Journal of Earthquake Engineering
ISSN: 1363-2469 (Print) 1559-808X (Online) Journal homepage: https://www.tandfonline.com/loi/ueqe20
Dynamic Tests of a Large-Scale Three-Story RC
Structure with Masonry Infill Walls
Ivica Guljaš, Davorin Penava, Lucas Laughery & Santiago Pujol
To cite this article: Ivica Guljaš, Davorin Penava, Lucas Laughery & Santiago Pujol (2018):
Dynamic Tests of a Large-Scale Three-Story RC Structure with Masonry Infill Walls, Journal of
Earthquake Engineering, DOI: 10.1080/13632469.2018.1475313
To link to this article: https://doi.org/10.1080/13632469.2018.1475313
Published online: 29 May 2018.
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JOURNAL OF EARTHQUAKE ENGINEERING
https://doi.org/10.1080/13632469.2018.1475313
Dynamic Tests of a Large-Scale Three-Story RC Structure with
Masonry Infill Walls
Ivica Guljaš
a
, Davorin Penava
a
, Lucas Laughery
b
, and Santiago Pujol
c
a
Faculty of Civil Engineering Osijek, Josip Juraj Strossmayer University of Osijek, Osijek, Croatia; bDepartment
of Architecture & Design, Nagoya Institute of Technology, Nagoya, Aichi, Japan; cDepartment of Civil
Engineering, Purdue University, West Lafayette, Indiana, USA
ABSTRACT
ARTICLE HISTORY
RC buildings with masonry infill walls are common throughout the
world. There is uncertainty about the effect of the infill walls on the
response of the building to earthquake demands. To shed light on
this issue, a three-story, scaled reinforced concrete frame with
masonry infill walls was subjected to two sequences of base motions
in two series of tests. In Series 1, hollow masonry units were used.
After completion of this series, the structure was repaired by replacing hollow masonry units with new solid units. Reinforced concrete
confining elements were also added along the vertical edges of
window and door openings in the first and second stories. The
repaired structure was then tested with the same sequence of base
motions in Series 2. On average, drift demands were 30% smaller in
Series 2 and were nearly half what the drift demands were estimated
to be for the bare RC frame alone. These results support the hypothesis that masonry infill panels can be used to control drift in low-rise
structures, provided they do not cause column failures and provided
they are restrained from out-of-plane collapse. Projections made
based on the test results also support the idea that a Wall Index
(ratio of masonry wall density to ten times the number of stories)
exceeding 0.2% leads to acceptable earthquake response in low-rise
buildings.
Received 23 October 2017
Accepted 7 May 2018
KEYWORDS
Dynamic Tests; Large-Scale;
Three-Story RC Structure;
Masonry Infill Walls;
Earthquake Action; Wall
Index
1. Introduction
Since the 1857 Naples Earthquake, the vulnerability of unreinforced masonry structures to
earthquakes has been a subject of repeated interest [Mallet, 1862]. Nevertheless, with the
advent of reinforced concrete (RC) frames, it became more common to use masonry walls
not as the primary structure, but as infills to separate functional spaces in buildings.
Compared with bare RC frames, these infill walls stiffen the structure and reduce its firstmode period. From the perspective of traditional strength-based design, this reduction in
period may appear to be detrimental. After all, we design for averages and, as Newmark
suggested [Newmark, 1973], in an average acceleration response spectrum reducing period
leads to an increase in acceleration demand in what Newmark referred to as the ranges of
CONTACT Davorin Penava
dpenava@gfos.hr
Faculty of Civil Engineering Osijek, Josip Juraj Strossmayer University
of Osijek, 3 Vladimir Prelog Street, 31000 Osijek, Croatia.
Color versions of one or more of the figures in the article can be found online at www.tandfonline.com/ueqe.
This article was originally published with errors. This version has been corrected/amended. Please see Corrigendum https://
doi.org/10.1080/13632469.2018.1499203.
© 2018 Taylor & Francis Group, LLC
2
I. GULJAŠ ET AL.
nearly constant velocity and displacement. Nevertheless, for the short structures in which
masonry infills are likely to be used, acceleration demand is likely to fall within what
Newmark called the range of nearly constant acceleration. From that point of view, it
follows that a shortening in first-mode period should not be expected to have detrimental
effects in terms of acceleration demand for an “average” spectrum.
In any case, displacement is often a better index to judge damage potential. A reduction
in period has time and again been observed to reduce displacement and as a consequence
reduce also damage. Decades of experience, the works of [Shimazaki and Sozen, 1984;
Bonacci, 1989; Lepage, 1997; Browning, 1998; Kalman Šipoš et al., 2013; Zovkić et al.,
2013; Matošević et al., 2014], the success of wall buildings in Chile [Sozen, 1989], and even
the early ideas of Howe [Howe, 1936] all suggest that the performance of RC structures
during earthquakes is related to drift demands, and that drift demand is related to
fundamental period. Drift has time been identified to be a reliable index for damage
potential [Sozen, 1980; Priestley, 2000], and is a key demand parameter in performancebased earthquake engineering [Moehle and Deierlein, 2004]. And for periods not exceeding 2 s in an average displacement response spectrum, mean drift demand tends to
decrease with decreases in period. This tendency implies that adding infill walls can
help reduce drift demand and, as a result, damage. Nevertheless, masonry infill walls do
more than just stiffen a structure: they also increase base shear strength. For short-period
structures and motions of high intensity, the work of Ozturk [Ozturk, 2003] suggests this
increase in strength is beneficial in helping reduce drift further.
