Journal of Structural Engineering

Journal of Structural Engineering
Journal of Structural Engineering
Seismic Simulation of Integrated Ceiling-Partition Wall-Piping System at E-Defense,
Part 1: Influence of 3D Structural Response and Base Isolation
--Manuscript Draft-Manuscript Number:
STENG-3259R1
Full Title:
Seismic Simulation of Integrated Ceiling-Partition Wall-Piping System at E-Defense,
Part 1: Influence of 3D Structural Response and Base Isolation
Manuscript Region of Origin:
UNITED STATES
Article Type:
Technical Paper
Section/Category:
Seismic Effects
Abstract:
The seismic response of a full-scale, five-story steel moment frame building in baseisolated and fixed-base configurations with an integrated suspended ceiling-partition
wall-sprinkler piping system that was shaken at E-Defense is critically assessed.
Horizontal floor accelerations were constrained by the isolation systems to relatively
low levels, which allowed observation of damage to the integrated system that was
directly related to the vertical component of input acceleration. The floor slabs exhibited
single mode vibration at their natural periods with widely varying effective damping.
Peak vertical accelerations were amplified by an average factor ranging from 3 to 6
from the table to the middle of the floor slabs, where amplification factors increased as
slab vibration periods lengthened. Damage to the ceiling-partion-piping components
initiated at slab accelerations of about 2g, and became extensive for slab accelerations
exceeding 5g. These metrics establish target vertical accelerations for achieving
desired performance objectives.
Corresponding Author:
Keri L Ryan, PhD
University of Nevada, Reno
Reno, NV UNITED STATES
Corresponding Author E-Mail:
[email protected]
Order of Authors:
Keri L Ryan, PhD
Siavash Soroushian
Emmanuel Manos Maragakis
Eiji Sato
Tomohiro Sasaki
Taichiro Okazaki
Suggested Reviewers:
Tara Hutchinson
University of California, San Diego
[email protected]
Hutchinson led another recent test program of a full-scale building outfitted with
nonstructural components and has desirable expertise for all aspects of the companion
papers. However, Hutchinson is also a Co-PI on the Grand Challenge project, and so it
might be considered a conflict of interest, although she was not involved in this aspect
of the project.
Amir Gilani
Miyamoto International
[email protected]
Gilani has published experimental research on seismic response of ceiling systems
and has expertise to evaluate structural and nonstructural responses observed in this
research.
Claudia Marin
Howard University
[email protected]
Marin has expertise in seismic isolation and was involved in the test program on fullscale building outfitted with nonstructural components that was led by Hutchinson.
Matthew and
Hoehler
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Cover Letter
Click here to download Cover Letter: Cover Letter_Revision 1.pdf
College of Engineering
University of Nevada Reno
Oct. 26, 2014
Sherif El-Tawil, Ph.D., P.E., F.ASCE,
Dept. of Civil and Env. Engineering
University of Michigan
Ann Arbor, MI 48109-2125
Ph (734) 764-5617 Fax (734) 764-4292
[email protected]
Managing Editor of ASCE Journal of Structural Engineering
Dear Professor El-Tawil:
On behalf of the NEES/E-Defense collaborative research project on base-isolation and
nonstructural components, I hereby submit revised versions of STENG-3259 and STENG-3260
for further consideration as Technical Papers in Journal of Structural Engineering. The titles of the
manuscripts have been revised to “Seismic Simulation of Integrated Ceiling-Partition Wall-Piping
System at E-Defense, Part 1: Influence of 3D Structural Response and Base Isolation”, and
“Seismic Simulation of Integrated Ceiling-Partition Wall-Piping System at E-Defense, Part 2:
Evaluation of Nonstructural Damage and Fragilities”. We wish the revised manuscripts to be
reviewed as companion papers.
The reviewer requested additional information in Part 1 so that the study could be better understood
without referring to other documents. Therefore, the length of the submitted manuscript for Part 1
has increased from 9 to 10 pages according to the sizing worksheet estimate. We believe the
additions have led to a more readable paper, and the additional length is justified. The length of
Part 2 remains at an estimated 10 pages. To reiterate from the original submission, the manuscripts
contain color figures, but all figures can be understood in black and white. Thus, our intention is
for figures to be published in color electronically and in black and white for the printed journal.
We look forward to learning the outcome of the manuscript peer review process. If there are any
problems with the submission, please let me know.
Sincerely,
Keri L. Ryan, Ph.D.
Associate Professor
Department of Civil and
Environmental Engineering
University of Nevada, Reno/258
Reno, Nevada 89557-0152
(775) 784-6937 office
(775) 784-1390 fax
Manuscript
Click here to download Manuscript: E-Defense companion paper 1 - Revision 1 - No Figures.docx
1
Seismic Simulation of an Integrated Ceiling-Partition Wall-Piping
2
System at E-Defense, Part 1: 3D Structural Response and Base
3
Isolation
4
Keri L. Ryan,a) Siavash Soroushian,b) E. "Manos" Maragakis,c) Eiji Sato, d)
5
Tomohiro Sasaki, e) Taichiro Okazaki f)
6
Abstract
7
The seismic response of a full-scale, five-story steel moment frame building in base-isolated
8
and fixed-base configurations with an integrated suspended ceiling-partition wall-sprinkler
9
piping system that was shaken at E-Defense is critically assessed. Horizontal floor accelerations
10
were constrained by the isolation systems to relatively low levels, which allowed observation of
11
damage to the integrated system that was directly related to the vertical component of input
12
acceleration. The floor slabs exhibited single mode vibration at their natural periods with widely
13
varying effective damping. Peak vertical accelerations were amplified by an average factor
14
ranging from 3 to 6 from the table to the middle of the floor slabs, where amplification factors
15
increased as slab vibration periods lengthened. Damage to the ceiling-partion-piping components
16
initiated at slab accelerations of about 2g, and became extensive for slab accelerations exceeding
a)
Assoc. Prof, Dept. of Civil and Environmental Engineering, University of Nevada, Reno, MS 0258, Reno, NV
89557-0258
b)
Post-doctoral Scholar, Dept. of Civil and Environmental Engineering, University of Nevada, Reno, MS 0258,
Reno, NV 89557-0258
c)
Dean of Engr., Dept. of Civil and Environmental Engineering, University of Nevada, Reno, MS 0258, Reno, NV
89557-0256
d)
Dr. Engr., National Research Institute for Earth Science and Disaster Prevention, 1501-21 Nishikameya, Mitsuta,
Shijimi-cho Miki, Hyogo, Japan 673-0515
e)
Dr. Engr., National Research Institute for Earth Science and Disaster Prevention, 1501-21 Nishikameya, Mitsuta,
Shijimi-cho Miki, Hyogo, Japan 673-0515
f)
Assoc. Prof., Graduate School of Engineering, Hokkaido University, Kita 13, Nishi 8, Kita-ku, Sapporo,
Hokkaido, Japan, 060-8628
17
5g. These metrics establish target vertical accelerations for achieving desired performance
18
objectives.
19
Keywords: seismic isolation, nonstructural components, seismic response, vertical ground
20
motion, floor slab vibration, shake table testing
21
Introduction
22
The performance of critical facilities such as hospitals and fire stations during an earthquake
23
depends not only on the structural systems, but also on the functionality of nonstructural
24
components. The shaking intensities that can cause damage to the nonstructural components are
25
typically lower than those that induce structural damage (Miranda, 2003), and the influence of
26
vertical shaking on nonstructural components has not been well studied. Several component and
27
subsystem level experiments of nonstructural components such as suspended ceilings, partition
28
walls, and sprinkler piping, have been conducted in recent years (e.g. Badillo-Almarez et al.,
29
2007; Gilani et al., 2010; Zaghi et al., 2012; Tian et al., 2012; Retamales et al., 2012). However,
30
these component and subsystem level experiments may not accurately reflect the following
31
influences of real buildings on nonstructural response: realistic input excitation, interaction
32
between different types of nonstructural components, realistic boundary conditions, and floor
33
system vibration.
34
Mitigation strategies like seismic base isolation are often chosen when the seismic
35
performance objective is post-earthquake continued functionality (Taylor and Igusa, 2004). Base
36
isolation has been shown to effectively protect structural systems from damage through the
37
substantial reduction of story drifts, but its ability to eliminate nonstructural damage has not been
38
proven conclusively. Nonstructural components in base-isolated buildings are subjected to a
39
different proportion of horizontal to vertical accelerations than prescribed by seismic
2
40
qualification tests, and so the observations from component tests referenced above may not
41
strictly apply. Specifically, ACC 156 prescribes that the horizontal acceleration is amplified over
42
the building height while horizontal acceleration in a base-isolated building is significantly
43
reduced relative to the ground acceleration (ICC 2010).
44
Observations of the response of numerous base-isolated buildings in past earthquakes, mostly
45
favorable, have been reported (e.g. EERI, 1996; Saito, 2006; Moroni et al., 2012; Gavin and
46
Nigbor, 2012; Kasai et al., 2013). Although problems with expansion joints and some disruption
47
of contents (e.g. sliding or toppling of furniture, items falling from shelves) have been observed,
48
significant nonstructural component damage has not been reported. For instrumented buildings,
49
authors confirmed that the isolation system generally attenuated the horizontal accelerations
50
recorded at the floor level relative to the ground, especially when the input ground acceleration
51
was significant enough to activate the isolation systems (Kasai et al., 2013). It does not appear
52
that any isolated building, constrained to low horizontal floor accelerations, has been subjected
53
to significant vertical accelerations; however, most reports do not mention vertical response.
54
Among the studies we examined, the peak vertical acceleration recorded in any isolated building
55
was 0.76g, which was observed in the base-isolated Nursing Home Building of the Niigata-ken
56
Cheutsu Earthquake of 2004 (Saito, 2006), for which no damage was reported. In the 2011
57
Christchurch Earthquake, free field vertical accelerations of 0.6g were recorded near the
58
Christchurch Women’s Hospital (Bradley and Cubrinovski, 2011), which was not instrumented.
59
Nonstructural damage was reportedly “limited to minor cracking of partitions around window
60
openings” (Mayes et al., 2012).
61
Although the accelerations were not large, significant amplification of vertical accelerations
62
from ground to roof level in a few buildings has been documented. Bozorgnia et al. (1998)
3
63
examined 12 buildings – including base-isolated and conventional with varying levels of vertical
64
instrumentation – that were shaken in the 1994 Northridge Earthquake. Three of these buildings
65
contained vertical accelerometers both at the ground level and on the roof slab away from a
66
column. The ratios of peak vertical accelerations recorded in the roof to those at the ground level
67
were 3.75, 6.4, and 2.4. In addition, Lew and Hudson (1999) concluded, by examination of
68
several instrumented buildings, that vertical accelerations transmitted from the ground to the
69
building were not affected by the presence of an isolation system.
