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 2: Evaluation of Nonstructural Damage and Fragilities
--Manuscript Draft-Manuscript Number:
STENG-3260R1
Full Title:
Seismic Simulation of Integrated Ceiling-Partition Wall-Piping System at E-Defense,
Part 2: Evaluation of Nonstructural Damage and Fragilities
Manuscript Region of Origin:
UNITED STATES
Article Type:
Technical Paper
Section/Category:
Seismic Effects
Abstract:
A full-scale, five-story steel moment frame building in base-isolated and fixed-base
configurations was subjected to a number of ground motions using the E-Defense
shake table. In these experiments, more than 84 m2 (900 sf) of suspended ceiling with
lay-in tiles, 90 m (300 linear ft) of partition walls with individual lengths varying from 1.5
to 9.8 m (5 to 32 ft), and 3 sprinkler branch lines were installed below the 5th and 6th
(roof) floors of the building. Since the horizontal floor accelerations were generally
constrained to relatively low values by the base isolation system, several damage
states related to vertical floor system acceleration were observed. One key observation
is that use of lateral bracing with compression posts did not improve the seismic
response of suspended ceilings when subjected to strong vertical excitation.
Acceleration amplification factors of the ceiling-partition-partition components relative
to structural floor accelerations were computed. The code prescribed amplification
factors for the design of nonstructural components was consistent with the observed
amplification in the horizontal direction, but unconservative in the vertical direction
because the code neglects the additional amplification produced by slab vibration.
Corresponding Author:
Keri L Ryan, PhD
University of Nevada, Reno
Reno, NV UNITED STATES
Corresponding Author E-Mail:
[email protected]
Order of Authors:
Siavash Soroushian
Emmanuel Manos Maragakis
Keri L Ryan, PhD
Eiji Sato
Tomohiro Sasaki
Taichiro Okazaki
Gilberto Mosqueda
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
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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 Hoehler
Hilti Corporation
[email protected]
Hoehler has been involved in experimental research on seismic response of ceiling
systems and has expertise to evaluate structural and nonstructural responses
observed in this research.
Opposed Reviewers:
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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 2 - Revision 1 - No Figures.docx
1
Seismic Simulation of an Integrated Ceiling-Partition Wall-Piping
2
System at E-Defense, Part 2: Evaluation of Nonstructural
3
Damage and Fragilities
4
Siavash Soroushian,a) E. "Manos" Maragakis,b) Keri L. Ryan,c) Eiji Sato,d)
5
Tomohiro Sasaki,e) Taichiro Okazaki,f) Gilberto Mosquedag)
6
Abstract
7
A full-scale, five-story steel moment frame building in base-isolated and fixed-base
8
configurations was subjected to a number of ground motions using the E-Defense shake
9
table. In these experiments, more than 84 m2 (900 sf) of suspended ceiling with lay-in tiles,
10
90 m (300 linear ft) of partition walls with individual lengths varying from 1.5 to 9.8 m (5 to
11
32 ft), and 3 sprinkler branch lines were installed below the 5th and 6th (roof) floors of the
12
building. Since the horizontal floor accelerations were generally constrained to relatively low
13
values by the base isolation system, several damage states related to vertical floor system
14
acceleration were observed. One key observation is that use of lateral bracing with
15
compression posts did not improve the seismic response of suspended ceilings when
16
subjected to strong vertical excitation. Acceleration amplification factors of the ceiling-
17
partition-partition components relative to structural floor accelerations were computed. The
a)
Post-doctoral Scholar, Dept. of Civil and Environmental Engineering, University of Nevada, Reno, MS 0258,
Reno, NV 89557-0258
b)
Dean of Engr., Dept. of Civil and Environmental Engineering, University of Nevada, Reno, MS 0256, Reno,
NV 89557-0256
c)
Assoc. Prof., Dept. of Civil and Environmental Engineering, University of Nevada, Reno, MS 0256, Reno,
NV 89557-0258
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
g)
Assoc. Prof., Dept. of Structural Engineering, University of California, San Diego, 9500 Gilman Dr. MC0085,
La Jolla, CA 92093-0085
18
code prescribed amplification factors for the design of nonstructural components was
19
consistent with the observed amplification in the horizontal direction, but unconservative in
20
the vertical direction because the code neglects the additional amplification produced by slab
21
vibration.
22
Keywords: nonstructural components, vertical ground motion, shake table testing, suspended
23
ceiling, sprinkler-piping, fragility functions
24
Introduction
25
The performance of critical facilities such as hospitals and fire stations during an
26
earthquake depends not only on structural systems, but also on the functionality of
27
nonstructural components (FEMA E-74, 2011). The shaking intensities that can cause
28
damage to the nonstructural components are typically lower than those that induce structural
29
damage (Miranda, 2003). Also, nonstructural systems almost always represent the major
30
portion of the total investment in buildings (Whittaker and Soong, 2003). Therefore,
31
economic losses in buildings due to nonstructural damage or malfunction can be much larger
32
than those directly related to structural performance (Taghavi and Miranda, 2003).
33
In the 2010 Chile earthquake, few buildings suffered structural damage, but economic
34
loss and damage of nonstructural systems was extensive in commercial, residential, office
35
and industrial buildings (Miranda et al., 2012). Damage to the acoustical suspended ceiling
36
systems including fallen ceiling panels was widespread; in some cases, 100% of the ceiling
37
panels fell. Several types of damage such as leakage at pipe joints, failure of braces, and
38
breakage of sprinklers were observed on fire sprinkler piping systems. Damage to partition
39
walls lacking adequate connections to the structure and/or adequate gaps was observed
40
during the earthquake (Miranda et al., 2012). Following the 2011 off the Pacific coast of
41
Tohoku Earthquake, 714 structures were inspected. Major portions of ceiling systems had
2
42
collapsed, and widespread water leakage was observed in inspected buildings due to
43
extensive damage to fire protection systems (Mizutani et al., 2012). These observations from
44
real earthquakes are insightful, but often lack measured responses (both structural and
45
nonstructural) that can correlate the complex system behavior and damage states to demand
46
parameters.
47
Several component and subsystem level nonstructural experiments have been conducted
48
in recent years. Examples of previously tested components include ceiling subsystems with
49
different sizes and aspect ratios (Badillo-Almarez et al., 2007; Gilani et al., 2010), piping
50
assemblies and subsystems (Zaghi et al., 2012; Tian et al., 2012), and individual partition
51
walls in different configurations and boundary conditions (Retamales et al., 2012).
52
However, these component and subsystem level experiments may not accurately reflect
53
the following influences of real buildings on nonstructural response: realistic input excitation,
54
interaction between different types of nonstructural components, realistic boundary
55
conditions, and floor system vibration. Furthermore, with base isolation, which is often
56
chosen to preserve functionality of critical buildings – the nonstructural components are
57
subjected to a different proportion of horizontal to vertical accelerations than prescribed by
58
seismic qualification tests, and so the observations from component tests may not strictly
59
apply. In summary, the limited quantitative data collected from past earthquakes and the
60
limitations of subsystem level experimental studies to reproduce structural and nonstructural
61
system interaction effects necessitates system level experiments and analytical tools to
62
facilitate a better understanding of the seismic response of nonstructural systems. The
63
response of nonstructural components observed in realistic building testbeds for various
64
structural systems that can replicate both horizontal modal effects and floor system vibration
65
is of particular interest.
3
66
As part of a collaborative research project between Network for Earthquake Engineering
67
Simulation (NEES) and the National Research Institute for Earth Science and Disaster
68
Prevention (NIED) of Japan, system-level full-scale shaking experiments of a 5-story
69
building were conducted at E-Defense. The investigation of the ceiling-partition-piping
70
(CPP) component performance in this building was led by the NEES Grand Challenge
71
project “Simulation of the Seismic Performance of Nonstructural Systems”. This is the
72
second of two related papers. In the first paper (Ryan et al., 2013a), the 3-dimensional (3D)
73
structural response in both base-isolated and fixed-base configurations was evaluated, and the
74
relation between severity of CPP damage and horizontal floor and vertical slab accelerations
75
was identified. In this paper, we evaluate the response of the integrated ceiling, partition, and
76
fire sprinkler piping systems, which were installed in the building throughout the test series.
77
Specifically, various observed CPP damage states – many of which were induced directly by
78
strong floor slab acceleration – are identified, and fragility data for these damage states are
79
developed where applicable.
80
Experimental Setup
81
The testbed structure (two base-isolated configurations and one fixed-base configuration)
82
that housed CPP components was described in Ryan et al. (2013a). A partition-ceiling-
83
sprinkler piping subassembly was designed and installed in nearly identical configuration
84
below the 5th and 6th floors of the testbed. Further details about these components are
85
described here.