The effects of stiffening and strengthening are not always positive: stiff and strong infill
walls can lead to column failures should the columns not have sufficient transverse
reinforcement [Louzi, 2014]. In addition, infill walls work well only up to modest levels
of drift: 1.5% or smaller [Negro and Taylor, 1996; Negro et al., 1996; Negro and Colombo,
1997; Pinto et al., 2002; Pujol and Fick, 2010; Stavridis et al., 2012]. If this threshold is
exceeded, damage can concentrate in a single story, creating the critical “soft story”
condition. Based on field observations, [Sezen et al., 2003] suggested soft stories caused
by failure of masonry infill walls may have contributed to building collapses in Turkey
during the 1999 Kocaeli earthquake. They also suggested that the presence of hollow clay
tile infills in the upper but not the lower stories of structures may have contributed to
other collapses. Although used for functional purposes in these cases, these infill walls
created a stiffness discontinuity in the structure. For reasons like these, the World Housing
Encyclopedia warns about possible detrimental effects if walls are not accounted for in
design [Murty et al., 2006].
Disagreement about the effects of masonry infill walls has prompted scores of investigations over the last several decades. Since the 1990s in particular, there has been more
emphasis on the use of shake-table tests of multistory structures to capture effects not
present in a static environment [Kwan and Xia, 1996; Žarnić et al., 2001; Lee and Woo,
2002; Dolce et al., 2006; Stavridis, 2009]. The test specimens in these investigations ranged
in number of stories from two to four and in scale from one-quarter to two-thirds, with all
but one specimen being one-third scale or smaller.
Kwan and Xia focused on the shear strength and damage in masonry infills of fourstory, one-bay structures. They concluded that the beneficial effects of infills would be
limited if out-of-plane failure is not prevented, and went on to suggest additional testing.
Žarnic et al. subjected one- and two-story structures to sine dwells with the goal of
JOURNAL OF EARTHQUAKE ENGINEERING
3
improving computational modeling approaches, and observed that RC buildings with
masonry infills designed in compliance with Eurocode provisions had significant overstrength that allowed them to sustain stronger motions that expected. Lee and Woo’s tests
of RC frames with non-seismic detailing and infills were similarly positive; they concluded
that infills were effective at increasing strength and at reducing drift without a detrimental
impact on the drift capacity of the structure. The tests of Dolce et al. were not only focused
on the use of shape-memory alloy retrofits in RC frames, but they also tested three-story,
two-bay bare RC frames and RC frames with masonry infills. Comparing the bare frame
with the infilled frame, they observed that the addition of “nonstructural” infills nearly
quadrupled the stiffness of the structure and led to a considerable reduction in drift ratio
for tests of similar intensity as well as the ability to sustain motions with larger PGAs.
The largest of these tests, by Stavridis et al., were two-third-scale, three-story, two-bay
structures with unconfined openings in one bay and poor details in the RC frames. These
details were meant to represent construction practice in California in the 1920s. Stavridis
et al. not only found that “infilled RC frames can behave in a safe manner during strong
earthquakes provided that sufficient infill walls are present,” but also observed the critical
soft story mechanism in the specimens. These results not only offer insight into the
response of RC-infill structures in a dynamic environment, but they also prompt
questions:
(1) Would confining elements around the openings in the solid infill walls tested by
Stavridis et al. or [Sigmund and Penava, 2014] have improved the performance of
the structure? Could these elements have prevented the out-of-plane failures mentioned by Kwan and Xia?
(2) Provided they are given sufficient confinement, could solid masonry infill walls
offer an inexpensive and effective retrofit in regions without access to more
advanced technologies?
The tests presented in this article address both of these questions.
2. Research Significance
Past studies suggest that masonry infill walls may offer an inexpensive retrofit option for
improving the seismic performance of bare RC frames with poor detailing. Nevertheless,
studies and field observations also warn against the possible detrimental effects of using
infill masonry: for instance, failure of adjacent RC columns that are not detailed for
increased demand and soft-story effects. In Taiwan, continuous infill masonry walls
confined by RC frame elements along all wall edges are being used successfully as
lateral-force resisting systems for school and government buildings [Pujol et al., 2017].
The use of infill for this purpose opens the door for an inexpensive and simple retrofit that
can be used in scores of vulnerable RC buildings in areas with high seismicity and
abundant resources to build masonry walls. Technology today can produce better solutions, but masonry may offer a cheaper alternative where clay and manpower are more
accessible than, say, dampers and base isolators. The objectives of this investigation are:
(1) to shed light on the effect of infill walls on seismic response through the shake-table
4
I. GULJAŠ ET AL.
testing of a 2/5-scale specimen, and (2) to show how masonry infills can be used as an
effective retrofit in RC frames for reducing drift demands.
These are still challenges that demand testing under plausible conditions for confident
analysis and generalization of the problem, in order to contribute to the safety of urban
populations in seismic regions and to the reduction of economic and human losses in
future earthquakes.
3. The Experimental Program
The test structure was a three-story, 1:2.5 scale RC frame with masonry infills (Fig. 1). Plan
and elevation views of the structure are shown in Figs. 2 and 3. For the first series of tests,
the infills were built using hollow clay masonry units. After the first series of tests, the
structure was repaired and tested again. These repairs consisted in patching the reinforced
concrete frames, replacing the hollow masonry units in the first two stories with solid
masonry units, and adding confining vertical elements. The objective of this second test
series was to investigate whether the repaired, framed-masonry system would respond
better to the applied earthquake demands. In the next section, details of the test specimen,
setup, and program are described.
3.1. Scaling
The specimen was built at a geometric scale of 1:2.5 scale. The maximum aggregate size in
the test specimen was 8 mm, compared with a typical size of approximately 20 mm for
full-scale RC construction. The masonry units were cut from full-scale masonry units to
preserve the relative area of voids. Mortar thicknesses also were reduced and were sized
for compliance with Eurocode 6, which governs the design of masonry structures in
Europe (EN 1996–1). Eurocode 6 Section 8.1.5 requires joints to be between 6 and
15 mm, corresponding to 2.4–6 mm for a scale factor of 2.5. For these tests, a joint
Figure 1. Structure on the shaking table before testing: (left) Series 1, with hollow masonry and no
vertical confining elements, and (right) Series 2, with solid masonry and confining vertical elements.