70
Besides field observation, the attenuation of input horizontal accelerations at upper levels by
71
base isolation has also been demonstrated numerous times by shaking table tests (Kelly et al.,
72
1980; Al-Hussaini et al., 1994; Clark et al., 1997; Wolff and Constantinou, 2004; Fenz and
73
Constantinou, 2008; to name a few). The vast majority of these have been conducted on bare
74
frame structures with added mass, and thus could not explicitly address the performance of
75
nonstructural components. Only recently, state-of-the-art large capacity shake table facilities
76
have provided opportunities to test full-scale base-isolated buildings constructed with realistic
77
floor systems and nonstructural components, such as the tests of a full-scale 4-story reinforced
78
concrete medical building at Hyogo Earthquake Engineering Research Center/E-Defense (Sato et
79
al., 2011). In a specific investigation of the effect of vertical excitation in these tests, Furukawa
80
et al. (2013a) reported that the peak vertical acceleration recorded at the table (0.5g) was
81
amplified to about 4g at the center of the roof slabs during a 3-dimensional (3D) El Centro
82
ground motion. Disruption and damage to building contents and equipment (the main focus of
83
the study) was significant, but damage to the Japanese-designed suspended ceilings, piping and
84
sprinkler systems, and partition walls/sliding doors was not observed.
4
85
This paper focuses on another major test of a full-scale base-isolated building incorporating
86
nonstructural components. As part of a collaborative research project between Network for
87
Earthquake Engineering Simulation (NEES) and the National Research Institute for Earth
88
Science and Disaster Prevention (NIED) of Japan, system-level full-scale shaking experiments of
89
a 5-story steel moment frame building were conducted at E-Defense. The building was tested in
90
three different configurations: 1) base-isolated with triple pendulum (TP) bearings (TP
91
configuration), 2) base-isolated with a hybrid isolation system (hybrid configuration), and 3)
92
fixed-base configuration. An integrated ceiling, partition wall, and fire sprinkler piping (CPP)
93
system was installed in the building for each test configuration, as led by the NEES Grand
94
Challenge project “Simulation of the Seismic Performance of Nonstructural Systems”. This is
95
the first of two related papers. In this paper, we evaluate the 3D structural response in all
96
configurations, and identify the relation between CPP damage and observed horizontal floor
97
accelerations and vertical slab accelerations. In the follow-up paper (Soroushian et al., 2013a),
98
the response of the integrated ceiling, partition, and fire sprinkler piping systems is discussed,
99
damage states are identified, and fragility functions are developed.
100
Experimental Setup
101
Building Testbed
102
The CPP system was installed in a 5-story testbed building (Fig. 1) that was utilized for this
103
collaborative NEES/E-Defense project. This building was approximately 16 m (53 ft) tall,
104
asymmetric in plan with dimensions of 10 m by 12 m (33 ft by 40 ft), 2 bays by 2 bays, and
105
weighed about 5220 kN (1174 kip) (Fig. 2(a)). Lateral resistance was provided by steel moment
106
frames in both directions. The columns were made from 350 mm x 350 mm (13.8 in x 13.8 in)
107
hollow square sections with thicknesses that varied from story to story. The beams were either
5
108
rolled or built-up I-sections. The primary beams, which were connected to the columns via fully
109
restrained moment connections, consisted of 3 x 400 mm (15.7 in) deep segments that were
110
bolted together at the approximate inflection points determined from gravity loading. The end
111
segments were haunched to improve their bending strength and the beam-to-column connection
112
strength. Secondary beams spanned between the primary beams as configured in Fig. 2(a); the
113
beams – with depths ranging from 200 to 400 mm (7.9 in to 15.7 in) – were bolted to the primary
114
members (idealized as pin connections).
115
The floor system in floors 2-5 consisted of 80 to 155 mm (3.15 to 6.10 in) thick concrete
116
slabs cast on 1.2 mm (0.047 in) thick corrugated metal decking. Typical slab reinforcement was a
117
single layer of 10 mm (0.39 in) diameter bars placed at the mid surface of the slabs, and spaced
118
150 mm (5.90 in) in both directions. The roof slabs were composed of 150 mm (5.90 in) thick
119
concrete slabs cast on a 1.2 mm (0.047 in) flat steel deck. Two layers of 13 mm (0.51 in)
120
diameter bars spaced at 200 mm (7.9 in) provided reinforcement in each direction. On average,
121
the roof slab was thicker than the floor slabs, to carry additional weight as described below.
122
Shear studs, which were covered by concrete for protection, connected the concrete slabs to the
123
primary beams to provide composite beam-slab behavior.
124
In addition to the weight of structural components, concrete and steel weights were installed
125
to simulate a realistic live load. Concrete blocks, which contributed an additional 175 to 257 kN
126
(39.3 to 57.8 kip) per floor, were installed on floors 2 to 5 with representative dimensions and
127
position as shown in Fig. 2(b). Where the blocks intruded the effective width of the composite
128
slab, polystyrene foam was inserted between the block and the slab so that the blocks did not
129
contribute to composite bending of the floor system. Steel plates weighing 535 kN (123 kips)
130
were installed on the east side of the roof to simulate the weight from a combination of roof
6
131
mounted equipment (e.g. air conditioner system or water tanks) and a penthouse (Fig. 2(c)).
132
Each weight included either 7 or 8 steel plates with dimensions 2.1 x 4.3 x 0.025 m (6.9 x 14.1 x
133
0.08 ft). The weight at the roof was altered from the testbed’s original design configuration to
134
introduce a significant mass eccentricity.
135
Two independent isolation systems were designed, incorporated beneath the building, and
136
tested as part of this experiment. Each isolation system design was developed by the project team
137
in conjunction with the bearing supplier to meet distinct objectives. The TP isolation system
138
incorporated 9 identical TP bearings (Fig. 3(a)), one beneath each column depicted in the plan
139
layout in Fig. 2(a). The properties of these bearings were chosen to accommodate the largest near
140
fault ground motions that could be replicated by the E-Defense shake table. The bearings were
141
1.4 m (4.6 ft) in diameter, with a displacement capacity of 1.13 m (3.7 ft) at a base shear
142
coefficient V/W = 0.275, where V = isolator shear force and W = axial force due to building self-
143
weight. The inner pendulum mechanism (period T = 1.84 s) is engaged in minor shaking, with
144
transition to the outer pendulum mechanisms (period T = 5.57 s) when V/W first reaches about
145
0.08. The isolation system reached a peak displacement of 0.7 m during testing. The TP bearings
146
were very stiff in compression, while providing no resistance in tension. The bearing vertical
147
stiffness was assumed to correspond to a rigid mass vertical frequency of 30 Hz (T = 0.03 s).
148
Given the history of investigation of elastomeric bearings for nuclear applications, the
149
hybrid isolation system evolved from the need to verify the stability and load-carrying capacity
150
of elastomeric bearings at displacements representative of the nuclear industry’s extended design
151
basis ground motions. The hybrid isolation system featured 4 lead-rubber (LR) bearings (Fig.
152
3(b)) for lateral resistance, each 0.7 m (2.3 ft) in diameter with a 102 mm (4 in) lead core and a
153
shape factor S = 29. The manufacturer provided nominal bearing properties were: elastic
7
154
stiffness = 6.5 kN/mm (37 kip/in), post-yield stiffness = 0.65 kN/mm (3.7 kip/in), characteristic
155
strength = 65.7 kN (14.8 kip), and nominal vertical stiffness in compression = 1500 kN/mm
156
(8566 kip/in). The design of the system was also heavily influenced by the building’s light
157
weight; to provide the desired isolation period shift while simultaneously accommodating the
158
overturning moments, the LR bearings (located beneath edge columns) were supplemented with
159
5 cross-linear (CL) bearings (Fig. 3(c)), which were located beneath center and corner columns.
160
The CL bearings were composed of two sets of perpendicular rails with top and bottom guided
161
plates separated by nearly frictionless ball bearings. These low friction sliders share the vertical
162
load with the LR bearings without increasing the horizontal stiffness or base shear of the
163
isolation system. Furthermore, the CL bearings enhance the global system stability by allowing a
164
redistribution of axial force between LR and CL bearings, while also providing tension
165
resistance. The CL bearings were stiff  relative to the LR bearings  in the vertical direction,
166
with manufacturer-supplied nominal vertical stiffnesses of 3471 kN/mm (19821 kip/in) in
167
compression and 245 kN/mm (1399 kip/in) in tension.
168
The LR bearings reached a peak displacement of 0.55 m during testing. The nominal
169
isolation system properties provided horizontal post-yield period Td = 2.6 s and characterized
170
strength Qd/W (summed over 4 bearings) = 0.050. Based on the sum of the nominal vertical
171
stiffnesses, the rigid mass vertical frequency of the hybrid isolation system was 30 Hz (T = 0.03
172
s). Additional details about the testbed, isolation system design, and overall setup are available in
173
Dao (2012) and Ryan et al. (2013a) for the interested reader.
174
Nonstructural Systems
175
A partition-ceiling-sprinkler piping subassembly was designed and installed in nearly
176
identical configuration over two complete floors of the testbed building. These components were
8
177
installed below the 5th and roof floor slabs, which were expected to draw the maximum floor
178
accelerations. The following components were included on each floor: approximately 83.6 m2
179
(900 sf) of lay-in-tile suspended ceiling designed per ASTM E580/E580M-11ae1 (ASTM, 2011),
180
a standard Schedule 40 piping system conforming to NFPA 13 (2011), and approximately 90 m
181
(300 ft) of typical light gauge steel studded gypsum partition walls with individual lengths
182
varying from 1.5 to 9.8 m (5 to 32 ft). Detailing variations were considered, such as: suspended
183
ceiling sections with or without seismic bracing; straight drops, armover drops, or flexible hose
184
drops for the piping sprinkler heads; and slip track or full connection details for partition walls.
185
Further details of the CPP components can be found in Soroushian et al. (2013a).
186
Instrumentation
187
The table accelerations and the responses of structural and CPP components were monitored
188
by nearly 400 sensor channels (not including the isolation system response, when applicable)
189
recorded at a sampling frequency of 1000 Hz. A 4-pole low-pass Butterworth filter with a cutoff
190
frequency of 25 Hz was applied to all recorded responses, unless otherwise mentioned.
191
Figure 4 shows the layout of accelerometers on a typical floor that are applicable to the
192
responses presented here. Achieved table accelerations were measured by triaxial accelerometers
193
mounted at the 4 corners of the shake table (Fig. 4). Floor accelerations (2 horizontal and vertical
194
components) were measured using triaxial accelerometers installed at the southeast (SE),
195
northeast (NE) and northwest (NW) corners of every floor (Fig. 4). Each accelerometer was
196
attached to the corresponding column face just above the floor slab. Vertical accelerations at
197
intermediate locations on the floor slab and at the center column were also recorded by uniaxial
198
accelerometers attached to the bottom of the slabs. Details of the instrumentation used to
199
measure the response of CPP components are provided in Soroushian et al. (2013a). All data
9
200
discussed in this paper is archived and publicly accessible through the NEES Project Warehouse
201
(Ryan et al. 2013b,c,d).