86
Suspended Ceilings
87
An approximately 84 m2 (900 sf) lay-in-tile suspended ceiling system was designed for
88
each floor that worked around existing boundaries (Fig. 1). However, the ceiling area was
89
interrupted at two locations (total area of 3 m2 or 34 sf) that were impeded by vertical trusses
4
90
used to measure story drifts. The ceilings were installed in the test frame per ASTM
91
E580/E580M-11ae1 standards (ASTM, 2011). The grid was constructed using the heavy-duty
92
USG DONN 24 mm (15/16 in) exposed tee system. Main runners and cross tees were aligned
93
as shown in Fig. 1. The main runners were supported by 12-gauge Hilti X-CW suspension
94
wires spaced 1.2 m (4 ft) apart; additional wires supporting all perimeter grid pieces were
95
placed within 8 in of the partition wall faces. The plenum height (the distance between the
96
bottom of the structural slab and the ceiling system) was 0.9 m (3 ft).
97
A 22 mm (7/8 in) wall molding was attached to the perimeter partition walls. On the
98
North and East sides, the main runners and cross tees were attached tight to the wall molding
99
using USG/ACM7 seismic clips with one partition attached screw and one top hole screw to
100
prevent movement of the ceiling grids (Fig. 2(a)). On the South and West sides, a 19 mm (3/4
101
in) clearance was provided between the main runners/cross tees and the wall molding. This
102
connection used the same seismic clip, but with the second screw attached at the middle of
103
the clip slot to allow the grid members to float freely (Fig. 2(b)). At the hatched areas in Fig.
104
1(a), heavier gypsum board panels were used to simulate the weight of light fixtures.
105
ASTM E580/E580M-11ae1 (ASTM, 2011) requires seismic bracing to be included for
106
ceiling areas larger than 93 m2 (1,000 sf). To compare the behavior of braced and unbraced
107
ceiling systems, seismic braces were installed only on the 6th floor ceiling, while all other
108
ceiling details were identical on both floors. From here forward, the 5th and 6th floor ceilings
109
are understood to be the components suspended from the 5th and roof slabs, respectively.
110
Each seismic brace consisted of: 1) a system of splay wires and 2) a USG/VSA30/40
111
compression post. The seismic braces were placed at 3.7 m (12 ft) on center, in each
112
direction, with the first set occurring within 1.8 m (6 ft) of the wall face. Four wires splayed
113
at 90° from each other were attached to the main runner within 51 mm (2 in) of the
5
114
compression post (Fig. 3). Due to the connection constraints, steel stud compression posts
115
were used instead of VSA30/40 compression posts when the posts were attached to structural
116
girders. In one location on each floor, a 2-way steel stud rigid brace was used in place of two
117
of the splay wires due to space constraints.
118
Fire Sprinkler Piping
119
A standard Schedule 40 piping system was attached to the testbed building per NFPA 13
120
(NFPA, 2011). The 5th and 6th floor piping systems (suspended from the 5th and roof slabs)
121
included one 64 mm (2.5 in) diameter main run and three (North-South) 32 mm (1.25 in) and
122
25 mm (1 in) diameter branch lines per floor connected through a 76 mm (3 in) diameter riser
123
pipe (Fig. 4). All connections on the riser and the main run were groove fit, while the rest of
124
the connections were threaded. Branch Lines 1 and 2, each with three 305 mm (12 in) drops,
125
incorporated armover and straight drops, respectively. For Branch Line 3, Drop 1 was a 305
126
mm (12 in) straight drop, while a Victaulic Aquaflex flexible hose was used at Drop 2 (Fig.
127
5(a)). At Drop 1 of each branch line, a 51 mm (2 in) oversized ring was used to separate the
128
ceiling panel from the sprinkler heads (oversized gap configuration, Fig. 5(b)), while a
129
minimal gap between the ceiling panel and sprinkler head was provided at the rest of the drop
130
locations (no gap configuration, Fig. 5(c)).
131
Lateral resistance was provided by: inclined 25 mm (1 in) diameter longitudinal and
132
lateral pipe sway braces on the main run near the riser pipe (Fig. 6(a)), a lateral pipe sway
133
brace at the end of the main run, and two longitudinal braces at the end of the riser pipe
134
below the 5th floor slab. The ends of the branch lines were restrained with two diagonal splay
135
wires to limit the lateral movement (Fig. 6(b)).
136
Partition Walls
6
137
Approximately 91 m (300 ft) of typical light gauge steel studded gypsum partition walls
138
with individual lengths varying from 1.5 to 9.8 m (5 to 32 ft) were installed in the testbed.
139
The full height partitions were approximately 2.7 m (9 ft) tall. The partition wall details were
140
selected based on the most commonly used commercial and institutional partition walls.
141
Figure 7 presents the plan view of 4th and 5th floor partition walls (understood to be partitions
142
between 4th/5th slabs, and 5th/roof slabs, respectively) of the testbed building. The black and
143
gray labels refer to 4th and 5th floor partition walls, respectively.
144
All partitions were constructed using either 18 mm (0.7in) (350S125-18 and 350T125-18)
145
or 30 mm (1.2 in) (350S125-30 and 350T125-30) studs and tracks with a single ply 13 mm
146
(1/2 in) thick gypsum board on each side of the wall. The 30 mm (1.2 in) thick studs and
147
tracks corresponded to institutional detailing (Retamales et al., 2012), while the 18 mm (0.7
148
in) thick studs and tracks corresponded to commercial detailing. The institutional T-wall and
149
corner connection details incorporated additional studs and more frequent placement of
150
screws than the commercial connection details, as shown in Fig. 8.
151
Full connection detailing was provided for 4th floor partitions while slip track connection
152
detailing was provided for 5th floor partitions. In a full connection, the studs and gypsum
153
boards are fully attached to the top and bottom tracks. Due to the deflection of floors (and top
154
tracks) under live loads, the screws attaching the gypsum boards or studs to the tracks can
155
tear out. To prevent this type of damage, slip track connections have been designed to allow
156
vertical deflection of the top track relative to the gypsum boards and studs. Slip track details
157
can also accommodate horizontal drift with minimal damage. The typical top and bottom
158
details of the full and slip track connections that were installed in this experiment are
159
presented in Fig. 9.
7
160
On the North-East side of the building, a room was constructed with self-standing, partial
161
height partitions on the 4th and 5th floors to house hospital and office contents, respectively
162
(Fig. 10(a)). Full connection detailing was used at the bottom of these partitions while the top
163
boundary, positioned below the ceiling, was unrestrained. Also, bulkhead partitions were
164
built around the drift measurement trusses (at two locations on each floor as shown in Fig.
165
1(a)) to provide realistic ceiling boundary conditions. For the bulkhead partitions, only the
166
corner studs were extended down to the floor to provide access to the trusses (Fig. 10(b)).
167
Instrumentation
168
The table accelerations and the responses of structural and CPP components were
169
monitored by nearly 400 sensor channels (not including the isolation system response, when
170
applicable) recorded at a sampling frequency of 1000 Hz. A 4-pole low-pass Butterworth
171
filter with a cutoff frequency of 25 Hz was applied to all recorded responses.
172
Details of the instrumentation used to measure the structural response (horizontal floor
173
and vertical slab accelerations) are provided in Ryan et al. (2013a). Accelerometers spaced
174
regularly on the ceiling grid members (Fig. 11(a)-(b)) measured grid acceleration. Additional
175
sensors measured accelerations on select ceiling panels and grid points that were seismically
176
braced (Fig. 11(c)-(d)). Displacement transducers on the floating side of the ceiling perimeter
177
measured the movement of the ceiling system relative to the partition walls (Fig. 11(e)).
178
Accelerometers were also placed directly on the sprinkler pipes to measure their
179
longitudinal and transverse movement (Fig. 12(a)-(c)). Displacement transducers located on
180
the main run pipe (near the second branch line) measured the movement of the pipes relative
181
to the structure (Fig. 12(d)).
182
Several uniaxial accelerometers were installed on the full height partitions at or near the
183
ceiling elevation. These accelerometers measured the out-of-plane partition accelerations,
8
184
which corresponded to the acceleration imposed to the ceiling system near the boundaries
185
(Fig. 13(a)). Accelerometers in each horizontal direction were installed on top of the partial
186
height self-standing partitions (Fig. 13(b)). All data discussed in this paper is archived and
187
publicly accessible through the NEES Project Warehouse (Ryan et al. 2013b,c,d).
188
Seismic Response of Suspended Ceiling System
189
The building was subjected to various ground motions over six total days of testing for
190
two base-isolated and one fixed-base building configuration. Out of 41 total excitations, 23
191
were 3D including a vertical component. The achieved or realized excitations encompassed a
192
wide range of shaking intensities and frequencies, which allowed the CPP vulnerability to be
193
critically addressed. The damage mechanisms and the extent of damage were very similar for
194
the three system configurations because, as discussed in Ryan et al. (2013), similar peak
195
demands were observed in the three system configurations and vertical slab vibration was
196
insensitive to the presence of an isolation system. Thus, the damage observations are
197
presented and discussed without further mention of the system configuration during which
198
they were observed, but rather interpreted relative to the horizontal floor acceleration and
199
vertical slab acceleration to which they are closely correlated. The ceiling-piping-partition
200
system was repaired after each test day, but it was never restored to its original configuration.
201
Thus, unless otherwise noted, the results pertaining to repaired CPP components are not
202
reported in this study.
