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SEISMIC DESIGN OF A MULTI-STORY CROSS LAMINATED
TIMBER BUILDING BASED ON COMPONENT LEVEL
TESTING
Shiling Pei1, Marjan Popovski2, John W. van de Lindt3
ABSTRACT: Cross laminated timber (CLT) is a new type of timber structural system that has just been introduced in
North America, but has been used successfully in Europe for over 20 years. There have not been any notable data sets
developed on the performance of tall CLT buildings during major earthquakes and there are no seismic design criteria
currently in place for CLT structures. Shake table testing of walls, assemblies and multi-storey CLT structures have
been performed, and development of dedicated nonlinear numerical models has been initiated. . In order to provide the
necessary information to develop practical seismic design procedure for mid-rise CLT buildings, a number of CLT
walls with different geometry and connector configurations were tested at FPInnovations, Canada to obtain the subassembly level hysteresis. Then, a simplified numerical model to predict the reverse cyclic behaviour of CLT walls was
developed and calibrated with the test results. Using the wall component capacity predicted by the model, a 10-story
CLT building was designed using a performance-based seismic design procedure known as direct displacement design
(DDD) to enable the structure to remain under the predefined drift limits at various seismic hazard levels. This design
procedure was adopted in a recent research effort to successfully design a 6-story light frame wood building tested at
Japan’s E-Defense shake table in 2009. In the present study described in this paper, the 10-story CLT building was
designed with 80% non-exceedance probability of remaining below 4% inter-story drift when subjected to a maximum
credible earthquake intensity level (2500 year return period) for the City of Los Angeles, California. The DDD of the
building was refined and verified with nonlinear time history simulation using a suite of bi-axial ground motions scaled
to the predefined hazard levels. Based on the performance-based design results and laboratory testing of individual CLT
shear walls, a response modification factor (R-factor) is proposed for structures with CLT wall components according to
current force-based design approach (i.e. ASCE 7), thus providing quantitative insight into CLT design using traditional
design procedures in North America .
KEYWORDS: Cross laminated timber, performance based seismic design, multi-story, numerical model
1 INTRODUCTION 123
In Europe, a structural system known as Cross
Laminated Timber (CLT) was introduced approximately
20 years ago for panelized construction of mixed use and
light-commercial buildings. The CLT building
components can be fully prefabricated in a factory
environment. CLT can be made with lumber as small as
1x4 inch nominal. When small lumber is layered
together to form the CLT panels, the final thickness of
the panel is such that the building has better fire
1
Shiling Pei, Assistant professor, South Dakota State
University, Brookings SD, U.S.A. Email:
[email protected]
2
Marjan Popovski, Principal Scientist, FP Innovations,
Vancouver, Canada. Email:
[email protected]
3
John W. van de Lindt, Professor and Garry Neil Drummond
Endowed Chair in Civil Engineering, University of Alabama,
Tuscaloosa AL, U.S.A. Email: [email protected]
resistance than light-frame wood, essentially similar to
heavy timber construction. The CLT system could be
classified as Heavy Timber (HT) in the International
Building Code [1] and can be a viable option for midrise buildings particularly.. However, most of the
existing tall CLT buildings do not consider seismic load
effects as they are not located in active seismic regions.
In fact the seismic resistance of CLT components and
systems has only been studied by a handful of
researchers around the world.
The current approach used to construct multi-story CLT
buildings relies on mechanical connectors to rigidly
connect the wall and floor panels together. Results from
quasi-static tests on CLT wall panels showed that the
connection layout and design has a strong influence on
the overall behavior of the wall [2] and the resulting
system can be very stiff [3]. At the component level,
quasi-static monotonic and cyclic tests have been carried
out on CLT walls to study the influence of boundary
conditions, magnitudes of vertical load, and the type of
anchoring systems (e.g. [4] [5] [6]) on performance and
capacity. It was shown that CLT panel walls can exhibit
significant levels of ductility if the boundary condition is
set up to allow rocking of the wall panels. At the system
level, a shake table test of a seven-story CLT building
was conducted by Ceccotti and colleagues at Japan’s EDefense facility in Miki, Japan, showing that the
structural panel members remained essentially elastic
and the accelerations in the higher levels were on the
order of 4g, compared to the earthquake input
acceleration of less than 1g [6]. Finally, a recent
handbook published by FPInnovations in Canada
summarizes recent development and practice in CLT
design and construction [7].