JOURNAL OF EARTHQUAKE ENGINEERING
5
Figure 2. Transverse, longitudinal, and plan views of the Series #1 structure (all dimensions in cm).
thickness of 6 mm was chosen for compliance with provisions and to improve constructability, as thin joints would have been difficult to construct and could have led to issues
with uneven joint thickness.
Similitude practices commonly used in reduced-scale earthquake simulation tests were
followed. Cauchy-Froude similitude was satisfied, which involved scaling mass and the
duration of the ground motion such that gm/EL2 in the prototype and the model were
equal (g = gravitational acceleration, m = mass, E = elastic modulus, and L = length). This
meant: (1) adding mass to the system and (2) reducing the time step of the full-scale ground
motion record (Sections 3.3 and 3.5). The first is done to increase the period of the system
and to produce meaningful axial stresses and vertical elements. The second is done with the
intent of matching key frequencies in the ground motion and test specimen. The addition of
6
I. GULJAŠ ET AL.
Figure 3. Transverse, longitudinal, and plan views of the Series #2 structure (all dimensions in cm).
mass and compression of records has been a common practice in earthquake simulation
tests of RC structures since they became more common in the 1970s [Gulkan and Sozen,
1971; Otani and Sozen, 1972; Hidalgo and Clough, 1974; Aristizabal and Sozen 1976].
3.2. Test Specimen
The reinforced concrete frames abutting the infill walls were designed in compliance with
Eurocode 2 and 8 provisions [CEN, 2004a; CEN, 2004b] as medium-ductility momentresisting frames. This Code governs the design of RC structures in Croatia. The maximum
JOURNAL OF EARTHQUAKE ENGINEERING
7
considered earthquake had a PGA of 0.3 g. Other design parameters used were: ground
type B response spectrum (type 1), 3.9 behavior factor, 5% damping, and building
importance class II.
In the longitudinal direction (Fig. 2), the frames had two bays with center-to-center
span lengths of 128 and 248 cm. The three frames in the transverse direction had one bay
with a center-to-center span length of 220 cm. Gross dimensions of the structure were
460 cm × 280 cm × 390 cm (length × width × height), including a 30 cm foundation. The
columns and beams were 12 cm wide and 16 cm deep. Each story was 120 cm tall
including an 8-cm-thick reinforced concrete slab. All concrete was class C25/30 with a
nominal cylinder compressive strength of 25 MPa (nominal cube strength of 30 MPa) and
a maximum aggregate size of 8 mm.
Reinforcement plans for the beams, columns, slabs, and foundation are given in Figs. 4–7.
The reinforcement (in the form of deformed bars for the frame and of welded wire fabric
reinforcement for the slab) was of Grade B500B with a nominal yield stress of 500 MPa.
Reinforcement ratios for frame elements are summarized in Table 1.
After the RC frames had cured for two weeks, 12-cm-thick masonry infill walls were
installed in the two longitudinal frames and the interior transverse frame. In the transverse
frame, these walls had a 40 cm × 104 cm door opening in the first story and 53 cm ×
67 cm window openings in the second and third stories. Infill walls in the longer span of
the longitudinal frames also had 40 cm × 104 cm door openings in the first story and
52 cm × 67 cm window openings in the second and third stories. All openings were
centered in the bays. The masonry infills in the shorter span of the longitudinal frames did
not contain openings.
Figure 4. Beam reinforcement details (units in cm except for bar sizes).
8
I. GULJAŠ ET AL.
Figure 5. Column reinforcement details (units in cm except for bar sizes).
The masonry units used in both series are illustrated in Fig. 8. Infills in Series 1 were
made of hollow clay masonry units measuring 12 cm × 25 cm × 6.5 cm (width × length ×
height). The volume of voids compared to the total volume of the unit was 68% (i.e., the
units were approximately 32% solid). The masonry units were laid with class M5 general
purpose mortar with a nominal strength of 5 MPa. Proportions by volume of cement,
lime, and sand were 1:1:5. Vertical and horizontal mortar joints were 6 mm thick. Material
properties for steel, concrete, and masonry are summarized in Table 2.
After the first series of tests, the structure was repaired and tested again. These repairs
consisted in replacing the hollow masonry bricks in the first and second stories with solid
clay bricks (12 cm × 25 cm × 6 cm). In addition, reinforced concrete confining elements
JOURNAL OF EARTHQUAKE ENGINEERING
Figure 6. Slab reinforcement details (units in cm except for bar sizes).
Figure 7. Foundation reinforcement details (units in cm except for bar sizes).
9
10
I. GULJAŠ ET AL.
Table 1. Reinforcement ratios of the frame columns and beams.
Member
Column
Beam
Mark
C1,C4
C2, C5
C3, C6
B1-7
Longitudinal reinforcement ratio
ρ
¼ AAcs
As
¼ bh
Tension side
As1
ρ ¼ bd
Compression side
ρ0 ¼ Abds2
2:26
¼ 1216
¼
4:02
¼
¼ 1216
2:26
¼
¼ 1216
1:2%
2:1%
1:6%
1:51
¼ 1214:2
¼ 0:88%
1:51
¼ 1214:2
¼ 0:88%
Shear reinforcement ratio
Within critical region
¼ 0:25
412 ¼ 0:52%
Outside critical region
0:25
¼ 0:21%
¼ 1012
Asw
Within critical region
ρw ¼ sb
w
¼ 0:25
412 ¼ 0:52%
Outside critical region
¼ 0:25
812 ¼ 0:26%
Asw
ρw ¼ sb
w
Figure 8. Hollow and solid clay masonry units used in this investigation (all units in mm).
were installed around window and door openings (see Fig. 3). Opening sizes were kept the
same. Details of the confining elements are presented in Fig. 9. These elements were
reinforced with two, deformed 8-mm longitudinal reinforcing bars. These bars were lap
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11
Table 2. Mean values of material properties.