203
A key aspect of the ceiling response is the acceleration of the ceiling components (grid,
204
compression post, and panel) relative to the structural systems to which they are attached
205
(column or slab). Based on the recorded sensor data described previously, Table 1 reports
206
acceleration amplification factors (peak ceiling member acceleration normalized by peak
207
column acceleration). Table 1 also reports peak slab acceleration normalized by peak column
9
208
acceleration to highlight the effect of slab flexibility. The statistics (max, min and median)
209
are based only on valid simulations, which include all 41 simulations for the 5th floor ceiling
210
and only the first 4 simulations for the 6th floor ceiling, because appreciable damage was
211
observed in the 5th simulation. In Table 1, “column” refers to three column sensors while
212
“slab” refers to slab sensors on the given floor, valid only for vertical acceleration.
213
The component amplification factor ap in Eq. 13.3-1 of ASCE 7-10 (2010) accounts for
214
possible amplification of component horizontal response relative to the attached structure due
215
to the inherent component flexibility. The maximum recommended amplification is ap = 2.5
216
for components that are considered flexible; ap can be interpreted as spectral amplification of
217
2.5 relative to the peak column acceleration. Table 1 indicates that the average horizontal
218
(XY) amplification observed during the experiment was 2.6 and 2.8 for the unbraced and
219
braced ceiling, respectively. The observed acceleration correlates well with ap considering
220
that horizontal pounding of the ceiling panels against the grid members amplified the
221
accelerations in some motions.
222
In the vertical direction, the maximum component amplification can be interpreted as: 1)
223
the same value as for the horizontal direction (ap = 2.5) per ICC-AC156 (ICC, 2010) or 2) ap
224
= 2.67, which is the ratio of the constant to short period spectral acceleration (0.8 CV SDS/0.3
225
CV SDS) (ASCE, 2010). As discussed in Ryan et al. (2013a), this component amplification
226
assumes that the vertical ground acceleration is transmitted directly to the nonstructural
227
components, and does not account for amplification of the vertical acceleration as it travels
228
from the ground through the structure to the attachment point of the nonstructural component.
229
In this study, we evaluate component amplification by comparing the peak vertical
230
acceleration of the CPP component to the peak acceleration recorded in the nearest column.
231
Even this definition of component amplification is unconservative, since the vertical
10
232
acceleration was somewhat amplified as it traveled from the shake table to the columns
233
(Ryan et al., 2013a).
234
The median amplification of ceiling vertical acceleration (sensors mounted on the ceiling
235
grid members compared to sensors mounted on the columns) observed in the experiment was
236
4.2 and 5 for the unbraced and braced ceiling, respectively, which is higher than ap by
237
approximately a factor of 2. However, the median vertical amplification of the slab relative to
238
the column was 3 and 5.7 for the 5th and 6th floor, respectively. This suggests that the
239
majority of the vertical acceleration amplification was due to vibration of the floor slab
240
relative to the columns and not increased acceleration in the ceiling relative to the slab. An
241
example of such a trend during 3D-Superstition Hills-Westmorland (TP configuration) was
242
presented in Fig. 14. Some variation in slab acceleration amplification is expected as a
243
function of the slab vibration frequency, which has been shown to significantly vary among
244
realistic floor systems, both in general and in this experiment (Ryan et al., 2013a). Therefore,
245
a constant amplification factor may be inadequate to account for vertical acceleration
246
amplification caused by slab flexibility. The median vertical amplification of the ceiling
247
panel relative to the column sensors was 3.6 and 12 for the unbraced and braced ceiling,
248
respectively. The higher ceiling panel amplification factors in the braced ceiling will be
249
shown to result from pounding of the panels on the ceiling grid, and suggests that anchorage
250
design forces of hanger wires may need to be increased to account for this pounding.
251
In any individual simulation, a maximum of three panels (1%) from the unbraced (4th
252
floor) ceiling were displaced or fell to the floor while up to 40% of the panels in the braced
253
(6th floor) ceiling were displaced and/or fell. Most of the damage was located below the
254
middle of the North-East and South-East slabs, which supported the supplementary roof
255
weight. The condition of the braced and unbraced ceiling after the first 3D-Northridge-
11
256
Rinaldi simulation is compared in Fig. 15. In this simulation, which produced the first
257
significant damage to the ceiling system, the vertical peak table acceleration was 1.2g, which
258
led to peak acceleration of 6.8g on the 5th slab and 6.4g on the roof slab. While the chance of
259
a life threatening injury due to the fallen ceiling panels was low, serious disruption was
260
observed in the office room. Most of the fallen panels were damaged to the extent that
261
replacement would be desirable. In addition, fallen ceiling panels weakened the grid system
262
and increased the chance that more extensive grid repair was necessary. However, over the
263
course of the test program, some of the cross tee sections failed but the main runners always
264
remained intact.
265
In general, the accelerations in all three directions at the slab level, which represent the
266
input excitation to the ceiling system, were slightly higher at the roof slab than the 5th floor
267
slab. However, the input acceleration alone does not explain the difference in damage;
268
observed accelerations in each ceiling suggest that in this experiment, the compression posts
269
used in the lateral bracing increased the damage to the ceiling system.
270
Figure 16 shows the vertical acceleration of a ceiling panel (C9) and ceiling grid (C4)
271
(see Fig. 11(a)) measured for a moderate excitation (3D-Superstition Hills Westmorland/TP
272
configuration) executed prior to the occurrence of ceiling damage. In the unbraced ceiling,
273
little amplification of the panel acceleration relative to the grid was observed (Fig. 16(a)),
274
which suggests that these two components moved together. However, in the braced ceiling,
275
the acceleration of the ceiling panel was significantly amplified relative to the compression
276
post attachment location (Fig. 16(b)), which suggests that the panel pounded on the grid
277
system. The median acceleration amplification values presented in Table 1 imply the same
278
trend.
12
279
The acceleration trends observed in Fig. 16 are explained as follows: consider the
280
diagram of the ceiling system in Figs. 16 and 17, where the vertical acceleration of the slab,
281
grid, and panel are labeled Aslab, Agrid, and Apanel, respectively. Figure 17 depicts the unbraced
282
ceiling, which was supported only by hanger wires. When the hanger wires were in tension
283
(case 1), the accelerations of the slab, grid and panel were the same. However, when the
284
hanger wires were loose (case 2), which could have been initiated by downward slab
285
acceleration of more than 1g while the panels and grid system were limited to a maximum of
286
1g downward acceleration, the slab acceleration differed from that of the panels and grid
287
system. Since the panel and grid system had almost the same acceleration, the panels
288
remained in place between the grid members and the probability of dislodging panels was
289
low.
290
Figure 18 depicts the braced ceiling with compression posts at regular intervals. Due to
291
the constraint imposed by the compression posts, the entire system (slab, grid, and panels)
292
generally moved together with equal accelerations, as depicted in case 1. However, during
293
downward slab acceleration of more than 1g, the grid system moved with the slab (assuming
294
the compression posts are rigid) at the compression post locations while the panels were
295
limited to 1g downward acceleration. As a result, the slab and grid accelerations differed
296
from the panel accelerations, causing a gap to form between the ceiling grid and panels.
297
Once the gap formed, the ceiling panels were no longer constrained by the horizontal forces
298
of the grid system, and hence the panels tended to “pop out” of the grid. Furthermore, the
299
ceiling panels impacted the grid system when they fell, weakening the grid members. Gilani
300
et al. (2010) observed this type of damage in ceiling component tests, but concluded that the
301
damage was atypical of that observed in the field due to large vertical accelerations
302
experienced at the roof. This study confirms that the vertical acceleration and associated
13
303
damage pattern can be observed in a realistic building system subjected to significant vertical
304
input shaking.
305
Ceiling Perimeter Attachment Damage
306
Figure 19 shows minor damage observed at the unattached joints between grid members
307
and wall molding in the first 3D-Northridge-Rinaldi simulation. The mechanism is
308
interpreted as follows: when the grid member moved away from the wall, the grid member
309
lost contact with the wall molding (Fig. 20(a)). Since the middle slot was large relative to the
310
screw dimensions, the grid member settled slightly due to a combination of vertical
311
movement of the grid member, rotation of the seismic clip over its attachment point, and
312
popping out of the middle screw (Fig. 20(b)). As the settled grid member moved back toward
313
the wall, it hit the wall molding to cause the observed damage (Fig. 20(c)). This damage
314
could perhaps be avoided by increasing the width, and hence seat length, of the wall molding.
315
Note that ASTM E580/E580M-11ae1 (ASTM, 2011) permits the use of either 22-mm (7/8-
316
in) or 51-mm (2-in) wall molding to support seismic clips, and therefore, the tested design
317
met code requirements.
318
Fragility Methodology
319
Due to the large numbers of applied motions, a fragility curve methodology was used to
320
interpret and extend the test results to assess the seismic vulnerability of the CPP systems.