202
Earthquake Simulation Plan
203
The building was subjected to various ground motions over three days of testing for the TP
204
configuration, two days of testing for the hybrid configuration, and one day of testing for the
205
fixed-base configuration. Out of 41 total earthquake simulations, 23 targeted 3D motion
206
including a vertical component. The achieved table accelerations encompassed a wide range of
207
shaking intensities and frequencies, which allowed the CPP system vulnerability to be critically
208
addressed. Table 1 lists a subset of motions that includes all 3D motions and any other motions
209
specifically mentioned in the papers. The majority are multi-component acceleration histories
210
recorded in previous earthquakes. Included for each motion is earthquake and station where
211
recorded (as applicable), basis for selection, theoretical scale factor (as applicable, described
212
below), system configuration, actual scale factor applied to the recorded motion during the test,
213
target acceleration or PGA, and ratio of recorded table acceleration to target acceleration. The
214
latter three statistics are provided separately for X/Y and Z components.
215
Different considerations factored into the selection of input ground motions for each test
216
configuration (Dao et al., 2012; Ryan et al., 2013a). This discussion focuses on the TP
217
configuration, which was the basis for many of the vertical ground motions applied throughout
218
the test program. For the TP configuration, a target spectrum was developed for a high
219
seismicity, Los Angeles site on site class D soil with SDS = 1.18g and SD1 = 0.71g. Ground
220
motions recorded in similar site conditions were selected to represent the service, design and
221
maximum considered earthquake (MCE) at the U.S. site (the first six shaded records of Table 1).
222
Amplitude scale factors (Theoretical Scale Factors in Table 1) were developed to minimize the
10
223
least square error between 1.3 times the 5% damped target spectrum and the SRSS of the 5%
224
damped spectra of the two horizontal components of motion, over a period ranging from 0.5 to
225
1.25 x Teff, where Teff = effective period of the isolation system. During testing, the theoretical
226
scale factors were applied to the service and design level motions, but MCE motions were
227
generally applied at a scale factor of 1.0 or the largest scale factor that could safely be
228
accommodated. Besides the motions for the U.S. site, additional motions were selected to
229
represent Japanese design practice, and to investigate the response of a base-isolated building
230
during long duration, long period subduction motions like those recorded in the 2011 Great East
231
Japan (Tohoku) Earthquake, which were of special interest to Japanese collaborators. Many of
232
these same ground motions were also applied to the hybrid configuration and the fixed-base
233
configuration to provide direct comparisons. Horizontal components applied to the fixed-base
234
configuration were generally applied at reduced scale to ensure the structural system did not
235
yield. In addition, two synthetic motions representative of the nuclear power industry’s extended
236
design basis for a Central and Eastern U.S. site (Vogtle) and design basis for a Western U.S. site
237
(Diablo Canyon) were selected and applied at various scale factors to the hybrid configuration.
238
While a vertical design spectrum was not developed for the test, the recorded components of
239
vertical ground motions were generally scaled proportionally to the horizontal components. In
240
some simulations, the vertical component was omitted (i.e. the target vertical acceleration was
241
zero) due to table limitations or to preserve the integrity of the CPP system. To interpret the
242
vertical motions that were imposed to the system, a vertical design spectrum and MCE spectrum
243
was developed to match the horizontal spectrum for the Los Angeles site mentioned earlier, on
244
the basis of NEHRP recommendations (FEMA, 2009). The NEHRP recommendations are based
245
on the findings of many studies investigating vertical to horizontal (V/H) spectral ratios such as
11
246
Bozorgnia and Campbell (2004). The 5% damped acceleration spectra of the realized (not target)
247
vertical motions applied during the test program are compared to the vertical design and MCE
248
spectra in Fig. 5. The realized vertical motions are mostly enveloped by the MCE spectrum with
249
the exception of the Northridge-Rinaldi ground acceleration, applied at a target scale factor of
250
88% (RRS88), which was replicated similarly in all three configurations. The realized RRS88
251
motion may be outside of what would normally be considered in design; however, larger vertical
252
components have been recorded in the near-fault region during prior earthquakes; for instance,
253
near the population center during the relatively moderate magnitude Christchurch Earthquake
254
(Bradley and Cubrinovski, 2011).
255
Testbed System Identification
256
White noise motions were applied to the building in the fixed-base configuration; these
257
motions included unidirectional (X or Y-direction) white noise at the beginning and end of the
258
experiment (2 repetitions), and 3D white noise before and after every earthquake simulation (10
259
repetitions). The normalized frequency responses of the recorded floor accelerations with respect
260
to the table accelerations (also known as transfer functions) were analyzed to estimate the
261
vibration periods and damping ratios of the natural modes of the testbed building in the
262
horizontal direction. The transfer functions were smoothed using a periodic Hamming window
263
with 50% overlap ratio. The periods and damping ratios corresponding to the fundamental
264
response modes were evaluated by curve fitting theoretical transfer functions to the measured
265
transfer functions using a least squares algorithm.
266
The average (over the repetitions) periods and damping ratios of the first three modes in each
267
horizontal direction, computed from the transfer functions of the average horizontal floor
268
acceleration (recorded at 3 columns, Fig. 4) relative to the average horizontal input acceleration
12
269
(recorded at 4 locations on the table), are listed in Table 2. These horizontal structural vibration
270
modes together comprise the majority of the mass participation in the horizontal direction. Slight
271
differences in the fundamental periods determined from 3D white noise compared to
272
unidirectional white noise were observed.
273
System identification in the vertical direction was performed using vertical accelerations
274
recorded at the center of 3 of the 4 slab quadrants (ASlabSE, ASlabNE and ASlabNW in Fig. 4), which
275
are referred to hereafter by location as SE, NE and NW. The transfer functions of these
276
accelerations (slab transfer functions) during 3D white noise were computed with respect to the
277
nearest accelerometer at the shake table level (ATSE, ATNE, ATNW in Fig. 4). Figure 6 illustrates
278
representative transfer functions. A single dominant peak can be observed in the transfer function
279
for the roof SE slab quadrant (Fig. 6(a)), which suggests that its response was dominated by
280
vibration in a single mode. However, multiple vibration modes contributed to the response at
281
some locations, as demonstrated by the multiple peaks in the transfer function for the 5th NE slab
282
quadrant (Fig. 6(b)).
283
The transfer function for each slab quadrant was processed using the previously described
284
techniques to identify a dominant participating modal period and damping ratio, but the
285
evaluation of vertical mode shapes was not attempted. This simplified approach was expected to
286
provide reasonable accuracy when the response is primarily single mode (Fig. 6(a)) but have
287
limitations for multi-mode response (Fig. 6(b)). The median (λ) and dispersion (β) – over the
288
white noise repetitions – of the natural period and damping ratio of each slab quadrant are
289
presented in Table 3. These results suggest that the first two vertical modes were localized
290
primarily to the roof and 5th floor east side slab quadrants, with periods of 0.13 sec (7.7 Hz) and
291
0.1 sec (10 Hz), respectively. This finding is consistent with Furukawa et al. (2013b), who
13
292
evaluated the vertical modes of this building using more precise techniques. The SE and NE roof
293
slab quadrants had a lengthened vibration period due to the substantial added mass on the east
294
side of the roof level. The cause of the increase for the 5th floor slab relative to lower floors is
295
unclear. The dominant period for the remaining slab quadrants was either 0.09 sec or 0.08 sec,
296
which suggests that one or two system modes contributed mainly to the slab vibration over the
297
rest of the building. Note that the dispersion in the observed vibration periods was close to zero,
298
meaning that little variation was observed from one white noise repetition to the next.
299
The slab damping ratios found by the approximate technique were more difficult to interpret,
300
with greater variation in the median damping in slab quadrants contributing to the same mode as
301
well as greater variation over the repetitions (based on the dispersion). The estimated damping
302
ratios were 1 – 1.5% for the first mode at 0.13 sec and 2 – 3.25% for the 2nd mode at 0.10 sec.
303
The estimated damping ratios for the vertical system mode(s) at 0.08 – 0.09 sec varied widely
304
from 3% to 10%, which suggests that the approximate technique was not very reliable. With
305
caution against drawing conclusions from this data, we observe a possible correlation between
306
slab damping ratio and modal period (damping ratios were lower for more flexible modes).
307
Overall, the observed periods of slab vibration modes ranged from about 0.08 to 0.13 sec (7.7
308
to 12.5 Hz). In comparison, the slab vibration frequencies of composite floor systems ranged
309
from 4 to 12 Hz in several newly constructed buildings, as measured through low level forced
310
vibration with an eccentric mass shaker (Hicks, 2004). Murray et al. (1997) and Allen and
311
Pernica (1998) suggested that natural frequencies of composite steel floor systems range from 5
312
to 9 Hz and 3 to 10 Hz, respectively, based on expert opinion. Boice (2003) found this range to
313
be 3 to 13 Hz based on floor response measured by heel drops, walking, and ambient vibration
314
tests for 103 case studies. Although the supplementary mass at the roof level may be viewed as
14
315
an unusual design configuration, the range of frequencies observed in this experiment is
316
consistent with the range of normal established in prior studies.
317
The slab vibration properties were also assessed during earthquake simulations. Because the
318
transfer function data was more difficult to interpret for earthquake simulations than for white
319
noise motions, a different procedure was used. First, the 5% damped response spectral
320
acceleration of the measured vertical acceleration at the middle of the slab quadrant, or
321
“measured slab spectrum” was computed. Next, the slab quadrant (Fig. 7(a)) was idealized as a
322
single degree-of-freedom (SDOF) oscillator (Fig. 7(b)) with period and damping ratio to be
323
determined. The response history and 5% damped response spectral acceleration of the idealized
324
SDOF system subjected to average vertical column acceleration (recorded at the center and
325
nearest corner column accelerometers, Fig. 4) were computed iteratively, with a search period
326
increment of 0.01 sec and damping ratio increment of 1%. (Since no sensor was installed at the
327
center column for the 2nd, 3rd and 4th floor slabs, the center column sensor at the 5th floor slab was
328
used instead.) The period and damping ratio of the idealized SDOF system that matched the
329
recorded slab acceleration were identified by a least squares fit of the measured slab spectrum to
330
the idealized SDOF system spectrum. This SDOF idealization procedure was expected to suffer
331
from similar limitations as the white noise transfer function procedure; nevertheless, the “best
332
fit” idealized SDOF system acceleration history and spectrum generally corresponded adequately
333
to the measured versions, where the correspondence was closer for slab quadrants with known
334
localized dominant modes at the 5th and roof level (e.g. Fig. 8(a)) compared to those with system
335
modes or multi-mode response (e.g. Fig. 8(b)).
