/3. In this case, the maximum shear in the slab is Vb/3.

Examining the above two extreme cases suggests that the URM infill wall increases the sheardemands on the diaphragm. Figure 19(a) shows the ratio of the maximum shear in the slab tothe maximum total base shear in the test structure for different levels of shaking. As expected,this ratio reduces as the infill wall is damaged during high levels of shaking. The time historyresults from the experiment confirm this as well. Figure 19(b) shows the time history plot of themaximum shear in the slab (absolute value) as well as the two limits of Vb/2 and Vb/3 for levelDUZ 7. As discussed previously, severe damage in the URM infill wall occurred during this levelat t = 15.4 s. Comparing the slab shear forces for intervals of time at the beginning (t = 7.0 s) andat the end (t = 17.0 s) of the motion shows that the shear demand in the slab changes from close toVb/2 before the URM infill wall damage to being close to Vb/3 after the URM infill wall damage.

In order to measure the slab deformations, a set of displacement transducers are installed asshown in Figure 20(a). The displacement transducers are mounted on aluminium plates securedto the top of the prestressing bars extending from the centre of each column. The transducersare arranged to form an imaginary statically determinate truss. In order to find the displacement



0.3

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DUZ 8 Case: VA=VB=VC → VS=Vb/3

Case: VA=VC=0 → VS=Vb/2

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[kN

]

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(a) (b)

VS

VA

VB

VC

VS

BVb

B

Figure 19. Variation of shear demand in the slab due to damage in the URM infill wall: (a) slab-sheardemand variation at peak base shear; and (b) partial time histories for DUZ 7.

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x

y

1P

q

Vqu Tˆ

V

(a) (b) (c)

V S =

(167

000)

Figure 20. Shear deformation of the RC slab: (a) displacement transducers; (b) nodes and virtual forcesystem in equilibrium; and (c) slab shear stiffness.

of a specific node, we invoke the principle of virtual force (PVF) by applying a virtual forceP in the direction of the required displacement u and solve the determinate truss to find thecorresponding virtual internal forces satisfying equilibrium, Figure 20(b). Arranging the measureddeformations and the corresponding virtual internal forces into vectors V and q, respectively, u isdetermined from PTu = qTV with P = 1. Once the relative node displacements u are determined,the shear deformation, �, of the slab can be calculated as the differential displacement along theopposite edges of the slab divided by the distance between the two measurements. It should benoted that using PVF yields more accurate results than the approximate method used for the sheardeformation of the URM wall, Figure 17. The approximation is suitable for the URM infill wallbut not for the RC slab where the shear deformation is an order of magnitude smaller than that ofthe URM wall.

Figure 20(c) shows a plot of the maximum shear in the slab versus its corresponding sheardeformation. Although it is observed during the experiment that with increase in the intensity of



B1 B2

C2C1

Denotes trans. reinf. yielding

Denotes long. reinf. yielding

A2A1

Figure 21. Damage in the RC frames after level DUZ 7.

shaking, the gravity-induced cracks become more visible, Figure 20(c) suggests that there is nosignificant drop in the shear stiffness of the slab. The slope of the fitted line through the data points isabout one-eighth of the elastic shear rigidity of the uncracked section of the slab (GA= 5783MN(1.3× 106 kips)) indicating that although the shear rigidity of the slab remains more or lessunchanged, its value is significantly reduced due to initial gravity cracks. The deformed shapes(amplification factor= 250) plotted on the graph are obtained using PVF at times correspondingto the maximum slab shear deformations at levels TAR 2, DUZ 7 and DUZ 8. These deformationsinclude the effects of both shear forces and bending moments in the plane of the RC slab.

RC frames. As mentioned earlier, up to level TAR 6 and the initial part of DUZ 7, the stiff URMinfill wall carries most of the base shear protecting the three RC frames. At incipient failure ofthe wall and after the major cracking occurs, there are large shear demands at the bases of thecolumns where there are contacts with the wall as well as large bending moment demands onthe frames as the test structure undergoes larger displacements. Figure 21 shows the locationsin the test structure where reinforcing bar yielding is recorded by the strain gauges during levelDUZ 7 after partial failure of the wall. It can be observed that all the yielding occurred in themiddle infilled frame and the bare frames are practically undamaged. Figure 21 also includes aphotograph of the column–footing joint taken after level DUZ 7 and indicating the onset of damageat this joint location.

As discussed before, during level DUZ 8 and the repeat of level DUZ 7, i.e. DUZ 7-2, theeffect of the URM infill wall on the behaviour of the test structure diminishes and the RC framestake the significant portion of the earthquake-induced forces. The high demands on the RC framesduring these levels cause some damage in the form of local spalling and cracking in the beam–column and column–footing joints. Figure 22 shows photographs of the RC frame joints taken afterDUZ 7-2.