321
The experimental results were used to estimate the seismic demands, or engineering demand
322
parameter (EDPs), on CPP systems. Seismic fragility curves are conditional probability
323
statements about the vulnerability of a system under the seismic loading. Vulnerabilities are
324
generally expressed in terms of damage states that are physically meaningful in terms of
325
repair (cost and/or time) and the system functionality, and the fragility statement shows the
326
probability that the seismic demand exceeds a threshold capacity associated with the damage
14
327
state. The conditioning parameter of these probabilistic statements is often a single seismic
328
intensity measure (IM) (e.g. horizontal peak floor acceleration or PFA, as used in this study).
329
Seismic fragility curves can be represented by a lognormal cumulative distribution function
330
(Nielson and DesRoches, 2007) as:

ln( Sd / Sc )
P  EDP  DS | IM    
  2  2
C
d IM

331
332




(1)
333
where Sd is the median seismic demand estimate as a function of IM, Sc is the median
334
estimate of the damage state capacity, βd|IM is the logarithmic standard deviation of the
335
demand estimate, βc is the dispersion of the damage state capacity, and Φ[·] is the standard
336
normal cumulative distribution function.
337
The fragility study reported here highlights the relation between seismic performance,
338
wall molding width, and clearance (seismic gap) of the grid members from the wall molding.
339
To do so, the relative ceiling-partition displacements demands were conditioned on the
340
experimentally observed PFA of the associated floor level. A regression analysis of this data
341
was used to estimate the parameters Sd and βd|IM of the probabilistic seismic demand models
342
according to (Cornell et. al., 2002):
343
(2)
344
(3)
345
where di is the peak demand at the ith floor. In developing the fragility curves, all test data
346
obtained from the unbraced ceiling was considered, while only the first five applied motions
347
for the braced ceiling were included.
15
348
The maximum observed forward or reversing relative displacement between the ceiling
349
and partition was evaluated with respect to PFA for each horizontal direction. The
350
displacement versus PFA trends are shown in Fig. 21 for both unbraced and braced ceilings
351
on a log-log scale along with regression lines. Figure 21(a) compares all observed data and
352
fitted curves for both ceilings, while Fig. 21(b) compares All Data and Limited Data
353
(undamaged only) and associated curves for the braced ceiling. Table 2 presents regression
354
parameters a and b along with the dispersion βd|IM .
355
As mentioned previously, the grid members were installed with a clear space of 19 mm
356
(3/4 in.) from the partition walls on the floating side, which left a travel distance of only 3
357
mm (1/8 in) before the grid members unseat from the wall angle. To identify the optimum
358
clearance of grid members from the partition walls, three different limit states were defined.
359
These limit states were classified according to their clearance [19 mm (3/4 in), 16 mm (5/8
360
in), and 13 mm (1/2 in)] between the grid members and partition walls, where 19 mm (3/4 in)
361
represents the tested condition. Each limit state was defined by the median displacement
362
for pounding (equal to the wall clearance) or grid unseating (computed as 22 mm or
363
7/8 in minus the clearance). A constant value of 0.4 was assigned to βC (Table 3), as
364
frequently used for nonstructural components (FEMA, 2012).
365
The fragility curves for the limit states defined above were obtained from Equation (1)
366
using the data of Fig. 21 as demands and the limit state displacements as capacities, and are
367
plotted in Fig. 22 for both the unbraced and braced ceiling configurations. Figure 22(a)
368
implies that the unbraced ceiling is slightly more vulnerable than the braced ceiling.
369
However, Fig. 22(b) shows that the braced ceiling based on Limited Data was more fragile
370
than the braced ceiling based on All Data. Therefore, the differences in vulnerability of the
16
371
braced ceiling may have been influenced by the effect of ceiling repair in the All Data case or
372
insufficient data points in the Limited Data case.
373
The median and logarithmic standard deviation (dispersion) of each fragility curve from
374
Fig. 22 is tabulated in Table 4. The results obtained for the tested condition – a 19 mm (3/4
375
in) clearance between the grid ends and the partitions and a 3 mm (1/8 in) seat length – were
376
considered as the reference point. The fragility data suggests that for a constant length wall
377
molding, reducing the clearance and simultaneously increasing the seat length reduces the
378
vulnerability of the ceiling. Since unseating has a greater probability of occurrence than
379
pounding, increasing the seat length from 3 mm (1/8 in) to 6 mm (1/4 in) and 10 mm (3/8 in)
380
increases the median PFA of the unseating damage state by 77% and 149%, respectively, in
381
the unbraced ceiling and by 93% and 185%, respectively, in the braced ceiling (All Data).
382
Increasing the seat length at the expense of clearance increases the probability of pounding.
383
However, increasing the overall width of the wall molding reduces the probability of
384
unseating without any increase in the probability of pounding. As an example, use of 32 mm
385
(1.25 in) wide wall molding with 19 mm (3/4 in) clearance and 13 mm (1/2 in) seat length
386
achieves the lowest probability of damage for both pounding and unseating (Fig. 22).
387
Seismic Response of Sprinkler Piping
388
Using a similar approach to that reported for the ceiling system, the accelerations
389
recorded on the piping systems and the amplification factors of the piping acceleration with
390
respect to column accelerations were evaluated and compared to the code component
391
amplification factors. Table 5 tabulates acceleration amplification factors for sensors
392
mounted on the main runs, branch lines, and sprinkler heads relative to sensors mounted on
393
the columns, with slab to column amplification included for reference. The median horizontal
394
amplification observed during the experiment was 2.6 and 2.28 for the 5th and 6th floor main
17
395
runs relative to the columns, respectively, which is comparable to the code component
396
amplification factor (ap = 2.5). However, these amplification factors were increased to 5.7
397
and 4.9 on the branch lines and 6.4 and 5.5 on sprinkler heads of the 5th and 6th floor piping
398
system, respectively. The responses of the branch lines, armovers, and drops were further
399
amplified since the acceleration of the main runs (already amplified relative to the structural
400
acceleration by a factor of 2.5) served as input excitation to the branch lines. In general, the
401
largest horizontal acceleration was observed in the main run perpendicular direction at the
402
middle compared to other main run locations. The transition from loose to tension wire
403
restrainers induced large pulse accelerations in the direction of restrainers.
404
The median vertical acceleration amplification relative to the column sensors was greater
405
than 4 for the main runs on both floors, 6.3 for the 5th floor branch line, and 4.9 for the 6th
406
floor branch line. Thus, the observed vertical amplification factors were more than twice ap.
407
On the other hand, the median vertical amplification of the slab relative to the column sensors
408
was 3 and 3.9 for the 5th and roof slabs, respectively, which implies that much of the piping
409
acceleration amplification relative to the columns can be attributed to structural slab
410
flexibility. Thus, design of the piping supporting elements and anchorages based on a vertical
411
component amplification factor ap = 2.5 may be inadequate.
412
Damage Near Sprinkler Heads
413
Wherever rigid drop pipes were used, the ceiling panels sustained damage from pounding
414
of the sprinkler heads regardless of whether the oversized gap configuration, which
415
conformed to code requirements (ASTM, 2011), or the no gap configuration was used.
416
During the XY-Tohoku-Iwanuma excitation/fixed-base configuration with PFA = 1.12g, up
417
to 203 mm (8 in) of material was knocked out of the ceiling panel (Fig. 23(a)-(b)), which is
418
much larger than the 51 mm (2 in) gap required by code. Tearing of ceiling panels due to the
18
419
piping interaction is a function of maximum piping movement as ceiling panels are
420
composed from very weak fibers. Therefore, this type of damage might not be sensitive to the
421
duration of a motion. However, 203 mm (8 in) tearing of ceiling panels was observed during
422
XY-Tohoku-Iwanuma excitation, which is a strong motion with a very long duration. On the
423
other hand, no damage was observed around the flexible hose fittings that were mounted at
424
the end of Branch Line 3 (Fig. 23(c)).
425
Similar to the ceiling-partition relative displacement, a fragility procedure was used here
426
to study the probability of ceiling-sprinkler head interaction. The peak ceiling-piping relative
427
displacement was determined as the peak relative displacement between the piping and
428
ceiling systems, both of which were recorded by displacement transducers. A regression
429
analysis of the relative displacement with respect to PFA in the direction perpendicular to the
430
main run was performed using Equation (1). Figure 24 plots the ceiling-piping (sprinkler
431
head) displacement demand on a log-log scale along with the regression curve. As shown in
432
this figure, higher heteroscedasticity was observed between the demands and IMs at PFAs
433
smaller than 0.1g. While linear regression was applied to the corresponding data to estimate
434
the median response, alternative approaches such as weighted linear least squares regression
435
can be used to accommodate such a non-constant variability (heteroscedasticity) (Ang and
436
Tang, 1975).
437
Several limit states have been defined herein to interpret the seismic response of ceiling-
438
piping interaction during the experiment (Table 6). These limit states were defined by
439
(median clearance of ceiling panel and sprinkler head) and βC (logarithmic standard deviation
440
of this clearance). Three values of 3 mm (0.1 in), 13 mm (0.5 in), and 25 mm (1 in) were
441
considered as the clearance limit states (
442
"no gap"), 25 mm (1 in), and 51 mm (2 in) diameter of ceiling panel holes around the
), which correspond to the 6 mm (0.2 in) (named
19
443
sprinkler heads. A value 0.4 was assigned to βC as before. The fragility curves for ceiling-
444
piping interaction computed by Equation (1) are shown in Fig. 25. Table 7 lists the median
445
and dispersion of the fragility curve for each limit state.