336
The above process, which was computationally intensive, was applied to a subset of the slab
337
quadrants for four earthquake simulations. The selected simulations – all from the hybrid
15
338
configuration – included the synthetic Vogtle site motion at target scale factors of 75%, 125%
339
and 175% (VOG75, VOG 125, VOG175) and the 88% Northridge at Rinaldi (RRS88), which
340
represent the range of vertical intensity observed in the test program. In particular, only intensity-
341
related variation was present in the Vogtle simulations since the same motion was scaled to
342
different intensities. The resulting periods and damping ratios determined by the aforementioned
343
process are listed in Table 4.
344
Recognizing that a different methodology was applied and a different configuration was used
345
for earthquake excitation, the best fit slab vibration periods (Table 4) did not deviate from those
346
determined for white noise motions (Table 3) by more than 0.01 sec. Therefore, the ground
347
shaking intensity and the building configuration were concluded not to affect the slab vibration
348
period. The estimated damping ratios in most floor slabs increased for earthquake simulations
349
compared to white noise motions. This increased energy dissipation during earthquake excitation
350
could have resulted from mild nonlinearities (cyclic opening and closing of cracks in the
351
concrete slabs) during vertical vibration. Nonetheless, the trends for relative damping in the
352
various modes based on white noise were upheld, so that the damping ratios for earthquake
353
excitation also appeared to be lowest on the East side of the 5th floor and roof slabs, where the
354
most vulnerable components of the CPP system were located during the experiment. The nature
355
of the recorded structural response during 3D shaking is explored next.
356
Structural Response
357
Horizontal Floor Acceleration
358
As is well known, base isolation attenuates horizontal floor accelerations throughout the
359
structure at great benefit to nonstructural components. Representative floor acceleration profiles
360
(peak vector sum acceleration vs. height) recorded for each system configuration are shown in
16
361
Fig. 9(a), and in Fig. 9(b) the peak floor accelerations at each level have been normalized by the
362
peak table acceleration (PGA). Figure 9 confirms that horizontal accelerations were attenuated in
363
the isolation configurations relative to the ground acceleration, while they were amplified over
364
the height in the fixed-base configuration. The range of peak horizontal floor accelerations
365
recorded at the 5th and roof level was 0.12 - 1.12g for the isolated configurations and 0.46 - 1.22g
366
for the fixed-base configuration. The largest floor accelerations in all configurations were
367
recorded during the RRS simulation, where the isolation configurations were subjected to
368
horizontal PGA = 1.21g and the fixed-base configuration to horizontal PGA = 0.41g.
369
Other than mild cracking in the concrete slabs, some of which was present in the testbed
370
building prior to this experiment, structural damage was not observed in any of the simulations.
371
Horizontal floor acceleration, generally constrained to peak values well below 1g in the isolation
372
configurations (with the exception of the RRS simulation), is not believed to be the primary
373
cause of damage to the CPP components. Rather, the vertical slab acceleration was more closely
374
correlated to the CPP component damage that was observed (evidence is presented later). Thus,
375
the nature of the vertical slab vibration is discussed next.
376
Vertical Column and Slab Acceleration
377
To provide an overview of how vertical acceleration propagated from the shake table up
378
through the building columns and into the floor slabs, the peak accelerations recorded in the
379
shake table, columns, and middle of the floor slabs for every 3D simulation are investigated. The
380
data in this section has been processed with a modified filter cutoff frequency of 50 Hz to point
381
out higher frequency effects that were observed. Figure 10 presents the peak column
382
accelerations – absolute and normalized by vertical PGA - for the NW column in Fig. 10(a)-(b)
383
and for the SE column in Fig. 10(c)-(d). Column accelerations are also representative of floor
17
384
slab accelerations that would be recorded near the columns. Figure 11 uses the same format to
385
present the corresponding accelerations recorded at the middle of the floor slabs. The normalized
386
accelerations are equivalent to amplification factors of the PGA. The simulations are numbered
387
chronologically, where the first 10 correspond to the TP configuration, the next 10 to the hybrid
388
configuration, and the last three to the fixed-base configuration. The last data point in every
389
subplot is an average over all the simulations.
390
Figure 10 suggests that in general, column acceleration amplification factors were low and
391
insensitive to the height; that is, column acceleration did not significantly increase as the seismic
392
waves propagated from the base to the roof. This behavior was expected since the columns were
393
relatively rigid and could transfer the motion with little distortion. The average column
394
acceleration amplification factor was about 2 for the NW columns (Fig. 10(b)) and 2-3 for the SE
395
columns (Fig. 10(d)). Some exceptions to these general observations were noted; for instance,
396
during a few motions (GM #3 and 4 for both columns, and 7 and 10 for the NW column), the
397
acceleration amplification was notably higher. These instances of amplification, which were
398
limited to the TP configuration, arose due to local uplift/impact excursions that caused high
399
accelerations to propagate through individual columns. Consequently, amplification factors were
400
larger and increased significantly over the height.
401
The trends for floor slab acceleration amplification factors (relative to the shake table
402
acceleration) were quite different for the NW and SE floor slabs (Fig. 11). The average slab
403
amplification factor for the NW slabs was about 3 and did not increase much over the height of
404
the building (similar to what was observed in the NW column). The amplification factors for the
405
SE slabs, on the other hand, were notably higher, and increased steadily over the height of the
406
building (Fig. 11(d)). The average amplification factor varied from about 3 on the 2nd floor to
18
407
about 6 on the roof. As mentioned earlier, an amplification factor exceeding 6 was recorded in
408
one building in the Northridge Earthquake (Bozorgnia et al., 1998). The differences in response
409
of the NW and SE slabs appear to be related to period and damping trends identified earlier. For
410
instance, the NW slab dynamic properties – lower period and larger damping ratios that did not
411
vary much with height – were more favorable to vibration suppression, while the SE slab
412
dynamic properties – period increasing with height and damping ratio decreasing with height –
413
closely corresponded to the intensification of acceleration amplification factors in the upper
414
floors of the building.
415
To provide greater insight into the differences between the two slab locations, the spectral
416
accelerations computed from accelerations recorded in the SE and NW slabs at the 2nd through
417
roof level (with reference to the spectral acceleration for the recorded table acceleration) are
418
plotted for 3 representative input accelerations in Fig. 12: VOG175 and RRS88 that were
419
mentioned earlier, and 80% of 1978 Tabas at Tabas Station (TAB80). For VOG175 (GM #16)
420
and TAB80 (GM #9), the shake table spectral acceleration was maximized between 0.1 and 0.2
421
sec, a trend that was reflected in the majority of the ground motions used in the test program. As
422
noted earlier, the natural vibration periods of the NW slabs were consistently near 0.07 sec, while
423
the periods of the SE slabs increased from 0.07 sec at the 2nd floor to 0.13 sec at the roof. Thus,
424
the SE slabs at upper floors, with periods more aligned to the maximum spectral content of the
425
earthquakes, were generally more vulnerable to the ground input. This is reflected in Fig. 12,
426
where the spectral peaks for the SE slabs were generally shifted to the right (Fig. 12(a)-(b))
427
compared to the NW slabs for the same motions (Fig. 12(d)-(e)), especially at the roof level. The
428
shake table spectral acceleration for RRS88 (GM #3) was approximately constant in the period
429
range of 0.08 to 0.2 sec. This uniformly large intensity shaking produced similar intensity
19
430
spectral peaks for many of the floors – regardless of vibration period (Fig. 12(c), 12(f)) – and
431
likewise uniform peak acceleration amplification factors ranging from 4-6 in the SE slab and 2-4
432
in the NW slab (GM #3 in Fig. 11(b), 11(d)). In summary, slab acceleration amplification factors
433
were somewhat dependent on the frequency content of the ground acceleration, where flexible
434
slabs with vibration periods exceeding 0.1 sec were more vulnerable than stiffer slabs.
435
Influence of Isolation System on Vertical Amplification Factors
436
For low intensity vertical table accelerations, column acceleration amplification factors were
437
generally insensitive to the presence of the isolation system (Fig. 10). This is consistent with
438
Lew and Hudson (1999), who identified three examples where the vertical accelerations recorded
439
below and at columns above the isolation system were essentially the same. However, during
440
high intensity vertical accelerations, column acceleration amplification factors were affected by
441
the presence of the isolation system. The effect is best illustrated by comparing the column
442
accelerations recorded during RRS88 in each of the three configurations. The RRS88 table
443
acceleration was characterized by a large near fault fling pulse that appeared in the vertical
444
direction at about 8.6 s, producing peak vertical table accelerations between 1.06 and 1.26g (Fig.
445
13(a)). This vertical motion was transmitted differently in the three configurations, as evidenced
446
by accelerations recorded in the 5th story column near the roof level (Fig. 13(b)). For instance,
447
the table acceleration was transmitted directly to the 5th story column with little amplification for
448
the fixed-base configuration. On the other hand, the table acceleration was amplified by nearly a
449
factor of 7 in the 5th story columns for the TP configuration. During this simulation, every
450
bearing uplifted, causing high frequency acceleration spikes from the bearing impact to
451
propagate through the building. Somewhat lower amplitude high frequency spikes were also
20
452
observed in the column accelerations for the hybrid configuration, suggesting some looseness in
453
the rails before the CL bearings engaged in tension.
454
The slab acceleration amplification factors, on the other hand, were not much affected by the
455
presence of the isolation systems. Figure 13(c) shows that during RRS88, the peak accelerations
456
recorded in the middle of the roof slab – where much of the damage was observed - were
457
approximately the same for each building configuration (variation was less than 20%). In
458
addition, the high frequency signals appearing in the column accelerations were not transmitted
459
to the slabs. Rather, the slab accelerations are characterized by a relatively lower frequency
460
vibration at the slab fundamental frequency, which is approximately 7-8 Hz according to Table
461
4. Thus, the slab amplification factors were basically unaffected by system configuration; that is,
462
normalized slab accelerations were approximately the same for the TP configuration (GM #1-
463
10), the hybrid configuration (GM #11-20), and the fixed-base configuration (GM #21-23). Since
464
the CPP component damage was closely related to the accelerations recorded in the center of the
465
floor slabs (demonstrated in the next section), the vertically rigid base isolation systems
466
considered in this experiment neither helped nor hurt the system response in the vertical
467
direction. This observation is consistent with the conclusions
468
Ceiling-Partition-Piping Component Damage vs Structural Response
469
While seismic design engineers have generally assumed that nonstructural components will
470
be protected from damage in an isolated building, these tests (similar to Furukawa et al., 2013a)
471
showed that vertical acceleration need also be considered. Thus, an understanding of the target
472
horizontal and vertical acceleration demands that induce nonstructural component damage, and
473
hence the demands that should be targeted to prevent such damage, is needed. The earthquake
21
474
simulations from this test series, which induced a varied combination of horizontal floor
475
accelerations and slab accelerations, produced useful data in this regard.