The objective of the study summarized in this paper was
to (1) perform a performance-based seismic design for a
10-story CLT apartment building which utilizes the
ductility of CLT panel walls observed in wall level
component tests, and (2) derive an appropriate strength
reduction factor (R-factor) for force-based design
procedures to achieve damage free performance during a
MCE (Maximum Credible Earthquake) level event. In
this study, a PBSD procedure termed the simplified
Direct Displacement Design (DDD) developed in the
NSF funded NEESWood project was used [8]. The load
resisting characteristics of the CLT walls during cyclic
loading were evaluated based on component tests
performed at FPInnovations. Based on the PBSD result,
a response modification factor (R factor) for force-based
design procedure is calibrated approximately to obtain a
design equivalent to the PBSD.
2 WALL LEVEL CALIBRATION
2.1 CLT WALL TESTS
In order to perform PBSD, the nonlinear deformationresistance characteristics of CLT walls must be obtained.
In this study, a nonlinear CLT wall model was developed
and calibrated based on wall component data from
FPInnovations.
Figure 1: Behaviour of Wall 11 during the testing
Figure 1 shows an example of the CLT wall test
conducted. Detailed description of the FPI test program
and specimen can be found in [6]. The impact of
Table 1: CLT wall tests conducted at FPInnovations
multiple factors including panel geometry, bracket type,
fastener type, and gravity loading on the lateral response
of CLT walls was evaluated using those tests. A brief
summary of the tested CLT walls and which test data
was used in this study is shown in in Table 1.
2.2 KINEMATIC MODEL
Because the shear deformation of the CLT wall panel
itself is not significant compared to the deformation at
the panel-to-floor or inter-panel connections, a
simplified model was developed that assumes all CLT
wall lateral deformation is a result of the CLT panels
rotating as a rigid body about the corners, as shown in
Figure 2. The lateral resistance of a CLT wall is
essentially a scaled summation of the load-slip resistance
of all the connectors engaged in the rocking movement
of the wall. The scale factor for each connector is a
function of their location and the geometry of the panel.
The resistance F at lateral displacement D may be
calculated as:
∑ (1)
Figure 2: CLT wall kinematics model
Figure 3: Hysteretic model for CLT connections
2.3 PARAMETER CALIBRATION
The “back-calibration” procedure used in this study can
be described as follows. The tested CLT walls were
modelled numerically using the simplified model and
were subjected to the same displacement protocols used
in the experimental tests. Then the model hysteresis was
compared with the experimental measurements. The
parameters for nails, screws, and hold-down connections
were adjusted systematically until the model closely
matched the observed experimental response. As a result
from this calibration, a group of connector parameters
were obtained. These connector parameters were then
used to develop the hysteretic responses for any given
CLT wall configurations (including configurations
different than those tested) based on the kinematic
assumption illustrated in Figure 2. These models were
also used to develop the design resistance values for the
CLT design tables. Figures 4 and 5 illustrated the
comparison between calibrated model prediction and test
results for two different wall configurations. Note the
effect of gravity is considered in the calibrated model as
well (Figure 5).
All parameters in equation (1) can be obtained from the
wall configuration (as shown in Figure 2) except for the
connector resistance. In this study, the connector
hysteresis was assumed to follow the CUREE 10parameter model, which has been widely adopted for
wood frame shear wall and connection modelling [9].
The behavior of the model and each control parameter is
shown in Figure 4. The parameters for each type of
connector were computed in this study by “backcalibration” using the wall test results.
Figure 4: Calibrated model compared with test (no
gravity)
and “DA” stands for Double All, meaning all brackets in
the panel are double sided. Based on Figure 6, the
designer can specify, for example, the “3DE”
configuration for the entire story, etc. It should be noted
that for the case of the wall panel with only 2 brackets ,
configurations DE and DA are identical. The maximum
number of 16d nails that can be put in a single bracket is
6, due to the limitation of the bracket to floor connection
strength. These typical walls will be used later in the
design of the 10-story CLT building.