Property
Concrete cylinder strength
Secant modulus of elasticity of concrete
Yield/ultimate tensile strength
Modulus of elasticity
Strain at yield/maximum force/ultimate
Value
D 4 mm
D 6 mm
D 8 mm
Masonry
Masonry
Masonry
Masonry
units, mortar, and masonry wallets
unit gross/net compressive strength
mortar compressive strength
wallet:
Tensile (diagonal) strength
Shear (diagonal) strength**
Series 1*
10.0/31.2
10.6
0.05
0.08
37
38800
753/780
204000
5.6/11.2/28.0
564/589
188000
5.0/24.8/60.8
591/621
199000
5.0/28.5/77.9
Series 2*
20.0/20.0
10.6
0.23
0.38
Units
MPa
MPa
MPa
MPa
‰
MPa
MPa
%
MPa
MPa
%
MPa
MPa
MPa
MPa
*Note: Series 1—Hollow clay block masonry with 68% voids. Series 2—Solid brick masonry.
**Shear strength is based on gross cross sections.
spliced to 8-mm postinstalled dowels embedded in the footings and the beams. The
dowels were installed by drilling 100-mm deep, 7-mm-diameter holes, and then hammering the dowels into them. No mortar or epoxy was used in the holes. In every mortar bed,
4-mm-diameter hoops were provided to anchor the infill wall to the confining element.
The objective of this second test series was to investigate whether the framed-masonry
would have better response to the applied earthquake demands. In a frame with limited
ductility and lateral stiffness, rehabilitation by introducing masonry infill is an attractive
solution as masonry can limit displacement demands by increasing stiffness.
3.3. Construction and Setup
The model building was constructed at the IZIIS laboratory in Skopje, Macedonia
[Necevska-Cvetanovska et al., 2015]. Two weeks after the final cast of the RC frame, the
hollow brick masonry infills were installed. The specimen was then placed on the shake
table and additional story masses were installed in the form of steel ingots. These ingots
were necessary to increase the period of the specimen by doubling its mass. The masses of
the structure and ingots are listed in Table 3. Testing began approximately four weeks
after installing the infill walls.
3.4. Instrumentation
Instrumentation of the structure is shown in Fig. 10. The setup was instrumented using
three types of sensors: displacement, acceleration, and strain. To measure absolute displacements, four string potentiometers were installed: one at the foundation and one at
each floor level. To measure acceleration, 49 accelerometers were installed throughout the
structure. To measure wall deformations, a total of 20 LVDTs were placed along diagonals
in the bays on the first two stories (Fig. 10). Strain gages with gage lengths of 50 mm were
12
I. GULJAŠ ET AL.
Figure 9. Confining element anchorage and reinforcement details in the cross-section (bottom), in
elevation (middle), and in structure (top).
attached to column longitudinal reinforcement near the column bases to measure axial
strains in columns. All sensors were sampled at 1000 Hz.
3.5. Loading Program
Base excitation was applied along a single-axis parallel to longitudinal frames. The applied
motion was modeled after the NS component of the record obtained at Herceg-Novi
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13
Table 3. Summary of masses.
Structural element
Foundation
Mass of the frame structure
Mass of the masonry infill
Mass of the structure
Additional masses (12 steel ingots per floor, 400 kg each)
Total mass
Series #1
mass (kg)
4,700
7,400
2,700
10,100
14,400
29,200
Series #2
mass (kg)
4,700
7,400
4,100
11,500
14,400
30,600
Figure 10. Typical instrumentation plan for the structure.
station during the 1979 Montenegro Earthquake. The time step of the original record was
reduced by a factor √2.5 to follow common similitude practices.
The compressed motion was then scaled in amplitude to reach different desired peak
ground accelerations (PGAs). Linear response spectra are shown in Fig. 11 for the
compressed motion scaled to PGA = 1 g. Also shown on this figure is the design spectrum,
0
5
10
15
0.0
0.1
0.2 0.3 0.4
Period, sec
0.5
0.6
Spectral velocity, cm/sec
0.2 0.3 0.4
Period, sec
0.5
0.6
0
0.1
0
0.0
1
2
3
4
5
6
50
100
150
200
250
300
0.0
0.1
Figure 11. Linear response spectra for compressed ground motion at PGA = 1 g (2% and 5% critical damping).
Spectral displacement,cm
20
Spectral acceleration, g
0.2 0.3 0.4
Period, sec
0.5
0.6
2%
5%
Design
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I. GULJAŠ ET AL.
JOURNAL OF EARTHQUAKE ENGINEERING
15
amplified to a PGA of 1 g to facilitate comparison. This figure also shows that the ground
motion had a broad range of frequencies, and that it represents the design spectrum well.
The specimen was subjected to ground motions of increasing intensity, with target
PGAs of approximately 0.05, 0.1, 0.2, 0.3, 0.4, 0.6, 0.7, 0.8, 1.0, and 1.2 g. These runs are
referred to as “5%” through “120%” hereafter in reference to the target PGA. In series two,
an extra motion was applied at 1.4 g. Before and after each test run, cracks and damage
were marked and recorded, and a free vibration test was conducted.
4. Results
4.1. Data
Data from the experimental investigation are available online at www.framed-masonry.
com. Data are in the form of TXT and MAT files. The data structure is described in the
reports on the website. Detailed drawings of the specimen and material properties also are
available on the website in PDF format.