476
We qualitatively evaluated the damage to CPP components by inspection of all available
477
video footage, and correlated the observed damage states to recorded peak demand parameters
478
for every earthquake simulation. The damage mechanisms and the extent of damage were very
479
similar for the three system configurations because, as discussed previously, similar peak
480
demands were observed in the three system configurations. Thus, the damage observations are
481
presented and discussed without further mention of the system configuration during which they
482
were observed. Note that, similar to other related studies (e.g. Ryu and Reinhorn, 2012), CPP
483
damage was not inspected or repaired after every earthquake simulation. Generally, repairs were
484
applied to the suspended ceiling and sprinkler piping at the end of each test day, but these
485
components were not restored to their original strength after RRS88 in the TP configuration (3rd
486
3D and 5th overall of 41 simulations), which inflicted significant damage. Furthermore, damage
487
to partition walls was not considered in the evaluation, for the following reasons. Partition wall
488
damage states are generally drift-sensitive, but the typical drift sensitive damage states were not
489
observed. New vertical induced damage states were observed, which are described in Soroushian
490
et al. (2013a), but these damage states could not be directly observed from the videos, and thus
491
could not be associated with a particular earthquake simulation.
492
For the purpose of damage evaluation, we developed three general damage ratings, each
493
classified by several behaviors that occurred alone or in combination. The behaviors associated
494
with the damage ratings are described in Table 5. Ceiling panel equivalent fallen areas were
495
based on Gilani et al. (2013), wherein the total ceiling area suspended from the 5th and roof
496
levels was about 83.6 m2 (900 sf) each. Partially dislodged ceiling panels were not considered
22
497
part of the equivalent fallen area. Soroushian et al. (2013b) defined piping damage to be
498
extensive when 15% of pipe hangers fail - equivalent to 1 pipe hanger in this experiment.
499
Likewise, Soroushian et al. (2013b) rated damage as extensive when the permanent rotation
500
across a grooved or threaded connection exceeds 2 to 4 degrees. Rotations were not measured in
501
these experiments, but since 2 to 4 degrees is very small, any visible permanent rotation was
502
considered to indicate permanent damage. Finally, note that the assigned damage rating was
503
generally determined by the most severe rating when the observed behaviors overlapped multiple
504
ratings.
505
Based on video inspection, damage to the ceiling-sprinkler piping systems suspended from
506
the 5th and roof floors was evaluated independently and assigned a damage rating from Table 5
507
for each earthquake simulation. All simulations that applied XY (horizontal only) input
508
acceleration to the building in an isolation configuration were excluded since these simulations
509
never induced any damage. The results of this inspection and ranking are plotted in Fig. 14;
510
distinct markers for each damage rating are plotted against peak horizontal floor acceleration (X-
511
axis) and peak slab acceleration recorded at the center of the NE or SE quadrant (Y-axis). As
512
shown in Figs. 10 and 11, the vertical accelerations at the slab centers were substantially larger
513
than those at the corners, and thus the vertical acceleration was not uniform throughout the floor.
514
However, the accelerations at the slab centers reasonably represented the observed damage in
515
these experiments since, for example, ceiling panels tended to become dislodged in large
516
concentrations near the slab centers (see Soroushian et al. 2013a for further information).
517
Figure 14 illustrates that both Slight and Moderate damage ratings occurred for horizontal
518
accelerations ranging from about 0.4g–1.25g (variation by a factor of 3); but for vertical
519
accelerations ranging from about 2g–4g (variation by a factor of 2). Furthermore, individual
23
520
damage rating data points overlapped the intensity measures in the horizontal direction more
521
than the vertical direction. These observations suggest that the damage ratings were more closely
522
correlated to the measured vertical accelerations than horizontal accelerations, because the
523
horizontal floor accelerations were constrained to relatively low levels, which is a typical
524
objective of base isolation. If comparable intensity accelerations had been measured in the
525
horizontal direction (e.g. 5g and above), damage to nonstructural components would surely have
526
been pervasive; however, such intensities are generally unlikely to be realized in the horizontal
527
direction due to yielding in the structural system.
528
Damage to the CPP components generally initiated at vertical accelerations in the range of 2-
529
3g. As a notable exception, moderate damage to the ceiling system was observed in the fixed-
530
base configuration during the Iwanuma (XY) simulation, which generated peak horizontal floor
531
accelerations of about 0.9g on the 5th floor and 1.15g on the roof with vertical accelerations very
532
near zero (Fig. 14). The Iwanuma motion, which was recorded during the 2011 off the Pacific
533
coast of Tohoku Earthquake, applied continuous strong shaking for several minutes that caused
534
eventual fatigue to the ceiling system. Aside from this outlier, the vertical acceleration ranges
535
associated with the initiation of the discrete damage ratings overlapped slightly. In summary, for
536
the system considered in this test, Slight damage initiated at vertical accelerations between 2-3g,
537
Moderate damage at vertical accelerations between 3-5g, and Extensive damage at vertical
538
accelerations above 5g. Therefore, the data preliminarily suggests that vertical slab accelerations
539
should be limited to 2g for continued post-earthquake operation of typical seismically detailed
540
suspended ceiling and piping systems. This conclusion should be validated with further data.
541
Significant differences in the response of the ceiling systems suspended from the 5th and roof
542
floors were observed. In particular, very few ceiling panels fell from the 5th floor and the
24
543
equivalent area of fallen panels rarely reached even 5% (for Slight damage) despite recording
544
similar intensity accelerations at the 5th and roof levels. The technical explanation for the
545
discrepancy in the ceiling response over the two floors is given in Soroushian et al. (2013a).
546
However, Moderate and Extensive damage ratings at the 5th floor were assigned based on
547
sprinkler head-ceiling panel pounding interaction, damage to the perimeter seismic clips/wall
548
molding, and damage to the piping system such as permanent rotations.
549
Discussion
550
Since CPP damage in this experiment has been directly associated with vertical slab
551
vibration, we examine whether this slab vibration was representative of realistic floor systems.
552
The slab acceleration amplification factors averaged 5 or 6 at locations directly above the
553
concentration of CPP damage, compared to 2 or 3 elsewhere in the structure (Fig. 11(b), 11(d)).
554
The slab vibration periods lengthened in the areas with greater amplification factors, wherein the
555
period lengthening was justified, and the vibration periods were within the range of normal.
556
However, as discussed, the data related to effective damping in the experiment was inconclusive.
557
Thus, further evidence is needed to conclude that the trends for slab damping ratios and related
558
acceleration amplification factors observed in the experiment was representative of realistic
559
systems.
560
With regard to current design, the current ICC-AC156 code (ICC, 2010) can be interpreted as
561
follows: nonstructural components are designed/qualified based on a vertical spectral
562
acceleration equal to 2/3 of the horizontal spectral acceleration. Flexible components are
563
subjected to a component amplification factor ap that represents the dynamic amplification of the
564
component relative to the attached structure. Since z/h is taken to be 0 in the vertical direction,
565
nothing in the code accounts for the amplification of the vertical acceleration from the ground
25
566
(table) to the middle of the floor slabs by a factor of 3 to 6, as was observed in the flexible-side
567
floor slabs in this experiment. However, selection of alternative levels of amplification for design
568
is difficult to justify until the factors that contribute to such amplification (slab vibration periods,
569
modal properties, and effective damping as a function of vibration intensities) are better
570
understood.
571
Conclusions
572
573
The major findings of this experiment are summarized below:

574
575
Vertical ground acceleration can be a significant source of damage to integrated ceilingpartition wall-piping (CPP) components.

During seismic input, the floor slab vibrations were dominated by single mode response,
576
with periods ranging from 0.07 to 0.13 sec. For many of the slabs, the slab spectral
577
acceleration was accurately estimated using a calibrated equivalent single degree-of-
578
freedom system. Because the slab response was determined by its vibration properties,
579
the associated CPP performance was insensitive to whether the building was base-
580
isolated.
581

The peak vertical acceleration was amplified by an average factor ranging from 3 to 6
582
from the table to the middle of the floor slabs. Current design/qualification procedures for
583
nonstructural components do not account for any such amplification, and are thus
584
unconservative.
585

The amplification factor increased as the slab vibration period increased. Slab vibration
586
periods were within normal ranges found by previous studies. The amplification factor
587
was also correlated to effective damping in the slab, but the interpretation of damping
588
was difficult and requires further validation.
26
589

From qualitative evaluation of available video footage, damage ratings for suspended
590
ceiling and piping systems were assigned for each 3D earthquake simulation in this
591
experiment. Because the horizontal floor accelerations were constrained to relatively low
592
levels by base isolation, the damage ratings were more closely correlated to vertical slab
593
acceleration than horizontal floor acceleration. CPP component damage initiated at slab
594
accelerations of about 2g. Damage was minimal for accelerations from 2-3g, moderate
595
for accelerations from 3-5g, and extensive for accelerations exceeding 5g. These metrics
596
establish vertical acceleration targets for achieving desired performance objectives.
597
Acknowledgements
598
This material is based upon work supported by the National Science Foundation under
599
Grants No. CMMI-1113275 and CMMI-0721399 and the National Institute for Earth Science
600
and Disaster Prevention (NIED) in Japan. Any opinions, findings, conclusions or
601
recommendations expressed in this document are those of the authors and do not necessarily
602
reflect the views of the sponsors. The authors recognize and thank the following companies for
603
providing product donations and technical support: Earthquake Protection Systems, Dynamic
604
Isolation Systems, Aseismic Design Company, USG Building systems, Victaulic, Tolco, Hilti,
605
Allan Automatic Sprinkler and CEMCO Steel.
606
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724
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725
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32
726
727
728
729
Tables
Table 1. Subset of Ground Motions Selected for Testing, along with Scale Factors, Target Peak
Acceleration (g), and Realized to Target Acceleration Ratios
Reason for
Earthquake Station
Selection
1979 Imperial Westmorland U.S. Service
Valley, U.S.
1994
Northridge,
U.S.
Rinaldi
Receiving
Station
(RRS)
1989 Loma
Prieta, U.S.
Theoretical System
Test Scale
Target PGA
Scale Factor Config.