Figure 5: Calibrated model compared with test (with
gravity)
The calibrated hysteretic parameters for three types of
connectors are listed in Table 2, including the Simpson
Strong-tie HTT-16 hold-downs installed at corners of
CLT walls and 16d spiral nails with D=3.9mm and
L=89mm with Simpson Strong Tie 90mm x 105mm x
105mm Bracket commonly used in CLT construction in
Europe. The parameters listed for the 16d nails are for a
single nail connection. The 16d with step joint is the
equivalent nail parameter to be used for multiple
panelled walls with step joints. With the numerical
model and connector parameters calibrated, the loaddeformation curve or the hysteresis curve for any CLT
wall configuration with specific connectors can be
developed. This provides a useful tool to generate CLT
wall backbone curves that will be used in PBSD and
tables with lateral load design values that are needed for
force-based design of CLT structures.
Hysteretic Parameters (N, mm)
Connector Type
HTT-16
16D-SN
16D-SN + step joint
3 DIRECT DISPLACEMENT DESIGN
(DDD) OF CLT BUILDING
3.1 CAPSTONE BUILDING FLOOR PLAN
Table 2: Calibrated CLT connector parameters
HTT-16
16D-SN
16D-SN+step joint
Figure 6: Typical wall configurations
K0
r1
r2
r3
r4
4378
140
158
F0
40032
3558
2669
0.002
0.005
0.001
F1
1779
178
89
-0.3
-0.2
-0.3
X
51
64
41
1
1
1
a
0.75
0.5
0.5
0.05
0.01
0.03
b
1.1
1.1
1.1
2.4 TYPICAL WALL CONFIGURATIONS
Several typical CLT wall configurations were considered
in this study as shown in Figure 6. It is assumed that a
structural CLT wall in the multi-story building will have
2, 3, or 4 brackets attached at the bottom of the wall,
providing connection between the wall and the floor
panel. The size of a single panel can vary from 0.96 m (4
ft) to 1.83 m (6 ft). It is assumed that for walls longer
than 1.83m, multiple1.22 m (4 ft) panels are to be
combined together and each panel will have bracket
configuration as shown in Figure 6. The notation “S”
stands for Single sided brackets for each location, “DE”
stands for Double sided brackets at the End of the panel,
The floor plan of the NEESWood Capstone structure, a
six-story light frame wood apartment building tested at
Japan’s E-Defense shake table in 2009 [10], served as
the model to develop the 10-story CLT structure, i.e.
similar floor plan. The elevation and floor plan are
presented in Figure 7. The building foot print is about 12
x 18 meters, with total height of 27.4 meter. All floor
plans are identical except for the top story, where a
penthouse unit may be integrated. The seismic weight of
the building was assumed to be 2.2kN/m2 for the first
story, 1.4kN/m2 for the roof, and 2.1kN/m2 for all other
stories. The total building weight is 4,537kN (462 metric
tons). Given the floor plan, the wall selection is
constrained in that only a limited amount of wall
segments can be placed in each story. The design began
by identifying the total usable wall segment length in
each direction from the architectural floor plan. The
numbers in Figure 7 show the maximum amount of CLT
wall panel segments one can put in any particular line of
the floor plan, which may or may not be fully utilized in
the design process as the designer may choose to select
some walls as “non-structural partition” and only apply
minimal connections. The exterior walls are shown with
window openings removed from the wall line. However,
for CLT panels the windows are typically pre-cut into
the wall panel, so the window opening will not
significantly affect the strength of the outside wall
segments. For the interior, the door openings do interrupt
the CLT walls (due to the height of the door openings)
so these walls are broken into smaller segments. Note
that the X direction is the longitudinal direction in the
floor plan while the Y direction is the shorter direction.
curve. The backbone curve for CLT walls used in the
design can be readily obtained using the simplified
model and calibrated connector parameters obtained
earlier. Thus DDD can be utilized to design the CLT
building presented in this paper. Detailed description of
the DDD procedure can be found in [12]. The basic
philosophy of DDD is to identify the required story
lateral resistance at prescribed drift levels enabling the
designer to select shear walls that satisfy this
requirement. Three performance objectives were adopted
in the design of the CLT building, limiting the maximum
inter-story drift under different hazard levels. Because
the inter-story drift level correlates well with seismically
induced damage to a building, the performance
objectives outlined in Table 3 will also ensure minimal
damage during these earthquake events.
Table 3: Performance Objectives
Seismic
Hazard
50%/50yr
10%/50yr
2%/50yr (MCE)
Performance Expectations
Inter-story Drift
Non-exceedance
Limit
Probability
1%
50%
2%
50%
4%
80%
Based on the performance objectives, the target point for
the backbone curve for each story was identified using
DDD. Design of the CLT building was conducted by
choosing the CLT wall configuration for each story to
produce a backbone curve that will satisfy (be larger
than) the corresponding target point. The CLT walls
selected for each story in both directions are listed in
Table 4. The resulted backbone curves for all stories
were plotted in Figure 8, together with the target points
for these stories resulting from DDD.