446
The fragility curves show that during this experiment, even a 25 mm (2 in) oversized ring
447
around the sprinkler heads was not sufficient to prevent ceiling-piping interaction for floor
448
accelerations exceeding 0.65g. The experiment showed that use of a flexible hose drop was
449
effective in eliminating this type of damage. Decreasing the spacing of lateral sway braces,
450
currently spaced no more than 12 m (40 ft) (NFPA, 2011), might reduce the ceiling-piping
451
interaction. Decreasing the spacing of braces along the main runs can also limit the piping
452
deflection due to bending and subsequently reduce the ceiling-sprinkler head interaction.
453
Permanent Rotation of Armover Drops
454
A vulnerability of armover drops compared to straight drops was observed in these
455
experiments. During the first 3D-Northridge-Rinaldi excitation (input vertical peak table
456
acceleration = 1.2g and resultant peak slab vibration = 6.4g), the entire 6th floor Branch Line
457
1 with three armover drop pipes twisted around its connection point to the main run (Fig.
458
26(a)). During vertical acceleration, a vertical inertia force was generated proportional to the
459
mass of the armover drop. The twisting moment around the branch line was the summation of
460
the torque generated at each drop (Fig. 26(b)). The current code (NFPA, 2011) permits the
461
connections along this branch line to be designed without torsional resistance when the
462
cumulative horizontal length of the unsupported armover < 610 mm (24 in), which was true
463
for the experimental test setup. However, the torsional resistance of the threaded joints was
464
not sufficient to resist the cumulative torsional demand generated in the large vertical
465
excitation, and permanent twisting of the branch line was observed.
20
466
Next, the twisting moment of the threaded connection of the armover branch to the main
467
run in the 6th floor piping system was estimated by a simple calculation. Based on videos
468
recorded during the experiment, the branch line was observed to twist at ~ 9.5-10.5 sec into
469
the excitation. The rotational acceleration of each armover drop was estimated from the
470
relative acceleration between the corresponding sprinkler head and the closest point on the
471
branch line, which happened to be the main run (Fig. 12(a)). The relative acceleration history
472
of each armover is plotted in Fig. 27. The instantaneous peak of the relative acceleration
473
summed over the three armovers was 6.51g at 10.37 sec. The corresponding rotational
474
acceleration  = 210 rad/s2 was used to compute the twisting moment (T) that triggered the
475
rotation of branch line around its connection to the main run:
T  I 
476
(4)
477
where I is the effective rotational inertia of armover, drop pipes, and sprinkler head masses
478
about the branch line, assuming a lumped mass approximation. The average twisting moment
479
resistance per unit (circumferential) length is defined as follows:
sT 
480
T
  D0
(5)
481
where D0 is the outside diameter of branch line pipe. The moment was estimated as 0.05 kN-
482
m (0.4 kip-in), leading to sT = 4 kN-m/m (0.08 kip-in/in). If sT is assumed to be independent
483
of the threaded pipe diameter, Equation (5) approximates the maximum torsional resistance
484
of the piping.
485
Seismic Response of Partitions
486
Past component-level testing suggests that drift-related damage to partition walls occurs
487
at median story drifts of 0.3% (Retamales et al., 2012). Maximum drifts observed during the
488
experiment were 0.78% on the 4th story and 0.62% on 5th story, which occurred during the
21
489
3D-Northridge-Rinaldi excitation on the fixed-base configuration. Drift-related damage to the
490
partition walls was not noticeable.
491
On the other hand, atypical damage states were observed that were attributed to the
492
relative vertical acceleration between the floors. After the first 3D-Northridge-Rinaldi
493
excitation, large diagonal and vertical cracks appeared on the gypsum wallboards of full
494
height partitions (Fig. 28(a)). Also, at a few locations on the 5th story partitions, which
495
utilized slip track connections, the top of the studs were observed to move laterally or “pop
496
out” permanently from their constrained position within the top tracks (Fig. 28(b)). This
497
damage state, which – to our knowledge – has not been observed prior to this experiment,
498
was due to the large differential movement between the top track (attached to the roof slab)
499
and the studs (mounted on the 5th floor slab).
500
Damage to bulkhead and partial height self-standing partitions exceeded that of the full
501
height partitions. During the first 3D-Northridge-Rinaldi excitation, the studs of bulkhead
502
partitions buckled, again due to the differential acceleration between the 5th floor and roof
503
slabs (Fig. 29(a)). Bulkhead partitions are not common in practice, but were designed to meet
504
test constraints. Also, during the final motion of the test program, the top portion of the door
505
opening separated from the adjoining wall, and the bottom connection of the track to the slab
506
failed (Fig. 29(b)). These failures were likely due to the cumulative pounding against the
507
walls of loose objects from inside the rooms. Table 8 lists the out-of-plane horizontal
508
acceleration amplification of the full height partition walls compared to the slabs that they
509
were mounted on or hung from, and amplification of partial height self-standing partitions
510
relative to the slabs below. The data for full height partitions was categorized into full
511
connections (Full) and slip track connections (Slip) with respect to slab below (Below) or
512
slab above (Above). Partial height partitions were categorized based on their institutional and
22
513
commercial details. Amplification factors were computed for each relevant ground excitation
514
along with statistics (maximum, minimum and median).
515
The median amplification of the partition out-of-plane acceleration observed during the
516
experiment was 2.7 and 2.8 for partitions with full connections, and 3.8 and 3 for partitions
517
with slip track connections relative to the below and above slab, respectively. These values
518
are slightly higher than the code component amplification factor (ap = 2.5), especially for slip
519
track partitions. As mentioned previously, the gypsum board and steel studs of slip track
520
partitions were not attached to the top tracks, which caused their out-of-plane acceleration to
521
be higher than those with full connection details. These results also suggest that ceiling tests
522
need to incorporate the flexible partition wall boundaries to represent realistic ceiling
523
behavior. Experimental ceiling specimens that are attached to rigid boundaries cannot
524
account for the influence of flexible partition walls on overall ceiling system performance.
525
The median acceleration amplification factor of the partial height partitions was > 5,
526
which exceeded that of full height partitions. This higher amplification could have been
527
caused by pounding of the unrestrained objects against the walls. The observed amplification
528
factors suggest that current code factors for attachment and anchorage design forces for
529
partial height partitions should be reconsidered.
530
Summary and Conclusions
531
532
The major findings of this experiment are summarized below:

In the horizontal direction, the median CPP component acceleration amplification
533
factors (peak component acceleration relative to corresponding floor acceleration)
534
observed during the experiments were closely correlated to the recommended code
535
factor ap.
536

The component amplification factors in the vertical direction observed during the
23
537
experiment were much greater than the recommended code factor, which can
538
primarily be attributed to flexibility of the floor slabs.
539

Use of lateral bracing with compression posts may not improve the seismic response
540
of the ceiling when subjected to strong vertical excitation. Further research is needed
541
to confirm the generality of this observation.
542

Seismic clips were used around the ceiling perimeter in conjunction with reduced
543
width 22 mm (7/8 in) wall molding and a 19 mm (3/4 in) grid-partition clearance.
544
This detail was found to be vulnerable to unseating followed by pounding due to load
545
reversal. Although the seismic clip detail is expected to be an improvement over
546
standard detailing, increasing the width of wall molding and/or or reducing the grid-
547
partition clearance may alleviate the behavior observed in the experiment.
548

Due to the twisting moment generated from the armover drops, branch line pipes with
549
several unsupported armover drops can twist around the branch line threaded
550
connection point. A simple approach was proposed to estimate the maximum
551
torsional resistance as a function of threaded pipe diameter.
552

The oversized gap configuration with 51 mm (2 in) ring was not effective to prevent
553
damage to ceiling panels resulting from sprinkler head pounding; however, the use of
554
flexible hose drops substantially reduced the piping-ceiling interaction.
555

These observed response mechanisms may be sensitive to specific circumstances of
556
the experiments, such as building configuration (e.g. slab vibration characteristics)
557
and acceleration demands (e.g. large vertical acceleration relative to horizontal
558
acceleration).
559
Acknowledgements
24
560
This material is based upon work supported by the National Science Foundation under
561
Grant No. CMMI-0721399. This GC project to study the seismic response of nonstructural
562
systems is under the direction of M. Maragakis from the University of Nevada, Reno and Co-
563
PIs: T. Hutchinson (UCSD), A. Filiatrault (UB), S. French (G. Tech), and B. Reitherman
564
(CUREE). Any opinions, findings, conclusions or recommendations expressed in this
565
document are those of the authors and do not necessarily reflect the views of the sponsors.
566
The input provided by the NEES Nonstructural Project Practice Committee, composed of W.