4.2. Vibration Tests
Vibration tests were conducted on the bare RC frame as well as the infilled model before
and after each simulated earthquake test. The period of the bare RC frame was approximately 0.25 s. The period measured after installation of infills (with hollow infill walls) was
approximately 0.11 s (f = 8.8 Hz) before test Series 1. A period of 0.15 s was detected
during the initial, low-amplitude portion of the first ground motion. After five earthquake
tests reaching a PGA of 0.4 g, the period had elongated to approximately 0.33 s (3 Hz) as a
result of damage. Vibration tests conducted at the start of Series 2 showed a period of
approximately 0.13 s, but the initial low-amplitude vibrations during the 5% motion put
the period closer to 0.1 s. It is plausible that this difference is related to differences in
amplitude. After the final ground motion (PGA = 1.4 g), the period of the structure had
elongated to approximately 0.32 s.
Figure 12. Observed damage to frame A on the first floor after the final tests in both series: (left) Series
1, (right) Series 2.
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I. GULJAŠ ET AL.
Figure 13. Observed in-plane damage after the 0.4 g, 0.8 g, and final test in frames A (left) and B (right)
in Series #1.
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17
Figure 14. Observed in-plane damage after the 0.4 g, 0.8 g, and final test in frames A (left) and B (right) in
Series #2.
4.3. Observed Damage
Photographs of damage after the final tests in both series are shown in Fig. 12. Crack maps
are provided in Figs. 13 and 14 to show the damage states of the specimen after the tests at
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I. GULJAŠ ET AL.
0.4 g, 0.8 g, and after the final test in each series. Before the test at 0.4 g, in both test series
the structure experienced negligible-to-slight damage in the first story walls (based on the
definitions in [ATC, 1998; Grünthal et al., 1998]). At 0.4 g in both test series, cracks began
to form around the perimeter of the infill walls with some crushing at the corners.
Figure 14 also shows that there was slight diagonal cracking in the mortar joints of the
first and second story walls in Series 2 at this time. These cracks suggest the solid walls
were being engaged in resisting the lateral demands. The same was not observed in
Series 1.
At 0.8 g in Series 1, the first and second story infill walls began to develop inclined
cracks at isolated locations. These cracks formed in both directions in the shorter,
confined bay, suggesting that it was helping resist lateral demands in both directions. In
Series 2 at the same intensity, cracking in the first and second story infills was much more
extensive, although the third story still had negligible damage.
At the end of Series 1, the infill wall adjacent to the door opening had collapsed.
Otherwise, the damage patterns in the infills were similar to what had been observed after
the 0.8 g test. In contrast, infill damage at the end of Series 2 was much more extensive
than was observed after the 0.8 g test, although the third story walls still had negligible
damage. Once again, this damage suggested that the solid infills were being engaged more
effectively in resisting the lateral demands.
In both test series, damage to the main RC frame was negligible for most of the test
program, with hairline cracks only forming near the end of the testing program in the
beams of the second story and in the columns surrounding the shorter bay. Repair mortar
was applied in these cracks after Series 1. None of this damage compromised the vertical
load-carrying capacity of the RC frame. In Series 2, damage also was observed around the
vertical confining elements starting at 0.8 g and it became more severe in the first story by
the end of the test program.
4.4. Observed Response
Roof drifts were calculated as the difference between the measurements of the string
potentiometers at the roof and the foundation. Similarly, interstory drifts were calculated
as the difference between string potentiometers on two consecutive floors.
Accelerations at the foundation were filtered to exclude signals with frequencies outside
the range 0.5–400 Hz before integration to obtain base velocity. Peak ground velocity
(PGV) was calculated from these velocities with one exception. Vibrations caused by local
failures and/or collapse of debris are presumed to have affected the acceleration readings
obtained in the run at 120% in Series 1. For this run, PGV was therefore estimated by
differentiating a base displacement record (filtered to exclude signals with frequencies
larger than 70 Hz). Although it is well-known that taking differences of small quantities
can lead to large relative errors, the velocities obtained by differentiating the base
displacement histories generally showed good agreement with the velocities obtained by
integrating the base accelerations, with errors not exceeding 13% on average.
Key test results are summarized in Table 4 and Figs. 15 and 16. The most striking result
is that the structure survived test Series 1 with repairable damage (as demonstrated by
Series 2) despite these observations:
JOURNAL OF EARTHQUAKE ENGINEERING
19
Table 4. Summary of selected peak response values.
Measured
PGV*,
Max. measured
Run
cm/s
roof drift ratio
TEST SERIES 1
5%
2.4
0.03%
10%
5.3
0.05%
20%
8.2
0.08%
30%
9.9
0.17%
40%
14
0.17%
60%
19
0.30%
70%
36
0.47%
80%
36
0.54%
100%
57
0.93%
120%
63**
1.29%
TEST SERIES 2 (after repair)
5%
2.6
0.03%
10%
5.4
0.06%
20%
7.5
0.08%
30%
11
0.12%
40%
13
0.11%
60%
20
0.13%
70%
26
0.18%
80%
36
0.27%
100%
56
0.64%
120%
87
0.87%
140%
84
0.99%
Max. measured
interstory drift ratio
1F
2F
3F
Estimated max.
roof drift ratio
Ratio of max. measured roof drift
ratio to max. story drift ratio
0.04%
0.07%
0.10%
0.21%
0.28%
0.43%
0.73%
0.93%
1.66%
2.55%
0.05%
0.07%
0.09%
0.15%
0.26%
0.26%
0.37%
0.47%
0.72%
0.84%
0.05%
0.06%
0.08%
0.23%
0.26%
0.28%
0.33%
0.41%
0.45%
0.78%
0.05%
0.11%
0.17%
0.20%
0.29%
0.40%
0.74%
0.73%
1.17%
1.28%
2.0
1.5
1.2
1.3
1.6
1.5
1.6
1.7
1.8
2.0
0.03%
0.05%
0.08%
0.12%
0.13%
0.14%
0.20%
0.30%
0.93%
1.25%
1.63%
0.04%
0.07%
0.11%
0.14%
0.12%
0.14%
0.17%
0.26%
0.50%
0.63%
0.76%
0.03%
0.06%
0.08%
0.13%
0.11%
0.12%
0.17%
0.26%
0.93%
1.22%
1.50%
0.04%
0.08%
0.11%
0.16%
0.20%
0.30%
0.40%
0.56%
0.86%
1.33%
1.29%
1.7
1.2
1.3
1.2
1.1
1.1
1.1
1.1
1.5
1.4
1.7
*Mean of PGV obtained from integrated acceleration records (filtered to exclude signals with frequencies outside the range
0.5–400Hz).