Table 4: Wall selection based on DDD
Figure 7: CLT building architectural plan and wall
segments location
3.2 PERFORMANCE OBJECTIVE AND DIRECT
DISPLACEMENT DESIGN
Direct Displacement Design (DDD [11]) was the design
approach employed in the NEESWood research project
to design the NEESWood Capstone building for
prescribed drift limits under different levels of seismic
intensity. It was validated through full scale shake table
test of a six-story light frame wood building. In fact, the
DDD method can also be applied to any lateral
resistance system which has a clearly defined backbone
ST
1
2
3
4
5
6
7
8
9
10
#
2
2
2
2
2
2
2
2
2
0
ST
1
2
3
4
5
6
7
8
9
10
#
4
4
4
4
4
4
4
4
2
0
CLT walls in X-direction
4.88 m
6.1 m
#
Con.
#
Con.
2
4DA
8
4S
2
4DA
8
4S
2
4DA
7
4S
2
4DA
7
4S
2
3DA
7
3S
2
3DA
7
3S
2
3DA
7
3S
2
2S
7
3S
2
2S
7
2S
2
2S
5
2S
CLT walls in Y-direction
1.53 m
1.83 m
2.44 m
Con.
#
Con.
#
Con.
4DA
2
4DA
14
4S
4DA
2
4DA
14
4S
4DA
2
4DA
14
4S
3DA
2
3DA
14
3S
3DA
2
3DA
16
3S
3DA
2
3DA
14
3S
3DA
2
3S
14
3S
2S
2
2S
14
2S
2S
2
2S
10
2S
-2
2S
8
2S
3.66 m
Con.
4DA
4DA
4DA
3DA
3DA
3DA
4S
2S
2S
--
#
1
1
1
1
1
1
1
1
1
0
6.1 m
Con.
4DA
4DA
4DA
4DA
3DA
3S
3S
3S
2S
--
In Figure 8, note that only targets for the level 3 design
are shown in the Figures since this case controls the
design. From the plots, one can see that all of the
backbone curves exceeded (are higher) the DDD target
points associated with them, both in the X and Y
directions. The design of the CLT building following
DDD is completed once all CLT walls are selected.
Practically, the other details of the building, such as
hold-down details, still need to be designed. Additional
PBSD procedures will be needed to design these details
[13]. In this study, it is assumed that there will be
adequate tie-down and over-turning restraints for each
story thus the design will only focus on the lateral force
resisting component.
Generally, it can be used to predict non-linear building
system seismic responses given the hysteretic response
of the components, which makes it suitable for the
purposes of this study.
The model for as-designed CLT buildings was
constructed in SAPWood and subjected to a suite of
earthquake ground motion records scaled to the
predefined hazard levels. Maximum inter-story drift
within the building from each simulation was extracted
to develop the distribution of the building maximum drift
for that corresponding hazard level. The probability of
exceedance associated with the design target can be
evaluated by plotting the empirical cumulative
distribution function curve based on all drift samples
within that hazard level. The ground motion suite
recommended for use in the U.S. Federal Emergency
Management Agency Document P-695 [15] was adopted
in this study, which includes 22 bi-axial far-field ground
motions scaled to three target hazard levels. The
response spectra of all these ground motions scaled to
the Maximum Credible Earthquake (MCE, level 3)
hazard level is shown in Figure 9. These bi-axial ground
motions were also rotated by 90-degrees and applied to
the model building. The building natural period
estimated through the numerical model is 1.12 sec. At
each performance level, the building was subjected to a
total of 44 ground motions. The maximum inter-story
drifts of the structure at any story and in either direction
were recorded and rank-ordered as empirical cumulative
distribution function curves shown in Figure 10. The
PBSD objectives are also shown in Figure 10 as square
shaped points for all three hazard levels. It can be seen
that the building satisfied all performance requirements,
i.e. the PNE (probability of non-exceedance) value for
the designated drift level is equal to or higher than the
performance requirements.