Factor (X/Y, Z) g (X/Y, Z)
0.8
TP, Hybrid, 0.8, 0.8
0.17, 0.17
Fixed
Recorded to
Target PGA
Ratio (X/Y, Z)
1.14, 0.83
U.S. Design
0.88
TP, Hybrid
TP, Hybrid
Fixed
Fixed
Fixed
0.88, 0
0.88, 0.88
0.35, 0
0.35, 0.35
0.35, 0.88
0.73, 0
0.73, 0.72
0.29, 0
0.29, 0.35
0.29, 0.72
1.63, NA
1.66, 1.75
1.37, NA
1.40, 1.21
1.41, 1.48
Los Gatos
Pres. Ctr
U.S. MCE
1.09
TP
0.7, 0.7
0.42, 0.64
1.07, 1.07
Sylmar
Hospital
U.S. MCE
1.22
TP
1.0, 1.0
0.87, 0.52
1.32, 1.05
U.S. MCE
1.03
TP
TP
TP
TP
0.5, 0.5
0.8, 0.8
0.9, 0
1.0, 0
0.45, 0.33
0.72, 0.52
0.81, 0
0.90, 0
1.30, 1.09
1.21, 1.13
1.15, NA
1.15, NA
1999 Chi-Chi, TCU065
Taiwan
U.S. MCE
0.89
TP
TP
TP
0.5, 0
0.7, 0
0.8, 0
0.41, 0
0.57, 0
0.65, 0
1.11, NA
1.13, NA
1.14, NA
1940 Imperial El Centro
Valley, U.S.
1995 Kobe, JMA
Japan
1995 Kobe, Takatori
Japan
Japan design
motion
Japan design
motion
Japan nearfault
NA
TP, Hybrid 1.3, 1.3
0.28, 0.26
1.08, 1.07
NA
TP
1.0, 1.0
0.60, 0.34
1.14, 1.20
NA
TP
TP
1.0, 1.0
1.15, 1.0
0.75, 0.29
0.86, 0.29
1.06, 0.90
1.09, 0.97
2011 Tohoku, Iwanuma
Japan
Long duration
subduction
NA
TP, Hybrid 1.0, 0
Fixed
0.7, 0
0.42, 0
0.29, 0
1.41, NA
1.28, NA
Synthetic
Vogtle
Central/Eastern
U.S. Extended
Design Basis
NA
Hybrid
Hybrid
Hybrid
Hybrid
Hybrid
0.75, 0.75
1.0, 1.0
1.25, 1.25
1.5, 1.5
1.75, 1.75
0.33, 0.22
0.44, 0.29
0.55, 0.36
0.66, 0.43
0.77, 0.50
1.19, 1.00
1.19, 1.04
1.25, 1.03
1.30, 1.02
1.34, 0.98
Synthetic
Diablo
Canyon
Western U.S.
Extended
Design Basis
NA
Hybrid
Hybrid
0.8, 0.8
0.95, 0
0.78, 0.46
0.93, 0
1.17, 0.99
1.20, NA
1994
Northridge,
U.S.
1978 Tabas,
Iran
Tabas Sta.
(TAB)
730
33
Table 2. Fixed-Base Building Natural Periods and Damping Ratios in the Horizontal
Direction
Mode 1 X
Mode 2 X
Mode 3 X
Mode 1 Y
Mode 2 Y
Mode 3 Y
White noise X
Period
Damping
(s)
ratio (%)
0.652
3.30
0.204
1.62
0.112
3.31
n/a
n/a
n/a
n/a
n/a
n/a
White noise Y
Damping
Period (s)
ratio (%)
n/a
n/a
n/a
n/a
n/a
n/a
0.677
2.54
0.211
1.65
0.113
2.64
White noise 3D
Damping
Period (s)
ratio (%)
0.677
4.09
0.205
1.95
0.112
3.74
0.686
3.49
0.212
1.93
0.113
3.61
731
732
733
Table 3. Median (λ) and Dispersion (β) of the Natural Periods and Damping Ratios of the
Individual Slabs, Computed from the Transfer Functions of White Noise Motions
SE
Floor
Level
NE
NW
Period
Damping
Period
Damping
Period
Damping
(s)
Ratio (%)
(s)
Ratio (%)
(s)
Ratio (%)
Roof
λ
0.13
β
0.00
λ
1.00
β
0.22
λ
0.13
β
0.00
λ
1.50
β
0.25
λ
0.08
β
0.00
λ
7.50
β
0.95
5th
0.10
0.00
2.00
0.00
0.10
0.00
3.25
0.34
0.08
0.00
4.50
0.63
4th
0.09
0.00
3.00
0.69
0.09
0.00
4.00
0.38
0.08
0.00
5.00
0.69
3rd
0.08
0.00
6.00
1.32
0.08
0.00
2.75
0.30
0.08
0.00
9.50
0.80
2nd
0.08
0.00
7.50
1.32
0.08
0.00
4.75
0.47
0.08
0.00
10.00
0.00
734
34
735
736
737
738
Table 5. Ceiling-Piping Damage Ratings and Associated Behaviors
Damage
Rating
Slight
Moderate
Extensive
Behavior Description
Ceiling panels up to 5% equivalent area fall. Slight damage to panels at sprinklerheads due to pounding interaction (hole enlarged by 1 inch in any direction). Pipe
hanger surge clips pops out.
Ceiling panels between 5 and 20% equivalent area fall. More significant damage to
panels at sprinkler heads due to pounding interaction (hole enlarged by 1-2 inches in
any direction). 1 or 2 cross tees may fail, and a ceiling hanger wire may break.
Damage to perimeter seismic clips and wall molding is visible.
Ceiling panels exceeding 20% equivalent area fall. Large sections of the ceiling grid
are compromised, due to buckling, misalignment or connection failure. Very significant
damage to panels at sprinkler-heads due to pounding interaction (hole enlarged by
more than 2 inches). 1 or more pipe hangers break or are permanently deformed.
Permanent rotation of armover pipes is visible.
739
35
740
Figures
741
Figure 1. 5-story steel moment frame testbed set on triple pendulum bearings
742
Figure 2. (a) Plan and elevation view of structural framing, (b) supplemental concrete weight on
743
floors 2-5, and (c) supplemental steel weight on roof (dimensions in mm)
744
Figure 3. (a) TP bearing, (b) LR bearing, (c) CL bearing
745
Figure 4. Typical layout of accelerometers at the floor levels
746
Figure 5. 5% damped spectral acceleration for the vertical components of 3D motions compared
747
to design and MCE spectra based on NEHRP provisions
748
Figure 6. Representative slab transfer function for: (a) roof SE slab, (b) 5th NE Slab
749
Figure 7. (a) Depiction of single floor slab between columns, and (b) associated single-degree-of
750
freedom (SDOF) system idealization
751
Figure 8. Measured and idealized SDOF slab acceleration history and spectrum for (a) Roof SE
752
Slab for VOG 75 and (b) 3 SE Slab for RRS 88
753
Figure 9. Representative peak (vector sum) horizontal floor acceleration profile (acceleration vs.
754
level) for TP, hybrid, and fixed-base buildings for XY (horizontal only) and 3D input motions
755
Figure 10. (a) Peak acceleration in table and NW column, (b) normalized NW column
756
acceleration, (c) peak acceleration in table and SE column, (d) normalized SE column
757
acceleration; for each floor level and every 3D simulation in the NEES/E-Defense test
758
Figure 11. (a) Peak acceleration in table and NW floor slab, (b) normalized NW slab
759
acceleration, (c) peak acceleration in table and SE floor slab, (d) normalized SE slab
760
acceleration; for each floor level and every 3D simulation in the NEES/E-Defense test
rd
36
761
Figure 12. 5% damped spectral acceleration at the ground through roof levels for (a) SE slab –
762
VOG 175, (b) SE Slab – TAB80, (c) SE Slab – RRS88, (d) NW Slab – VOG175, (e) NW Slab –
763
TAB80, (f) NW Slab – RRS88
764
Figure 13. Recorded vertical accelerations at (a) table, (b) NE column near roof, and (c) NE roof
765
slab in each building configuration during RRS88.
766
Figure 14. Damage rating for select XY and every 3D simulation plotted against peak vector
767
horizontal acceleration in X and peak vertical acceleration in Y for (a) 5th floor and (b) roof floor
37
Figure
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Copyright Agreement
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STENG-3259
Journal of Structural Engineering (ASCE)
Keri Ryan
[email protected]
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Response to Reviewers Comments
Click here to download Response to Reviewers Comments: Response to Reviewers - Paper 1.pdf
Manuscript STENG-3259: Seismic Simulation of Integrated Nonstructural Systems at E-Defense, Part 1:
Influence of 3D Structural Response and Base Isolation
Authors: Keri L Ryan, PhD; Siavash Soroushian; Emmanuel Manos Maragakis; Eiji Sato; Tomohiro
Sasaki; Taichiro Okazaki
The authors thank the reviewers for their comments, and have made a sincere effort to address each
comment. We believe that the revisions generated by these comments have improved the quality of the
paper. Each comment is addressed in turn.
Reviewer #2: This paper discusses a full scale shake table study of a building equipped with an
integrated ceiling, partition, and fire sprinkler piping system. The building is tested as 1) isolated with triple
friction pendulum bearings, 2) isolated with a combination of lead plug and cross linear bearings, and 3)
fixed base. The study focus heavily on the effect of the vertical excitation.
The paper is well written and presents interesting and valuable experimental information from full scale
shake table testing. Nonetheless, before the reviewer can recommend publication, various issues need to
be addressed by the authors.
Technical Comments
1. The term "Integrated Nonstructural Systems" in the title of the paper is too broad. There are
innumerable types of nonstructural components in a building, but the present study focuses on a very
specific one: an "an integrated suspended ceiling-partition wall-sprinkler piping system." The title should
be revised to reflect exactly what is discussed in the paper. Similarly, the broad term "nonstructural
components" in the abstract, body, and conclusions of the paper should be replaced with a more
descriptive one.
The reviewer is correct that to generalize the response of the integrated ceiling-partition-piping system to
all nonstructural systems is misleading. Where used in the context of the specific system tested in this
paper, we have replaced the term “nonstructural component” or “nonstructural system” with “ceilingpartition-piping” or the abbreviation CPP, which is defined in the last paragraph of the introduction.
2. The reviewer does recognize that this is a study with a large scope; however, the fact that this paper
constantly refers to other papers by the authors for further test setup details and findings makes is very
inconvenient and frustrating to read. At times, it seems that this paper is an overview that discusses the
entire project thinly, rather than a research article that discusses a specific aspect of the study in depth.
The reviewer urges the authors to take this into consideration towards a more focused and improved
revised paper.
The reviewer’s impression has been carefully considered. We hope the reviewer appreciates the challenge
of both overviewing the experiment and discussing the specific aspect in depth within a reasonable length
paper. Somehow, the request to add more general details while at the same time suggesting the paper
reads like an overview rather than an in depth discussion was a little confusing. In the original
submission, we attempted to include all details necessary to understand the study. It would have been
helpful if the reviewer identified what details seemed to be missing. The reviewer has identified some
issues in later comments, which we have addressed, and we hope will also improve the readability of the
paper. Specifically, additional details have been added to describe the isolation systems (Lines 146 to
170) and the earthquake simulation plan (Lines 199 to 233) so that the description of the experiment may
stand on its own. Additional discussion of the uplift observed in the triple pendulum system and its
significance has been added (Lines 436 to 454). The discussion of horizontal-vertical coupling has been
removed, as it is not essential to the hypothesis of the paper (see response to comment 11).