567
Holmes (Chair), D. Allen, D. Alvarez, and R. Fleming; by the Advisory Board, composed of
568
R. Bachman (Chair), S. Eder, R. Kirchner, E. Miranda, W. Petak, S. Rose and C. Tokas, has
569
been critical for the completion of this research. The authors recognize and thank the
570
following companies for providing product donations and technical support: USG Building
571
systems, Victaulic, Tolco, Hilti, Allan Automatic Sprinkler and CEMCO steel.
572
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573
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576
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604
605
National Fire Protection Association (NFPA), (2011). NFPA 13: Standard for the Installation of
Sprinkler Systems." National Fire Protection Association, 2010 Edition, Quincy, MA.
Nielson, G. B. and DesRoches, R. (2007). “Analytical Seismic Fragility Curves for Typical Bridges in
the Central and Southeastern United States”, Earthquake Spectra, 23(3), 615-633.
606
Retamales, R., Davies, R., Mosqueda, G., Filiatrault, A. (2012). “Experimental Seismic Fragility of
607
Cold-Formed Steel Framed Gypsum Partition Walls”, Journal of Structural Engineering, ASCE,
608
139, Special Issue: NEES 2: Advances in Earthquake Engineering, 1285-1293.
26
609
Ryan, K. L., Soroushian S., Maragakis, E. M., Sato, E., Sasaki, T., Okazaki, T. (2013a). “Seismic
610
simulation of integrated nonstructural systems at E-Defense, Part 1: Influence of 3D structural
611
response and base isolation”, Under Review in J. Struct. Eng.
612
Ryan K, Sato E, Sasaki T, Okazaki T, Guzman J, Dao N, Soroushian S, Coria C (2013b). "Full Scale
613
5-story Building with Triple Pendulum Bearings at E-Defense", Network for Earthquake
614
Engineering Simulation (database), Dataset, DOI:10.4231/D3X34MR7R.
615
Ryan K, Sato E, Sasaki T, Okazaki T, Guzman J, Dao N, Soroushian S, Coria C (2013c). "Full Scale
616
5-story Building with LRB/CLB Isolation System at E-Defense", Network for Earthquake
617
Engineering Simulation (database), Dataset, DOI:10.4231/D3SB3WZ43.
618
Ryan K, Sato E, Sasaki T, Okazaki T, Guzman J, Dao N, Soroushian S, Coria C (2013d). "Full Scale
619
5-story Building in Fixed-Base Condition at E-Defense", Network for Earthquake Engineering
620
Simulation (database), Dataset, DOI:10.4231/D3NP1WJ3P.
621
Taghavi, S., and Miranda, E. (2003). “Response Assessment of Nonstructural Building Elements”,
622
PEER Rep. 2003/05, Pacific Earthquake Eangineering Research Center (PEER), Univ. of
623
California, Berkeley, CA.
624
Tian, Y., Filiatrault, A., Mosqueda, G. (2013). “Experimental Seismic Study of Pressurized Fire
625
Sprinkler Piping Subsystems”, Technical Report MCEER-13-0001, Multidisciplinary Center for
626
Earthquake Engineering Research, State University of New York at Buffalo, Buffalo, NY, USA,
627
2013.
628
Whittaker, A.S. and Soong, T.T., (2003). “An Overview of Nonstructural Component Research at
629
Three U.S. Earthquake Engineering Research Centers”, in Proceedings of Seminar on Seismic
630
Design, Performance, and Retrofit of Nonstructural Components in Critical Facilities, Applied
631
Technology Council, ATC-29-2, pp. 271-280, Redwood City, California.
632
Zaghi, A. E., Maragakis, E. M., Itani, A., and Goodwin, E. (2012). “Experimental and Analytical
633
Studies of Hospital Piping Subassemblies Subjected to Seismic Loading”, Earthquake Spectra,
634
Earthquake Engineering Research Institute. 28(1):367-384.
635
27
636
637
Tables
Table 1. Maximum and Minimum Ceiling to Column or Slab to Column Acceleration
Amplification
5th Floor
Motion
Number
1
2
3
4
5
...
41
Max=*
Min=*
Median*
Grid
Column
Slab
Column
6th Floor
Panel
Column
Grid
Column
Slab
Column
Panel
Column
XY
Z
Z
XY
Z
XY
Z
Z
XY
Z
1.02
1.25
2.55
3.27
3.00
...
1.37
7.19
1.00
2.57
0.94
1.71
5.14
4.82
2.23
...
4.41
6.25
2.09
4.24
1.00
1.13
4.51
4.26
2.45
...
1.25
4.95
1.50
2.97
1.00
1.16
3.25
3.23
3.72
...
NA
6.57
0.97
2.49
0.93
1.01
5.61
3.72
2.26
...
NA
5.61
1.94
3.58
1.03
1.15
5.75
4.52
4.27
...
2.29
5.75
1.03
2.83
1.31
1.43
5.27
4.67
2.03
...
6.44
5.27
4.67
4.97
1.08
1.69
6.46
4.88
2.14
...
1.35
6.46
4.88
5.67
1.01
1.08
7.70
5.88
16.78
...
NA
7.70
1.01
3.48
1.83
2.89
12.8
10.9
2.54
...
NA
12.8
10.9
11.9
638
639
640
641
Table 2. Demand Parameter
Estimations for Ceiling Displacements
System
5th Floor Ceiling
a
4.83
b
1.20
βd|PFA
0.53
6th Floor Ceiling-All
6th Floor Ceiling-Limited
4.32
7.37
1.05
1.40
0.72
0.78
Table 3. Limit States of Grid Perimeter Pounding or
Unseating (mm)
Ceiling System
(5th and 6th Floor)
19 mm (3/4 in) Clearance
16 mm (5/8 in) Clearance
13 mm (1/2in) Clearance
Pounding
Unseating
θdisp
βc
θdisp
βc
19
16
13
0.4
0.4
0.4
3
6
10
0.4
0.4
0.4
642
643
644
645
Table 4. Medians and Dispersions for 3 Different Ceiling Perimeter on the Floating Side
System
3/4" Clearance
Median
PFA(g) Dispersion
5th Floor Ceiling-Pounding
5th Floor Ceiling-Unseating
6th Floor Ceiling- Pounding
6th Floor Ceiling-Unseating
3.16
0.71
4.00
0.74
0.66
0.66
0.82
0.82
6th Floor Ceiling- Pounding
6th Floor Ceiling-Unseating
1.95
0.54
0.88
0.88
5/8" Clearance
Median
Difference
PFA(g)
Dispersion
(%)
All Data Points
2.71
0.66
-14
1.26
0.66
77
3.43
0.82
-14
1.43
0.82
93
Limited Data Points
1.71
0.88
-12
0.89
0.88
65
646
28
Median
PFA(g)
1/2" Clearance
Difference
Dispersion
(%)
2.25
1.77
2.78
2.11
0.66
0.66
0.82
0.82
-29
149
-31
185
1.46
1.19
0.88
0.88
-25
120
Table 5. Maximum Piping/Floor Acceleration Amplification
5th Floor Piping
Motion
Number
1
2
3
4
5
...
41
Max=*
Min=*
Median=*
647
648
Main
Column
Slab
Column
6th Floor Piping
Branch
Column
Head
Column
Main
Column
Slab
Column
Branch
Column
Head
Column
XY
Z
Z
XY
Z
XY
XY
Z
Z
XY
Z
XY
1.11
1.78
4.37
4.14
3.21
...
1.98
7.74
0.97
2.57
1.66
2.17
5.78
6.12
2.45
...
15.9
6.91
2.45
4.44
1.00
1.13
4.51
4.26
2.45
...
1.25
4.95
1.50
2.97
1.37
2.21
14.8
12.1
7.87
...
3.43
14.9
1.37
5.74
1.19
1.49
11.48
10.20
2.26
...
38.12
11.85
2.26
6.31
1.38
3.25
15.10
16.17
5.24
...
2.76
16.59
1.38
6.38
1.08
1.72
3.05
3.80
3.40
...
2.28
4.71
1.08
2.28
2.11
2.36
8.02
11.9
3.79
...
23.5
12.0
2.50
4.16
1.08
1.69
6.46
4.88
2.14
...
1.35
6.46
1.55
3.88
1.19
1.98
11.17
10.50
8.30
...
3.85
14.35
1.19
4.91
1.39
2.00
11.9
9.10
2.22
...
54.7
11.9
2.22
4.85
1.43
2.46
19.48
12.22
7.37
...
4.58
19.48
1.44
5.45
* Damaged Pipes (Highlighted numbers, first branch line data) were not included.
* Only 3D motions were considered for z direction calculation
Table 6. Limit States of Ceiling-Piping Interaction
Clearance
Ceiling - Sprinkler Head
3 mm (1/10 in) Clearance
13 mm (1/2 in) Clearance
25 mm (1 in) Clearance
θint
βc
3
13
25
0.4
0.4
0.4
649
650
651
Table 7. Medians and Dispersions of Ceiling-Piping Interaction
3 mm (1/10 in) Clearance
Median
PFA(g)
Dispersion
0.07
0.77
1.10
1.34
13 mm (1/2 in) Clearance
Median
PFA(g)
Dispersion
0.33
0.77
3.55
1.34
25 mm (1 in) Clearance
Median
PFA(g) Dispersion
0.64
0.77
N/A*
1.34
* Estimated median values are much larger than can be appropriately extrapolated from regression analyses.