Figure 8: DDD target points and design backbone curves
4 PERFORMANCE EVALUATION OF
THE BUILDING WITH A
NUMERICAL MODEL
4.1 SAPWood PROGRAM
The CLT building designed using simplified DDD was
subjected to different levels of seismic hazard through
numerical simulation to verify that the design objectives
have been achieved. The analysis was conducted using
the software program Seismic Analysis Package for
Woodframe Structures (SAPWood) [14]. SAPWood is a
numerical tool specially developed to conduct nonlinear
time history analysis for wood frame buildings but uses
general enough hysteresis models that it has been used
for modelling other types of system as is being done in
the present case. It was part of the deliverables from the
NEESWood project and has been validated by numerous
component and system level shaking table tests,
including the 6-story NEESWood Capstone building.
Figure 9: Response spectra of ATC63 ground motions
scaled to MCE hazard level
design, i.e. the higher factor one uses, the higher the
computed R factor. Since at this point the design values
for CLT walls are not defined neither in US nor in
Canada, it was decided to do the analyses with the
design level equal to the ultimate divided by a factor of
2.5; that is utilizing only 40% of the wall ultimate
strength in the design. The resulting design resistance
values for standard CLT wall configurations using the
16d spiral nails is shown in Table 5.
Table 5: CLT wall design resistance values (kN)
Figure 10: Performance objective target points and
simulated structural performance
5 APPROXIMATE RESPONSE
MODIFICATION FACTOR R FOR
THE CLT BUILDING
The equivalent lateral force procedure (ELFP) is one of
the most commonly used design procedures for seismic
design of multi-story buildings in many force based
design codes (e.g. ASCE7-10 [16], NBCC [17]). The
lateral force demand for each story is obtained by
calculating the total base shear, and re-distributing it to
each story. This study utilized the recommended ELFP
procedure in the 2010 ASCE7 to calculate the level of
story shear forces required in traditional force-based
design for the CLT Capstone building. The objective of
this analysis is to identify a suitable response
modification factor (R-factor in ASCE7) to be used in
the ELFP design calculation, so that the force-based
design will result in a final design will be similar to what
has been obtained from PBSD procedure earlier. As the
design using PBSD has been verified by the nonlinear
time history simulation using the state-of-the-art
computer model, it is assumed that the codified design
using the calibrated R-factor will lead to satisfactory
building performance. In other words, when subjected to
an MCE event in Los Angeles, California , the CLT
building will have an 80% chance of not exceeding 4%
inter-story drift (experience only minor damage) as was
also observed in the NEESWood Capstone Building
shaking table tests.
Bracket#
Config
0.92 m
1.22 m
1.53 m
1.83 m
2x1.22 m
3x1.22 m
4x1.22 m
5x1.22 m
2
DE
17
23
29
35
32
43
55
67
DA
17
23
29
35
32
43
55
67
S
14
18
23
28
24
32
41
49
3
DE
25
34
42
51
44
58
72
86
DA
21
28
36
43
41
57
73
89
S
16
21
27
32
29
40
50
60
4
DE
28
37
46
55
49
65
82
98
DA
26
34
43
52
50
71
92
113
With the CLT wall selection for the building already
determined from PBSD, the equivalent force-based
design storey shear resistance can be obtained by adding
the resistance of all the walls at that storey based on
individual wall resistance listed in Table 5. This step can
be easily performed for each story in both the X and Y
directions. The resulting minimum story shear resistance
(minimum between the two directions) is the demand
that should be produced by force-based design
procedure, i.e. ELFP with the appropriate R factor. The
required story resistance is listed in Table 6 in the
“Target” column.
Following the ELFP in section 12.8 of ASCE 7-10, the
Importance Factor (I) was taken as 1.0; the building
period was calculated based on empirical formula
(ASCE7) to be 0.58 sec. The seismic hazard map design
values for the city of Los Angeles were obtained from
the USGS Design maps application (conforms to ASCE
7-10) as SDL=0.57g and SDS=1.62g. The R value was
changed manually until the final resistance distribution
matched the target resistance. The calibrated R factor
and the detailed calculation results are listed in Table 6.