Details of the testbed building configuration, structural system, and composite floor system provided in
the original submission are believed to be sufficient. Details of the ceiling-partition-piping system are
appropriately deferred to the companion paper, as the link to the 2nd paper is clear and our expectation is
that most readers will consult the papers jointly. Generally, we left in place references directing the reader
to other sources of information; these are provided as a courtesy and not considered a necessity to
understand the study.
3. "The shaking intensities that can cause damage to the nonstructural components are typically lower
than those that induce structural damage, and nonstructural components may be more sensitive than
structural components to vertical shaking." Provide references.
A reference has been added for the first part of the statement. Since we do not know of a good reference
for the second part, we have modified the statement, so that it now reads “The shaking intensities that can
cause damage to the nonstructural components are typically lower than those that induce structural
damage (Miranda, 2003), and the influence of vertical shaking on nonstructural components has not been
4. Line 36. Unsubstantiated. Has this been shown elsewhere? or is it a conclusion of the present
study? If the latter, then this sentence probably does not belong in the introduction.
Line 36 states: “Nonstructural components in base-isolated buildings are subjected to a different
proportion of horizontal to vertical accelerations than prescribed by seismic qualification tests, and so
the observations from component tests referenced above may not strictly apply.” We believe this
statement can be made in confidence based on the present state of knowledge, independently of the data
later presented in the paper. To clarify our thinking, we have added the following “Specifically, AC 156
prescribes that the horizontal acceleration is amplified over the building height while horizontal
acceleration in a base-isolated building is significantly reduced relative to the ground acceleration (ICC
2010)” (Lines 36 to 41). The vertical acceleration as prescribed by AC 156 should not be significantly
altered by the presence of an isolation system.
5.
Line 62 points to the work of Lew and Hudson (1999) that found that the vertical accelerations
transmitted to the building from the ground were not affected by the presence of base isolation system.
Compare and contrast the findings of the present study to that study.
Not much more can be said about the Lew and Hudson study. Three different examples were identified –
all characterized by small amplitude vertical shaking - suggesting that vertical accelerations recorded at
the foundation and at the columns above the isolation units were essentially the same. This is consistent
with our observations at low intensities, but not high intensities. However, our more important
observation is that the presence of the isolation system did not affect the slab vibration and thus the
nonstructural response. The buildings considered by Lew and Hudson did not have vertical
accelerometers located anywhere on the floor slabs.
The following statement has been added (Lines 430-435): “For low intensity vertical table accelerations,
column acceleration amplification factors were generally insensitive to the presence of the isolation
system (Figure 9). This is consistent with Lew and Hudson (1999), who identified three examples where
the vertical accelerations recorded below and at columns above the isolation system were essentially the
same. However, during high intensity vertical accelerations, column acceleration amplification factors
were affected by the presence of the isolation system.” The statement appears at the beginning of the new
section titled “Influence of Isolation System on Vertical Amplification Factors”.
6. Line 74. It is stated that a study by Furukawa et al. (2013a), also at E-Defense, and also full scale,
found that despite the large vertical acceleration amplification observed at mid-slab, no damage to
suspended ceilings, piping and sprinkler systems was observed. Clarify the novelty of the present study,
in comparison to that by Furukawa et al.
Also, offer an explanation as to why the study by Furukawa et al. observed no damage to these
nonstructural systems, while the present study did.
The study by Furukawa discussed a different building that incorporated Japanese-designed ceilingpartition wall-piping systems. Japanese design practice is very different from U.S. design practice and so
the response of the two types of systems is not comparable. Furthermore, the response of the CPP system
was not a focus of the study by Furukawa et al, which focused on anchored and unanchored medical
equipment. No details about the CPP system is given; it is only mentioned in passing. In addition, as
mentioned in the original submission, the building studied in Furukawa et al. is a reinforced concrete
building, which has different floor vibration characteristics than a steel frame building. We feel that the
study by Furukawa et al is quite clearly of different focus than our manuscript, and further discussion of
the distinction or novelty of our study is not warranted. However, the above points are clarified in the
revised statement “Disruption and damage to building contents and equipment (the main focus of the
study) was significant, but damage to the Japanese-designed suspended ceilings, piping and sprinkler
systems, and partition walls/sliding doors was not observed.” (Lines 79-82 of the revised manuscript).
We believe that damage was not observed to CPP in the Furukawa et al study due to differences in the
design of U.S. and Japanese CPP systems. However, without knowing details about the design of the
Japanese style CPP systems, it is not appropriate to speculate.
7. Line 176. "Many observers of the test program felt that this level of vertical input acceleration was
extreme or unrealistic." The reviewer is not convinced by the argument that follows this statement, i.e.
discussion on the vertical PGA and spectral accelerations of two Christchurch motions. The seismic
vulnerability of a system needs to be evaluated based on a number of motions for a specific hazard level
and site. What is the rationale for the selection of the ground motions used in this study? For example,
the Rinaldi Receiving Station motion from the 1994 Northridge earthquake is a near fault motion that is
well known for its high destructive potential. Fig. 4 shows vertical spectral accelerations in the order of 3g
over the 0.1-0.2 s period range; consequently, it would be a big surprise if this motion didn't result in
extensive damage to the ceiling system, which has its fundamental period in this range. But how realistic
is it to base conclusions regarding the seismic vulnerability of nonstructural components on an input
motion like this? Is there a realistic design vertical spectral acceleration that the authors are working with
to select/scale motions? The authors have yet to establish that the selected group of ground motions are
realistic/representative since this heavily bears upon the validity of the conclusions drawn in the study.
The authors may find the work of Bozorgnia and Campbell on v/h ratios and vertical response spectra
relevant.
The realism of the selected ground motions, especially the Rinald Rec. Station motion, has been a point of
concern among project investigators and stakeholders since the completion of the tests. This concern was
obviously manifested in some of the comments in the original submission, which perhaps came across as
“defensive”. We have completely reworked the presentation of the vertical ground motions taking into
account the reviewers’ comments. In the revised section, we have:



provided a basis for the selection and scaling of ground motions in the study;
on the basis of NERHP provisions (FEMA, 2009), developed vertical design and MCE spectra
corresponding to the horizontal design and MCE spectra that were used for the ground motion
selection; and
showed how the vertical ground motions realized in the experiment compared to the vertical
design and MCE spectra.
Please see the response to comment 8 for discussion regarding drawing conclusions to unrealistic ground
motions.
8.
The selection of realistic ground motions is one thing; the ability of the shake table to accurately
reproduce these is another. If the table does not track properly, the resulting motion on the table may be
significantly larger (and, consequently, even more unrealistic) than the input. For example, what is the
input vertical PGA of the RRS88 motion which resulted in measured vertical PTA ranging from 1.05 to
1.25g? or, how does the output vertical spectrum for RRS compare to the input one? Discuss the tracking
performance of the E-Defense shake table and its ability to reproduce the selected ground motions.
Realized to target peak acceleration ratios are provided in Table 1 (a new table) for completeness.
However, the ability of the shake table to reproduce the input motions is not the subject of this paper, and
should not be a criterion for evaluation of the technical merit of the paper. The question of whether the
realized motions were realistic may be relevant if conclusions are based on responses observed during
those motions. We feel that our revised presentation of the ground motion establishes that the realized
vertical ground motions were realistic except for Rinaldi. Regarding the Rinaldi motion, we have
provided the relevant facts and presented a neutral perspective, allowing the reader to draw his/her own
conclusion.
The remaining question is whether the conclusions in the paper depend on the potentially unrealistic
Rinaldi motion. This is not the case, since the conclusions are drawn on the basis of observations over the
range of horizontal and vertical input intensities, not on a single ground motion. For instance, damage to
CPP systems was observed in several motions aside from the Rinaldi motion. The slab amplification
factors were consistent across the range of shaking intensities. Furthermore, the damage thresholds and
damage ratings were established at intensities well below the Rinaldi motion.
9.
Lines 201-203. Was table rocking observed but ignored? Or, was table rocking not observed at all?
“Kasai et al. (2011) observed that rocking of the shake table can affect the natural frequencies and mode
shapes; however, the effect of rocking was ignored in the analysis presented here.”
We believe the comment about table rocking is misleading and therefore have opted to remove it. To
clarify, the paper by Kasai presented a new method to analyze the natural frequencies and mode shapes of
the building that accounted for the rocking of the shape table. Kasai’s method was first demonstrated on
the same testbed building examined in this paper, since the testbed was also used in an earlier experiment
by Kasai and collaborators. However, we chose to present the experimental frequencies and modes of the
building determined by traditional methods, because they more closely matched the frequencies/modes
determined by eigenvalue analysis of the calibrated building computational model.
To answer the reviewer’s question about rocking, the research team initially investigated table rocking in
detail as a possible source of “unusual” structural response, but ultimately dismissed it as insignificant.
Specifically, the recorded table rocking was input to the building computational model through multiple
support excitation, in addition to the horizontal/vertical shaking. No difference was observed in the
computational response of the building with and without table rocking input.
10.
Discuss the two isolation types in a bit more detail, focusing specifically on the aspects of their
design that effects the vertical response of the columns and slabs. For example, the TFP is practically
rigid in compression but provides zero tensile resistance, which, as the authors noted, can result in uplift
and pounding, and consequent acceleration spikes that propagate up the columns. How about the hybrid
lead plug + cross linear bearing system? What are these cross linear bearings like? Do they provide any
tensile resistance? If so, is their stiffness in tension equal to that in compression? How stiff are they in
comparison to the lead plug bearings? When the lead plug bearings displace laterally, they also move
downward; how do the cross linear bearing respond? Do they follow this motion, do they restrict it, etc.?
Additional details have been provided about the isolation systems (Lines 146 to 170), including those that
will clarify the vertical response characteristics. Furthermore, we have strengthened our justification that
for these vertically stiff isolation systems, the vertical accelerations transmitted to the CPP system
through the floor slabs was not affected by the presence of an isolation system – an issue that was raised
in the review of the companion paper. Specifically, we have demonstrated that high frequency
accelerations propagating through the columns were not transmitted to the floor slabs (lines 448-454 and
new Figure 13).
With regard to the compliance between the LR bearings and the CL bearings, this is a complex issue that
we believe was inconsequential to the overall vertical response of the isolation system. Therefore, we
choose not to address it specifically in the paper.
11.
Lines 310-312. "These sources of coupling have been investigated (Dao, 2012; Ryan 310 et al.,
2013a); they are not strictly an artifact of the test setup and they can be predicted by numerical simulation
(Dao and Ryan, 2013)". Could these sources of coupling be due to irregularities in the distribution of mass
and/or stiffness of this specific design? To the best of the reviewer's knowledge, other authors have not
noted such a coupling before.