**No sensible estimate of PGV was obtained from acceleration records obtained in Run 120%, Series 1. PGV for this run was
therefore estimated from a base displacement record (filtered to exclude signals with frequencies larger than 70 Hz).
Target
Series 1
Series 2
Spectral displacement,cm
PGA=0.6 g
PGA=1 g
20
20
15
15
10
10
5
5
0
0
0.0
0.1
0.2
0.3
0.4
Period, sec
0.5
0.6
0.0
0.1
0.2
0.3
0.4
Period, sec
0.5
0.6
Figure 15. Linear displacement response spectra for target and measured ground motions.
●
roof drift reached 1.3%,
the infill used is quite brittle and often assumed harmful to structural response, and
● the demands as discussed in the next section were high.
●
20
I. GULJAŠ ET AL.
1.4%
Seri es 1 - 120%
Measured Peak Roof Drift Ratio
1.2%
1.0%
0.8%
Series 1
0.6%
Series 2
0.4%
0.2%
0.0%
0
20
40
60
Peak Ground Velocity, cm/sec
80
100
Figure 16. Measured peak drift ratio at the roof versus peak velocity measured at foundation.
Figure 16 shows that drift was (1) nearly proportional to PGV and (2) was smaller for
solid masonry infills (Series 2).
Histories obtained from the string potentiometer indicate that there was some slip
relative to the shake table platform. This slip ranged from 0.2 cm (20% run, both
series) to 3.3 cm (120% run, both series). Although this slip led to differences between
the target displacement response spectra and the spectra from measured accelerations
(for motions with PGA in excess of 0.6 g), the spectra for motions of the same
amplitude were comparable between the two test series (Fig. 15), suggesting that
demands were comparable.
5. Analysis
To facilitate analysis of the test results, the initial (before cracking) period of the test
structure as tested in Series 1 was estimated as:
2
3
T¼
5 6 2π
7
4 qffiffiffiffiffiffiffiffiffi5
3 3:5 Em I
mw
(1)
H4
Em is modulus of elasticity of the masonry (discussed in the next paragraph),
I is moment of inertia, which was taken as 30% of the sum of moments of inertia of
continuous gross cross-sections in the first story,
H is building height (3.6 m),
mw is total mass (24.4 × 103 kg) divided by H, and
JOURNAL OF EARTHQUAKE ENGINEERING
0
Mortar strength (MPa)
10
15
20
5
25
12000
Only full-bed,
cement-sand-lime
masonry prisms
80
4:1
Masonry prism net strength (psi)
60
8000
50
6000
40
1.5:1
30
20
Masonry panel net strength (MPa)
70
10000
4000
21
Cement/Sand
0.4
0.3
0.2
0.1
2000
10
0
0
1000
2000
3000
0
4000
Mortar strength (psi)
Figure 17. Masonry prism net strength versus mortar strength based on past tests.
5/3 is a factor used to account for shear deformations and observed deviations between
calculated and measured stiffness of RC frames with masonry infill.
Masonry prisms were not tested by the contractor who ran the material tests, so two
approaches were taken to estimate prism compressive strength and modulus of elasticity:
(1) by consulting building code methods for estimating these values and (2) by compiling
a database of past experimental tests of prisms with similar mortar. In the first approach,
American and European Codes for the design of masonry structures were consulted: TMS
402/602 and Eurocode 6 [CEN, 2005; The Masonry Society (TMS), 2016]. For the properties of the mortar and masonry units used in this investigation, the elastic modulus
estimated using these codes ranged from approximately Em = 6–9 GPa.
To understand better the range of material properties plausible, results from more than 200
clay masonry prism compression tests were compiled into a database[Brown and Borchelt,
1990; Thomas & Scolforo, 1995; Tomaževič, 1999; Aryana, 2006; Pujol & Fick, 2010; Penava,
2012; Gazić, 2014; Matošević, 2015; Sorić, 2016]. In this database, the ratio of brick length-towidth ranged from 1.5 to 3.3 (compared with 2.1 for the bricks in this investigation), and the
mortar mixes had similar ratios of cement-to-sand-to-lime. The ratio of prism height-to-width
ranged from 2 to 5.6. For the tests in this database, Fig. 17 shows masonry prism net strength
plotted against mortar strength. Net strength is defined as the gross compressive strength of
the prism divided by the percent of the cross-section that was solid. Figure 18 shows that
prism strength can range from approximately 1.5 to 4 times the strength of the mortar, with
an upper limit equal to the net strength equal to the masonry units. For the given compressive
strength of the mortar in this study, these factors lead to estimated prism compressive
22
I. GULJAŠ ET AL.
MEASURED PEAK ROOF DRIFT RATIO
2.0%
1.5%
SERIES 1- 120%
SERIES 1, EM=10GPA
1.0%
SERIES 1, EM=21GPA
SERIES 2, EM=10GPA
SERIES 2, EM=21GPA
0.5%
0.0%
0.0%
0.5%
1.0%
1.5%
ESTIMATED PEAK ROOF DRIFT RATIO
2.0%
Figure 18. Measured versus estimated peak roof drift ratio.
strengths ranging from f’m = 16–30 MPa. Following TMS 402/602 guidance, the corresponding moduli would be Em = 700f’m, or 10–21 GPa (or 40% larger using Eurocode). The lower
bound of 10 GPa is similar to estimates from the building codes. Both Em = 10 and 21 GPa
will be considered next when estimating period.