MODEL VALIDATION

The results of the experiment are compared with the simulated results of the OpenSees modelof the test structure to validate the prototype modelling and the established relationship between



(a) (b)

A1

B2

C2C1

A2

Denotes trans. reinf. yieldingDenotes long. reinf. yielding

(a)

(c)

(b)

(d)

(c) (d)

B1

Figure 22. Damage in the RC frames at the completion of level DUZ 7-2: (a) bare frame,column base; (b) bare frame, beam–column joint; (c) infilled frame, column base; and

(d) infilled frame, beam–column joint.

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ar [

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erim

ent

base

she

ar [

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]

(b)

DUZ 7 TAR 6 TAR 4 TAR 3 TAR 2

DUZ 7TAR 6TAR 4TAR 3TAR 2

Figure 23. Comparison between the experimental and simulated results: (a) concatenated input ground(shake-table) motion for the simulation; and (b) time histories of base shear.



96 98 100 102 104

96 98 100 102 104

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f di

sp.[

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]

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e sh

ear

[kip

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[kN

]

[mm]

(a) (b)

Figure 24. Comparison between the experimental and simulated results: (a) partial timehistories for DUZ 7; and (b) peak response.

the prototype substructure and the test structure as described at the beginning of the paper. Themeasured accelerations at the base of the test structure during levels TAR 2 thru DUZ 7 areconcatenated as a single input ground motion for the simulation, Figure 23(a). For this groundshaking, the simulated results are compared with the experimental results. Time history plots oftotal base shears are shown in Figure 23(b) with detailed partial plots for total base shears andfloor displacements (level DUZ 7) in Figure 24(a).

For levels TAR 1 through DUZ 7, the relationships of peak floor displacement versus the corre-sponding total base shear for the experimental and simulated results are compared inFigure 24(b). Considering the simple model used with a single strut for the URM infill wall,the OpenSees simulation results are in good agreement with the experimental results. How-ever, the computational model overestimates the peak floor displacement by 21% and under-estimates the corresponding total base shear by 16% at level DUZ 7 leaving the door open forfuture refinement of the computational model to obtain closer agreement with the experimentalresults.

CONCLUDING REMARKS

The URM infill wall has a significant role in the strength and ductility of the test structure andshould be considered in both analysis and design. Globally, it makes the test structure stiffer bya factor of 3.8, shortens the natural period of the test structure by 50%, increases the dampingcoefficient depending on the level of shaking from about 4 to 5–12% and increases the dissipatedenergy in the system. Such changes significantly affect the level of demand forces on the structureand generally reduce the displacement demands. Locally, the URM infill wall changes the loadpath and the distribution of forces between different elements of the test structure by increasingthe demand forces on its adjacent elements, e.g. the top and bottom of the RC columns and theRC slab. Quantitatively, the URM infill wall causes about 30% increase in the demand forces onthe diaphragm and collector elements in the test structure.



The benefits and limitations of the analytical model of the test structure is discussed andverified by comparing the simulation results with the shake-table test results. Using this model, therelationship between the prototype structure and the test structure is established. The test resultsalso serve as a basis for future modelling efforts to capture the behaviour of the infilled frame moreaccurately especially after initial cracking and through the transition phase. Efforts are underwayto introduce models that represent the cyclic behaviour of the masonry infills both for the in-planeand the out-of-plane directions. Such models can be used to analyse the behaviour of the prototypebuilding more reliably.

ACKNOWLEDGEMENTS

The authors gratefully acknowledge the financial support provided by the National Science Foundationunder NSF Contract No. CMS0116005. The reinforcing bars for the project were generously donated byMr T. Tietz, Western Regional Manager of the Concrete Reinforcing Steel Institute (CRSI). The help ofMr T. Elkhoraibi during the experiment is greatly appreciated. Special thanks are due to EERC laboratorypersonnel especially Mr W. Neighbour for his assistance.

REFERENCES

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15. ASTM A 615/A 615M. Standard Specification for Deformed and Plain Carbon-Steel Bars for ConcreteReinforcement. ASTM: West Conshohocken, PA, 2001.



16. ASTM C 873-99. Standard Test Method for Compressive Strength of Concrete Cylinders Cast-in-Place inCylindrical Molds. ASTM: West Conshohocken, PA, 1999.

17. ASTM C 270. Standard Specification for Mortar for Unit Masonry. ASTM: West Conshohocken, PA, 2003.18. ASTM C 1314. Standard Test Method for Compressive Strength of Masonry Prisms. ASTM: West Conshohocken,

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