Upon further consideration, we have decided to remove any mention of horizontal-vertical coupling from
this paper. While it is helpful to mention the horizontal-vertical coupling in the context of the overall big
picture of the experiment, the hypothesis of this paper is that the damage to the integrated CPP system
principally resulted from direct vertical shaking, and the coupling was of minor or no consequence to this
outcome. Since the coupling cannot be described, demonstrated, and validated satisfactorily in this paper,
it is better not to confuse the issue. Another paper focused on the horizontal-vertical coupling is currently
in review. The description of the observed horizontal acceleration has been modified accordingly (Lines
353 to 370 of the revised manuscript).
12. It is noted that damage to the integrated ceiling system was not evaluated or repaired after each
motion was run on the table, but rather at the end of each day. Furthermore, the system was never
restored to its original strength after RRS was run (which was very early on, based on Figs 9 and 10). If
so, then how can the authors know that it was a particular motion that caused damage, and it was not
compounding damage over the course of several tests? After all, the authors do point out the case of the
Iwanuma motion where fatigue and damage was caused to the ceiling system due the prolonged
shaking—even if the intensity of the shaking was very low.
The reviewer is correct that we cannot be sure that the damage states observed after each motion were not
influenced by cumulative damage over several tests. However, because of the effort and expense involved
in fabricating nonstructural systems for experimental testing, it is common to perform a series of shaking
tests where the initial state of the system is not restored after each test (e.g. Ryu and Reinhorn, 2012). We
feel that it is appropriate to report and interpret observations from sequential tests as long as the
conditions are clearly stated, which they have been here. The passage that states these important
conditions has been slightly reworded to state: “Note that, similar to other related studies (e.g. Ryu and
Reinhorn, 2012), CPP damage was not inspected or repaired after every earthquake simulation. Generally,
repairs were applied to the suspended ceiling and sprinkler piping at the end of each test day, but these
components were not restored to their original strength after RRS88 in the TP configuration (3 rd 3D and
5th overall of 41 simulations), which inflicted significant damage.”
Regarding the statement about the Iwanuma motion, the authors believe that fatigue was a factor in the
Iwanuma motion, but the intensity was not low. The paper states: “The Iwanuma motion, which was
recorded during the 2011 off the Pacific coast of Tohoku Earthquake, applied continuous strong shaking
for several minutes that caused eventual fatigue to the ceiling system.” The duration of strong shaking in
all of the other motions applied was very short by comparison.
13.
Did the authors investigate other intensity measures, besides horizontal and vertical acceleration
(Fig. 12), to see if those correlate better with damage to the ceiling system? It is not clear to the reviewer
what the failure modes of the ceiling system are.
The reviewer raises a valid question regarding the desire to find the best response measure to correlate
with the ceiling system response. Other intensity measures that we might have considered are: peak
spectral acceleration, spectral acceleration at the natural period of structure, peak floor velocity, etc.
However, we have several reasons for restricting the consideration to peak horizontal and vertical
accelerations recorded in the floors: 1) The component fragility functions that have been developed for
performance-based design and for use in the Performance Assessment Calculation Tool (PACT) work
with very general response measures, generally floor accelerations and story drifts, for simplicity. 2) We
wanted a general response measure, and not a response measure like period specific spectral acceleration
that would only be valid for this specific system. 3) We were trying to show, in a simple way, that the
CPP damage is more closely correlated to vertical shaking than to horizontal shaking. The failure modes
of the ceiling systems (damage states) are clearly stated in Table 4.
14. The establishment of damage states (Line 459) based on a very limited sample appears somewhat
arbitrary and unreliable. Please comment. Are there any other studies with data that could be used to
enhance confidence in the set damage state limits?
The reviewer is correct that it is premature to present conclusive damage states based on the limited data
considered in this paper. The text starting from Line 459 (now 528) has been reworded as follows: “In
summary, for the system considered in this test, Slight damage initiated at vertical accelerations between
2-3g, Moderate damage at vertical accelerations between 3-5g, and Extensive damage at vertical
accelerations above 5g. Therefore, the data preliminarily suggests that vertical slab accelerations should
be limited to 2g for continued post-earthquake operation of typical seismically detailed suspended ceiling
and piping systems. This conclusion should be validated with further data.”
Unfortunately, we do not think that other existing data sets can validate the vertically-induced damage
states at this time. The data set discussed in this paper appears to be unique with respect to the intensity of
vertical accelerations that were observed and/or the independence of horizontal and vertical accelerations.
For instance, for the data set discussed in Ryu et al. (2013), the specimen reached maximum frame
vertical accelerations of only 1.54g at mid-bay of the roof level. The data set discussed in Gilani et al.
(2010) did include larger vertical accelerations (e.g. nearly 6g recorded in the frame) but they were
accompanied by large horizontal accelerations, since horizontal and vertical accelerations were scaled up
proportionally (i.e. data points that would appear as a straight line in Figure 12 (now Figure 14)).
15. Paragraph starting on Line 482. "rigid components are designed/qualified for a vertical acceleration
equal to the vertical PGA, and flexible components are designed/qualified for a vertical acceleration equal
to 2.5 times the vertical PGA." The authors are urged to re-examine carefully Section 6.5.1 of AC156 and
revise this paragraph accordingly.
We hope that we have correctly interpreted the reviewer’s comment. There is no direct reference to
vertical PGA in AC156, but rather the horizontal spectral acceleration is the basis for deriving both
horizontal and vertical input motion. The statement has been revised to “nonstructural components are
designed/qualified based on a vertical spectral acceleration equal to 2/3 of the horizontal spectral
acceleration. Flexible components are subjected to a component amplification factor ap that represents
the dynamic amplification of the component relative to the attached structure. Since z/h is taken to be 0 in
the vertical direction, nothing in the code accounts for the amplification of the vertical acceleration from
the ground (table) to the middle of the floor slabs by a factor of 3 to 6, as was observed in the flexibleside floor slabs in this experiment.” (now lines 553-559)
16. Paragraph starting on Line 482. "Since 2.5 is a spectral amplification factor, nothing in the code
accounts for the amplification of the vertical acceleration from the ground (table) to the middle of the floor
slabs by a factor of 3 to 6, as was observed in the experiment. Thus, the current design/qualification
procedures for nonstructural components appear to be unconservative." As this conclusion leans on the
ground motions used in this study, the reviewer urges the authors to exercise caution.
We have removed the very strong statement “The current design/qualification procedures for
nonstructural components appear to be unconservative.” However, we wish to preserve the comment that
current codes do not account for amplification of vertical acceleration in the floor slabs. The amplification
factors of 3 to 6 that were observed on the flexible side of the building were not limited to one or two
motions of strong intensity; rather they were observed consistently over the diverse set of ground motions
with regard to horizontal and vertical intensity, duration, and frequency characteristics.
17. Table 4. Why were the damage states defined as such? Is there a monetary correlation to these
damage states?
We believe the rationale for the selection of the damage states is clearly described in the original
submission. “Ceiling panel equivalent fallen areas were based on Gilani et al. (2013), wherein the total
ceiling area suspended from the 5th and roof levels was about 83.6 m2 (900 sf) each. Partially dislodged
ceiling panels were not considered part of the equivalent fallen area. Soroushian et al. (2013b) defined
piping damage to be extensive when 15% of pipe hangers fail - equivalent to 1 pipe hanger in this
experiment. Likewise, Soroushian et al. (2013b) rated damage as extensive when the permanent rotation
across a grooved or threaded connection exceeds 2 to 4 degrees.” Basically, the damage states for our
CPP system are based on precedents established for individual components in previous experiments. We
chose not to use the commercial damage states in the PACT software, because they were not wellcorrelated to the range of observed damage in the experiment. For instance, the PACT extensive DS is
outside of the range of what we observed. In addition, there is a precedent to fine tune the damage states
to fit the patterns observed in the experiment. (e.g. Gilani et al. (2013) defined the transition to extensive
ceiling damage at 20% equivalent area of fallen ceiling panels, which matches well with their data).
18.
Figure 2. Check the dimensions in the E-W direction. They don't add up.
The correction to the figure has been made.
Editorial Comments
Line 229. "The SE and NE roof slab quadrants were more flexible due to the substantial added mass on
the east side of the roof level." Is it the flexibility or the period that the added mass increases?
The sentence has been reworded “The SE and NE roof slab quadrants had a lengthened vibration period
due to the substantial added mass on the east side of the roof level.”
References:



Federal Emergency Management Agency (FEMA), (2009). “NEHRP recommended seismic
provisions for new buildings and other structures”, FEMA P-750. Building Seismic Safety Council.
Miranda, E. (2003). ”Building Specific Loss Estimation for Performance Based Design”, Pacific
Conference on Earthquake Engineering. University of Canterbury, Christchurch, New Zealand,
2003.
Ryu, K.P., Reinhorn, A.M. and Filiatrault, A., (2012). “Full Scale Dynamic Testing of Large Area
Suspended Ceiling System”, 15th World Conference on Earthquake Engineering (15WCEE),
Lisbon, Portugal.
Reviewer #3: Seismic Simulation of Integrated Nonstructural Systems at E-Defense, Part 1: Influence of
3D Structural Response and Base Isolation
This paper is the first of two papers that summarize an impressive experimental study to investigate
dynamic response of nonstructural components. It is well written and comprehensive. A few questions
and comments are listed below:
* Was there a specific reason as to why the typical iterative feedback control algorithms were omitted
while replicating the target ground motions?
We thank the reviewer for the overall assessment of the series of papers. Regarding this first question,
during the test program, we were limited to 6 total days of testing and about 7 trials per day. We had to
choose between slowly ramping up one or two ground motions to reach a final target, or applying several
different motions at their target intensity. However, upon reworking this paper we have decided to
remove the statement about the feedback control algorithms as we believe it is unnecessary. Rather we
have chosen to provide additional data from which the reader can evaluate how the target motions were
replicated by the table.
* How was the scaling applied to obtain the given vector sum of horizontal accelerations? Was there
any scaling applied to the vertical component?
A common scale factor was applied to all components of the ground motion (including vertical) to
achieve desired targets in the horizontal direction. In the revised manuscript, the ground motion selection
and scaling is discussed in further detail, and the scaling procedure is clarified in lines 217-221.
*
Please check if in line 406 it is meant "structural" or "nonstructural damage"
We confirm that the intent is “non-structural damage” (now changed to “CPP damage”.)
*
A discussion about the structural damage, if any, observed in the experiments will be useful.
The following statement has been added for clarification: “Other than mild cracking in the concrete slabs,
some of which was present in the testbed building prior to this experiment, structural damage was not
observed in any of the simulations.” (lines 364-365)
*
Was there any correlation between the sudden acceleration peaks due to uplift and damage?
No, there was not correlation because the damage to CPP systems was determined by the slab vibration,
which was not affected by the uplift in the isolation system. An expanded discussion of the comparative
response of the three configurations to the largest vertical acceleration input, which includes the uplift
excursion in the TP configuration, has been added (lines 436 to 454).
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