Table 6: ELFP design of the 10-story CLT building
Cs
0.23
Story
For light frame wood buildings, the selection of wood
shear walls can be conducted based on the codified shear
wall design capacity tables (NDS Wind and Seismic
Supplement) once the story shear demand is obtained
using ELFP. The design resistance values for CLT walls
in this study were developed in a similar way to those for
wood-frame shear walls, which is by taking the ultimate
strength of the CLT wall backbone curve under
monotonic pushover test (which in this case was
generated using the simplified model and the parameters
obtained earlier), and dividing it by a factor. Note that
this factor will eventually affect the R factor used in
S
12
16
20
24
20
26
32
37
1
2
3
4
5
6
7
8
9
10
I
1.00
H
(m)
2.7
5.5
8.2
11.0
13.7
16.5
19.2
21.9
24.7
27.4
Tn
0.58
Wi
(kN)
489
467
467
467
467
467
467
467
467
311
SDL
0.57
Fi
(kN)
19
37
57
77
98
119
140
161
182
136
SDS
1.62
Vi
(kN)
1027
1009
971
914
837
738
620
480
318
136
R
4.30
Target
(kN)
788
788
743
677
597
597
562
448
319
204
Figure 11 illustrates the impact of the variation of the Rfactor to the storey shear distribution and how will the
shear resistance profile compare to the target one. The
story resistance of the PBSD design configuration for
stories 1 through 10 was plotted along with the
resistance calculated based on the ELFP, with different
R values. The dashed lines approximately represents the
appropriate lower and upper bounds for the R values
equal to 3 and 5.5, with the line in between representing
the calibrated chosen R-factor (4.3) in this study. Note
that the bounds presented here are not an indication that
the building should be designed using these R factors,
but just to illustrate the sensitivity of story demand
distribution to the R factor. In fact, the numerical
simulation indicated that the controlling drift level
occurs at higher stories. It is critical for force based
design to capture the PBSD demand at higher stories. If
the design only satisfies the lower stories, such as the
case when designing with R=5.5, such design will not
satisfy the drift requirement based on numerical
simulation results at the upper floors.
Figure 11: Story shear demand distribution with different
R factors
6 CONCLUSIONS AND DISCUSSION
Based on the analysis conducted in this study, CLT is a
viable option for mid-rise wood buildings up to ten
stories in moderate to high seismic regions, given the
buildings are correctly designed. It may also be
reasonable to expect only limited damage under MCE
level earthquake excitation in high seismic regions. By
adjusting the response modification factor, equivalent
force-based design that will meet the PBSD performance
objectives can be developed.
Based on the results of this study a value of R=4.3 can
be assigned to the analysed building and similar
symmetrical buildings in ASCE7 as examples of
structures with CLT wall panels that utilize spiral nails
in the brackets, provided that the design values for such
walls implemented in the material design standard are
similar to the values included in Table 5. The holddowns and adequate overturning restraints should always
be installed at both ends of a wall in order to ensure the
desired performance. From the comparison between the
story shear demand from ELFP and PBSD, it appears
that over-design for lower stories in the force-based
design procedure will result if the PBSD performance
objectives need to be satisfied using current ELFP
method.
It should be kept in mind that both the design resistance
value and the actual as-built resistance value were
simulated indirectly in this study through a numerical
model calibrated with limited number of wall level tests.
However, the conclusions on the R factors drawn from
this study are believed to be representative of a typical
CLT system. Also, it should be noted that the results in
this study are based on the assumption that the walls in
the system will be installed in a way that enables them to
rotate as in the tests conducted and generate similar
backbone curves. If the boundary conditions of the walls
differ significantly to the ones used in the analyses (the
walls are not allowed to rock), the backbone curves of
such wall configurations may be different and the results
of this study might not fully apply. However, the
kinematic assumption used in this study should be valid
for the inter-story range (0~3%) of interest in this study.
Thus the results are reliable unless one needs to
extrapolate the conclusions to very large inter-story drift
levels. In the future, if component testing is conducted
with different boundary conditions and backbone curves
are obtained for CLT walls at large deformation range,
the same procedure utilized in this study can be used to
derive the R-factor for that situation.
Finally, the response modification factor computed in
this study was based on one ten-story building with a
given floor plan and one type of fasteners used in the
brackets and the hold-downs. A variation of the R value
may exist if different fasteners are used in the brackets
and the hold-downs. Such variation may also exist as a
function of the floor plan and the building height. If
buildings with different heights are analysed, it is
unlikely that the R-factor will change significantly,
provided that the building floor plans remain
symmetrical in both directions. However, it is
recommended that further studies with a wider scope
look into issues related to R-factors for structures with
different archetypes and non-symmetrical floor plans
according to FEMA P-695 guidelines.
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