652
653
654
Table 8. Maximum Partition/Floor Acceleration Amplification
Partial Height Self
Standing Partitions
Full Height Partitions
Motion
Number
1
...
34
35
...
41
Max=*
Min=*
Median=*
Full
Below
Full
Above
Slip
Below
Slip
Above
Commercial
Institutional
1.51
...
5.08
3.14
...
1.40
6.35
1.16
2.62
1.15
...
4.91
2.29
...
1.16
7.43
1.00
2.79
1.87
...
8.40
2.24
...
1.24
8.40
1.17
3.77
1.31
...
7.63
1.93
...
1.11
8.09
0.92
2.95
3.34
...
7.69
4.75
...
3.58
13.69
1.40
5.12
3.66
...
7.00
3.41
...
3.35
15.54
2.75
5.75
* Damaged partitions (highlighted numbers) were not considered.
655
29
656
657
List of Figures
658
Figure 1. Overall view of ceiling system layout
659
Figure 2. Connection of runners/cross tees and wall molding: (a) attached detail, and (b)
660
unattached
661
Figure 3. Connection of seismic bracing to the ceiling grid
662
Figure 4. Overall plan view of piping system
663
Figure 5. Sprinkler heads and drops: (a) flexible drop, (b) 50 mm (2 in) oversized gap
664
configuration, (c) no gap configuration
665
Figure 6. Bracing for piping system: (a) lateral and longitudinal brace near riser, and (b)
666
diagonal splay wires and pipe hanger at the end of each branch line
667
Figure 7. Overall partition plan view; black labels: 4th story, gray labels: 5th story
668
Figure 8. Typical institutional (a) T and (b) corner; and commercial (c) T and (d) corner
669
connections
670
Figure 9. Typical institutional connection details for full connection: (a) top, (b) bottom; and
671
slip track connection: (c) top, (d) bottom
672
Figure 10. (a) Partial height self-standing partitions, (b) bulkhead partitions
673
Figure 11. (a) Summary of accelerometers on ceiling system (plan view); accelerometer on:
674
(b) ceiling grid, (c) ceiling panel, (d) compression post; (e) displacement transducer at ceiling
675
perimeter
676
Figure 12. Instrumentation view: (a) plan view of accelerometers placed on piping system;
677
accelerometer on: (b) sprinkler head, (c) branch line; (d) displacement transducer on main run
678
Figure 13. Accelerometers installed on (a) full height partitions, (b) partial height partitions
679
Figure 14. Vertical acceleration histories recorded at slab, ceiling, and column locations in
680
roof SE slab during 3D-Superstition Hills-Westmorland (TP configuration)
30
681
Figure 15. Condition of (a) 6th floor braced ceiling and (b) 5th floor unbraced ceiling after
682
3D-Northridge-Rinaldi (TP configuration)
683
Figure 16. Vertical acceleration in panel (C9) versus grid (C4) in (a) 5th floor unbraced and
684
(b) 6th floor braced ceiling due to 3D-Superstition Hills-Westmorland (TP configuration)
685
Figure 17. Vertical dynamics of unbraced ceiling
686
Figure 18. Vertical dynamics of braced ceiling
687
Figure 19. Ceiling perimeter attachment failure after 3D-Northridge-Rinaldi (TP
688
configuration)
689
Figure 20. Grid-wall molding interaction mechanism
690
Figure 21. Ceiling-partition relative displacement seismic demand: (a) All Data for braced
691
and unbraced (b) All Data and Limited Data for braced
692
Figure 22. Ceiling perimeter fragility curves on the floating side: (a) unbraced and braced -
693
All Data, (b) braced - All and Limited Data
694
Figure 23. Comparison of pounding damage on 6th floor after XY-Tohoku-Iwanuma
695
excitation: (a) conventional sprinkler head in no gap configuration and (b) oversized gap
696
configuration; (c) flexible hose sprinkler head
697
Figure 24. Ceiling-piping relative displacement perpendicular to the main run
698
Figure 25. Ceiling-piping interaction fragility curves in direction perpendicular to main runs
699
Figure 26. (a) Armover permanent rotation following 3D-Northridge-Rinaldi (TP
700
configuration), and (b) torsional demand on armover branch line
701
Figure 27. Sprinkler head-branch line relative acceleration at armover branch line in 6th floor
702
during 3D-Northridge-Rinaldi excitation (TP configuration)
703
Figure 28. (a) Typical cracks on full height partitions, and (b) lateral movement of studs
704
from the top tracks on slip track partitions
31
705
Figure 29. (a) Stud buckling of bulkhead partitions, (b) damage to partial height self-
706
standing partitions
707
32
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Copyright Agreement
Click here to download Copyright Agreement: Copyright Transfer - Paper 2.pdf
Sizing worksheet (.xls)
Click here to download Sizing worksheet (.xls): Sizing Worksheet_Paper 2.xls
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STENG-3260
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Keri Ryan
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Response to Reviewers Comments
Click here to download Response to Reviewers Comments: Response-to-reviewer-comments-Paper 2.pdf
Seismic Simulation of Integrated Nonstructural Systems at E-Defense, Part 2:
Evaluation of Nonstructural Damage and Fragilities
Reviewer #2:
1) One of the unique features of this study is that it is a system-level investigation. However, in
the presentation and analysis of the nonstructural data, the results of the three tested
configurations (fixed base, isolated with LRB+cross linear bearings, and isolated with TFP
bearings) are all thrown into one bin. Doesn't the type of isolation system, or lack thereof,
have an effect? This needs to be addressed.
One purpose of Part 1 of the companion papers was to describe the structural response
observed in the different configurations and its relation to the nonstructural response.
Certainly, horizontal accelerations are generally different in isolated buildings compared to
fixed-base buildings. However, the variation of horizontal floor accelerations observed in the
study was minimized due to the need to apply different scale factors to the motions for the
isolation and fixed-base configurations. More importantly, Part 1 demonstrated that the
response of the CPP system (at the amplitudes observed in the experiment) were more
closely tied to the vertical slab acceleration than the horizontal floor acceleration, and that
slab vibration was insensitive to the presence of an isolation system. A new section titled
“Influence of Isolation System on Vertical Amplification Factors” with an expanded
discussion of the propagation of vertical motion through the structure has been added to
strengthen this argument.
As a result, the presence of the isolation system did not have a noteworthy effect on the CPP
system response in this experiment. We have correlated the CPP damage to the observed
horizontal and vertical accelerations recorded at the floor level, which is the best indicator of
damage to those components. To clarify for the reader, we have expanded the discussion
related to this issue: “The damage mechanisms and the extent of damage were very similar
for the three system configurations because, as discussed in Ryan et al. (2013), similar peak
demands were observed in the three system configurations and vertical slab vibration was
insensitive to the presence of an isolation system. Thus, the damage observations are
presented and discussed without further mention of the system configuration during which
they were observed, but rather interpreted relative to the horizontal floor acceleration and
vertical slab acceleration to which they are closely correlated.” (lines 193 to 199).
2) As for part 1, the title of paper needs to be more specific. What types of nonstructural
systems are investigated?
A correction has been made. The title of paper was changed to:
“Seismic Simulation of an Integrated Ceiling-Partition Wall-Piping System at E-Defense, Part 2:
Evaluation of Nonstructural Damage and Fragilities”
3) The first paragraph of the Introduction needs references.
The following references have been added to the text:
FEMA E-74, (2011). “Reducing the Risks of Nonstructural Earthquake Damage: A Practical Guide”,
Federal Emergency Management Agency, Washington, D.C.
Miranda, E., (2003). “Building Specific Loss Estimation for Performance Based Design”, 2003
Pacific Conference on Earthquake Engineering, Christchurch, New Zealand.
Taghavi, S., and Miranda, E. (2003). “Response Assessment of Nonstructural Building Elements”,
PEER Rep. 2003/05, Pacific Earthquake Eangineering Research Center (PEER), Univ. of California,
Berkeley, CA.
Whittaker, A.S. and Soong, T.T., (2003). “An Overview of Nonstructural Component Research at
Three U.S. Earthquake Engineering Research Centers”, in Proceedings of Seminar on Seismic
Design, Performance, and Retrofit of Nonstructural Components in Critical Facilities, Applied
Technology Council, ATC-29-2, pp. 271-280, Redwood City, California.
4) There is a logical gap between the last two sentences of the first paragraph in the
Introduction. Something needs to be said about the value of structural vs nonstructural
components.
A correction has been made. The following sentence has been added:
“Also, nonstructural systems almost always represent the major portion of the total investment in
buildings (Whittaker and Soong, 2003).”
5) Line 165. Many types of nonstructural components have high natural frequencies. Is a 25 Hz
cutoff frequency appropriate for the Butterworth filter?