For the conditions of the structure before test Series 1, the initial periods corresponding
to Em = 21 GPa and Em = 10 GPa were 0.07 and 0.1 s. Measurements of initial period ranged
from 0.11 to 0.15 s (Sec. 4.2). It is not uncommon in tests of scaled RC specimens to measure
initial periods ranging from one to nearly two times periods calculated assuming linear
response and nominal material properties [Laughery, 2016]. This discrepancy is likely to be
related to the effect of cracking caused by shrinkage in concrete. In the case at hand,
shrinkage is also likely to have caused cracking in mortar joints. In Series 2, the infill had
solid bricks. Assuming the full cross-sectional area of walls was effective in resisting bending
and shear leads to an initial period between 0.04 and 0.06 s for the range of moduli described
earlier.
These estimates of initial period can be used to judge and project the drifts measured.
We concentrate on drift (or drift ratio) because we believe drift is a better index for both
demand and capacity than strength or ductility. The design method proposed by Sozen
[Sozen, 2003] to estimate drift implies roof drift is:
pffiffiffi
1
Δ ¼ PGV T 2 Γ
2
(2)
PGV = peak ground velocity,
T = initial period, and
Γ = modal participation factor calculated for first mode and for a normalized mode
shape with a unit value associated with the roof level (approximately 1.3 for a structure
with three equal lumped masses and a linear mode shape).
JOURNAL OF EARTHQUAKE ENGINEERING
23
pffiffiffi
The factor 2 in Eq. (2) was proposed by Shimazaki [Shimazaki and Sozen, 1984] to
produce reasonably conservative results. We are using this expression here to evaluate and
project the test results so conservatism is unwarranted. Nevertheless, given the uncertainties related to period calculation mentioned above, it seems unnecessary to ignore
Shimazaki’s factor to modify an expression that has worked well before [Laughery, 2016].
Using the estimates of PGV listed in Table 4, for the run at 100% amplitude
in Series 1
pffiffiffi
and 2, the obtained roof drift is approximately Δ ¼ 12 56 cm
T
2
Γ
or
Δ ¼ Sd s
pffiffiffi
cm
2 Γ with Sd ¼ 28 s T. Sd is linear spectral displacement in the region of nearly
constant spectral velocity for a damping ratio of 2%. Considering that the time step in the
pffiffiffiffiffiffi
records used for simulation was compressed by 2:5, the associated spectral displacement
44cm
would be Sd ¼ 44cm
s T in the absence of said “time compression.” Sd ¼ s T is nearly
twice displacement demand (for a linear SDOF) considered in regular design for common,
stiff soil and site conditions in the United States [Lepage, 1997]. In other words, the
demand in the described tests was not trivial at all.
Figure 18 compares measured and estimated peak roof drift ratios, using periods
calculated for both Em = 10 GPa and Em = 21 GPa. The comparison suggests that the
test results obtained did not deviate far from previous experience. This figure also
compares the responses of the test structure as tested in Series 1 and 2. The test structure
after repair performed better than before the repair in terms of mean (or roof) drift, as was
observed in Fig. 16. Roof drifts for similar demands were nearly 30% smaller in Series 2.
Table 4 also reveals another advantage in the structure as tested in Series 2: the ratio of
maximum inter-story drift to maximum roof or mean drift was smaller in Series 2. The
mean of this ratio was 1.6 in Series 1 and 1.3 in Series 2. The smaller this ratio is, the
smaller is the risk of formation of a “soft story,” which can occur if damage concentrates
in a single story. In addition, the end elements added to the structure before test Series 2
helped prevent out-of-plane collapse of infills not confined along all four edges by RC
frame elements.
To try to create a frame of reference to judge the obtained test results, the following
exercise was conducted. The initial period of the bare frame was estimated assuming:
(1) The modulus of elasticity of concrete is: Ec ¼ 4 104 MPa
(2) Cross-sectional moment of inertia is equal to moment of inertia of gross section
(3) Effective beam cross-sections include the slab segment directly above a 45-deg
projection of the beam web (producing an effective flange width of 20 cm).
(4) The effects of beam-column joint stiffness were negligible
(5) Masses were lumped at beam mid-height and were 8200 kg for the lower two levels
and 6500 kg for roof level
The estimated period of the bare frame was 0.2 s, that is, more than twice the longer
period estimated for the structure with infills. According to Eq. (2), drifts would have
nearly doubled in the bare frame. At a mean drift of 2 × 1% = 2%, inter-story drift can
reach 3%. The detailing required for a reinforced concrete column to survive cycles at that
drift is uncommon in structures built in the 1950s, 1960s, and 1970s.
24
I. GULJAŠ ET AL.
5%
Estimated Drift Ratio, DR
4%
3%
2%
1%
0%
0%
2%
4%
6%
Wall Density, ρ
8%
10%
Figure 19. Variation of expected drift ratio with wall density.
The test results suggest that the building with infills is better off than an ordinary,
nonductile bare frame as long as: (1) masonry does not fall out of plane and (2) masonry
does not cause failures in columns. Despite being uniaxial, the tests suggested that to
reduce out of plane instability all one has to do is to confine the masonry infills. The other
aspect of the second round of tests that is interesting is that rather simple repairs made the
structure as good as new (or better).