The reviewer raises a valid question. To clarify, the cutoff frequency of 25 Hz was used in
the previous publications of this experiment by the authors, and therefore, the same cutoff
frequency was chosen in this study to provide consistency of presented results. While 25 Hz
may be on the low side for nonstructural components, it is larger than the cutoff frequency
for rigid components (16.7 Hz, according to ICC-AC156), and thus encompasses the
frequency range of the interest. Knowing that the use of larger cutoff frequencies may result
in signals with high frequency noise, the authors decided to stay with cutoff frequency of 25
Hz to process the data.
ICC Evaluation Service (2010). AC 156 Acceptance Criteria for Seismic Certification by Shake Table
Testing of Nonstrucutral Components, ICC Evaluation Service.
6) Paragraph starting on line 223. A time history window of the involved responses, for a
typical case, might aid the discussion.
As suggested by the reviewer, a time history window of SE slab, ceiling, and columns was
added to show the trend discussed in the mentioned paragraph.
7) What are the consequences of falling ceiling panels. Do they pose a credible threat to the
safety of building occupants? Do they break easily? If not, can they be easily popped back
into place, or is repair associated with this type of damage costly? Etc.
Offering some perspective may be useful.
The following statement was added to the text in order to address the reviewer’s concern:
“While the chance of a life threatening injury due to the fallen ceiling panels was low, serious
disruption was observed in the office room. Most of the fallen panels were damaged to the extent that
replacement would be desirable. In addition, fallen ceiling panels weakened the grid system and
increased the likelihood that more extensive grid repair was necessary. However, over the course of
the test program, some of the cross tee sections failed but the main runners always remained intact.”
8) Line 233. For what type of building and site location are the median vertical amplification
factors from this study suitable for design or evaluation? This goes back to the choice of
ground motions used in this study, being criticized as uncharacteristic in the vertical
direction (see review of STENG-3259).
As shown in Figure 5 of Part 1 (STENG-3259), a wide range of vertical shake intensities and
frequencies were included in the test program. In addition, the variation of amplification
factor was found to be independent of the shaking intensity. Therefore, the median values
presented in this study may not correspond to specific types of buildings or site locations.
This is also consistent with the usual interpretation of ap, which is a function of component
type (or their flexibility) rather than site location or building types.
9) Line 308. This sentences does not quite make sense. The relationship between the demand
and the IM is expressed through a functional dependence similar to Equation 2 but with
randomness attached to it.
The correction has been made. Below is the modified sentence:
“Seismic fragility curves can be represented by a lognormal cumulative distribution function (Nielson
and DesRoches, 2007) as:”
10) The numerator in Equation 3 does not look correct. Hopefully this is just a typo.
The correction has been made. It was a typo mistake.
11) Line 394 and Figure 22. The damage described and shown in this case corresponds to the
Iwanuma motion, a strong motion with uncharacteristically very long duration. The motion
resulted in the highest PFA and EDP. Isn't it fair to say that damage like this is atypical?
Duration of motion may not be an important factor on causing damage to ceiling panels.
Experimental studies by Soroushian et al. (2014) and Tian et al. (2013) showed similar
damage in motions with 10 sec. and 48 sec. in durations. As shown by Soroushian et al.
(2014), 8 in. tearing of ceiling panels happened during a 10 sec motion with 1.06g PFA in a
full scale two-story experiment, which is very comparable to the observation of this study.
Note that ceiling panels are composed from very weak fibers and this type of damage is more
function of piping movement. The effect of motion amplitude on the extent of panel tearing
is presented in Figure 23. The following statement is added to the paper to address the
reviewer’s comment:
“Tearing of ceiling panels due to the piping interaction is a function of maximum piping
movement as ceiling panels are composed from very weak fibers. Therefore, this type of
damage might not be sensitive to the duration of a motion. However, 203 mm (8 in) tearing
of ceiling panels was observed during XY-Tohoku-Iwanuma excitation, which is a strong
motion with a very long duration.”
Soroushian, S., Rahmanishamsi, E., Ryu, K. P., Maragakis, E. M., Reinhorn, A.M., “A Comparative Study of
Sub-System and System Level Experiments of Suspension Ceiling Systems”, Tenth U.S. National Conference
on Earthquake Engineering, July 2014, Anchorage, AK.
Tian, Y., Filiatrault, A., Mosqueda, G. (2013) “Experimental Seismic Study of Pressurized Fire Sprinkler Piping
Subsystems,” Technical Report MCEER, State University of New York at Buffalo, NY, MCEER 13-0001.
12) Figure 23 shows that the dispersion of the data in log-log space is nonuniform. The
dispersion becomes huge in the lower range of the independent variable. Ordinary least
squares, which is used in this study to estimate the parameters of the regression model, is
problematic in this case. The reviewer suggests that the authors look up "heteroscedasticity"
in statistics literature for a description of the phenomenon and for possible ways to deal with
it.
The following statement has been added to address the reviewer’s comment:
“ As shown in this figure, higher heteroscedasticity was observed between the demands and IMs at
PFAs smaller than 0.1g. While linear regression was applied to the corresponding data to estimate
the median response, alternative approaches such as weighted linear least squares regression can be
used to accommodate such a non-constant variability (heteroscedasticity) (Ang and Tang, 1975).”
13) Figure 8 in Part 1 (STENG-3259) shows PFAs that exceed 2g for the fixed-base case.
However, in Part 2 the highest PFA appears to be 1.1 g, which is high for an isolated
building, but rather low for a fixed base building. Is the data for higher PFA omitted because
it corresponds to repaired ceiling systems (line 190)? If this is the case, then the upper end of
the IM used in the regression analysis is low and extrapolating to more realistic values of
PFA for fixed base buildings may or may not be applicable. This need to be noted.
Figure 8 in Part 1 (STENG-3259) shows the ratios of PFA/PGA that exceed 2 for the fixedbase case. This ratio is unitless and the (g) provided in the label of the x axis was included by
mistake. Please note that this figure was replaced (now Figure 9) in the revised version of
part 1, and now shows both absolute acceleration and acceleration ratio.
14) Lines 409-411. Unclear.
The sentence referred to was rephrased as follows:
“Three values of 3 mm (0.1 in), 13 mm (0.5 in), and 25mm (1 in) were considered as the
clearance limit states (θint), which correspond to the 6mm (0.2in) (named "no gap"), 25 mm
(1 in), and 51 mm (2 in) diameter of ceiling panel holes around the sprinkler heads.”
Reviewer #3:
1) Was there any correlation between the amplification of vertical acceleration and the number
of stories (i.e., how high up in the structure the accelerations are measured)?
We believe the reviewer is referring to amplification of CPP component acceleration relative
to floor slabs (or other appropriate measure) and not the amplification of column and slab
vibration relative to the ground. The latter was discussed in Part 1 (STENG-3259) (now
Figures 10 and 11). However, the amplification of CPP component response with respect to
height cannot be well addressed since the nonstructural components were only installed on
the top two floors. Also, as shown in new Figure 14 of this paper, vertical acceleration
amplification in the nonstructural components are related to the slab vibration, which may
vary more as a function of slab vibration frequency than building height.
2) The data has been processed as "All data" and "Limited data" based on the damage state. No
specific conclusions were drawn based on the existence of the isolation. Does that mean that
the base condition of the structure does not have any effect on the nonstructural damage?
Can this be due to the order of the 6 day testing without full repair of the damages?
One purpose of Part 1 of the companion papers was to describe the structural response
observed in the different configurations and its relation to the nonstructural response.
Certainly, horizontal accelerations are generally different in isolated buildings compared to
fixed-base buildings. However, the variation of horizontal floor accelerations observed in the
study was minimized due to the need to apply different scale factors to the motions for the
isolation and fixed-base configurations. More importantly, Part 1 demonstrated that the
response of the CPP system (at the amplitudes observed in the experiment) were more
closely tied to the vertical slab acceleration than the horizontal floor acceleration, and that
slab vibration was insensitive to the presence of an isolation system. A new section titled
“Influence of Isolation System on Vertical Amplification Factors” with an expanded
discussion of the propagation of vertical motion through the structure has been added to
strengthen this argument.
As a result, we believe the presence of the isolation system did not have a noteworthy effect
on the CPP system response in this experiment. We have correlated the CPP damage to the
observed horizontal and vertical accelerations recorded at the floor level, which is the best
indicator of damage to those components. To clarify for the reader, we have expanded the
discussion related to this issue: “The damage mechanisms and the extent of damage were
very similar for the three system configurations because, as discussed in Ryan et al. (2013),
similar peak demands were observed in the three system configurations and vertical slab
vibration was insensitive to the presence of an isolation system. Thus, the damage
observations are presented and discussed without further mention of the system
configuration during which they were observed, but rather interpreted relative to the
horizontal floor acceleration and vertical slab acceleration to which they are closely
correlated.” (lines 193 to 199).
The lack of full repair certainly has some effect on the vulnerability of the system but we
believe it does not change the observations in a major way. Since we cannot quantify the
influence of the lack of full repair, it is stated in the paper as a limitation, and care is taken
not to draw conclusions based on specific components that were not repaired.
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