6. Projections
Equations 1 and 2 were found to produce reasonable estimates of mean drift ratio for the
reduced-scale structures tested in this investigation. Combined, they may also provide a
useful vehicle for estimating the amount of wall area that may be needed to keep drift
demands below a given threshold. This would offer a valuable starting point for owners
interested in using masonry infills for retrofits. Substituting Eq. (1) into Eq. (2) and
assuming the effective moment of inertia of a masonry wall is 40% the gross moment of
inertia leads to:
rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi
γ
DR ¼ 12 PGV Γ α
(3)
ρ Em h
where DR is roof drift ratio Em is masonry modulus of elasticity. Here, the lower bound
value of approximately 10 GPa is used because conservativism is desirable in projections,
h is story height, assumed to be 2.5 m for a typical building,
PGV is peak ground velocity (assumed not to exceed 2/3 m/s),
α is ratio of wall height to wall length (aspect ratio) if all walls have the same thickness
P
P
0:5
(tw) and aspect ratio. If not, α ¼
twi α3
,
twi α1
i
i
Γ is participation factor for a mode shape factor of 1 at roof level (taken as 1.3),
γ is mass per unit area (assumed to be 1100 kg/m2 = 200 psf/g), and
JOURNAL OF EARTHQUAKE ENGINEERING
25
ρ is wall density (ratio of total gross cross-sectional area of wall in one direction to
tributary floor area for one floor).
Figure 18 shows how drift ratio varies with wall density for a wall aspect ratio of 1.5
(e.g., 5-m-long walls in a three-story building with a total height of 7.5 m). The figure
suggests that if the limiting drift ratio of the RC-masonry composite is 1.5%, as observed
by Pujol and Fick [Pujol and Fick, 2010], then hollow RC-masonry walls need to occupy
approximately 4.5% or more of the floor area in each direction near the building base
(unless the wall aspect ratio is larger than 1.5). A smaller area would be needed if the wall
aspect ratio is smaller than 1.5.
In Fig. 19, it is interesting that the rate at which drift decreases with increases in wall density
seems to “flatten out” near a wall density of 4.5%. For a wall thickness of 25cm (20-cm-wide units
with 2.5-cm-thick plaster on each face), a wall density of 4.5% would require a wall every 5–6 m (a
reasonable distance between partition walls in residential construction). If needed, double-wythe
masonry can be used to allow for longer distances between walls. Solid masonry can be used in
smaller densities, but strong infill may be more likely to induce column failure and may require
additional column transverse reinforcement near column ends [Mehrabi, 1994].
For buildings with more than three stories, it may be more efficient to use RC walls. If
masonry infill is preferable, to maintain an aspect ratio of 1.5 so that wall density can be kept at
4.5%, wall length would need to be increased beyond 5 m by nearly 1.5 m for every story in
excess of 3. Buildings with more than five stories should be examined with more care and may
require more reliable means to control drift than masonry infills. The use of reinforcement
embedded in the masonry and anchored in RC frame elements is bound to help the RCmasonry composite and reduce the hazard created by falling pieces of masonry.
Figure 20. Taiwanese government building with masonry infills (Pujol et al., 2017).
26
I. GULJAŠ ET AL.
7. Correlation with Field Data
Results from this investigation and expected drift ratios from Eq. (3) also are consistent
with observations of building damage after major earthquakes. Figure 20 shows a
Taiwanese government building with an equivalent masonry wall density of nearly 3.5%
(and no other source of lateral stiffness) in its short direction. With a wall aspect ratio of
1–1.5 and PGV in the area not exceeding 60 cm/s, Eq. (3) would yield a drift estimate of
1% of building height. Observations of damage from the tests presented here suggest that
such a drift ratio doesn’t always lead to severe damage. Observations from the field were
consistent with this expectation: during the earthquake of February 6, 2016 near Tainan,
Taiwan, this building and a number of similar buildings sustained only minor structural
damage in their short directions.
Over one thousand RC buildings like the one shown in Fig. 20 have been surveyed after
major earthquakes since 1992 for the purpose of evaluating vulnerability indices [Hassan
and Sozen, 1997; Lepage, 1997; Villalobos et al., 2018; Pujol et al., 2000; Pujol et al., 2017;
Ozcebe et al., 2003; Donmez and Pujol, 2005; O’Brien et al., 2012; Zhou et al., 2013; Shah
et al., 2017]. One index similar to masonry density is the Wall Index: the ratio of wall
density to ten times the number of stories [Hassan and Sozen, 1997]. For the buildings
listed in the aforementioned references, a Wall Index of 0.2% has been observed to lead to
acceptable performance. For a three-story building, this WI corresponds to a masonry wall
density of 6% which is similar to the limit at which expected drift ratio started to “flatten
out” in Fig. 19. Overall, these findings lend credence to Eq. (3) as a reasonable way to
approximate the amount of wall area needed to avert damage.
8. Conclusions
The test results support the hypothesis that masonry infill panels can be used to control
drift in low-rise structures. Their effect is beneficial as long as they do not cause column
failures and as long as they are constrained against out of plane collapse as was done in
test Series 2. This observation implies that existing vulnerable RC structures can be
strengthened simply by adding masonry infill, and/or rearranging their partitions (if
they are already masonry), so that:
(1) Captive columns created by discontinuous walls are eliminated
(2) Discontinuities in masonry infill, both in plan and elevation, are also eliminated
(3) The resulting structure has a period (from Eq. 1) short enough to limit drift (from
Eq. 2) to less than 1%–1.5% of building height.
The interpretation of the results presented also agrees with reported observations that a
Wall Index (ratio of masonry wall density to ten times the number of stories) exceeding
0.2% leads to acceptable earthquake response in low-rise buildings.
Acknowledgments
The authors are grateful to Professor Vladimir Sigmund (1956–2016) for his significant contribution and direction toward the research described in this article.We also would like to acknowledge
JOURNAL OF EARTHQUAKE ENGINEERING
27
Jonathan Monical, a Graduate Research Assistant at Purdue University, for his work preparing
revisions for this manuscript.
Funding
The research presented in this article is a part of the research project FRAmed-MAsonry composites for
modeling and standardization [HRZZ-IP-2013-11-3013] supported by Croatian Science Foundation
and its support is gratefully acknowledged.
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