expectation meets reality: seismic performance of …...frames running parallel to the structural...
TRANSCRIPT
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EXPECTATION MEETS REALITY: SEISMIC PERFORMANCE OF POST-
TENSIONED PRECAST CONCRETE SOUTHERN CROSS ENDOSCOPY
BUILDING DURING THE 22ND
FEB 2011 CHRISTCHURCH EARTHQUAKE
Stefano Pampanin (1), Weng Yuen Kam (1), Gary Haverland (2) and Sean Gardiner (3)
(1) Dept. of Civil and Natural Resources Eng., Uni. of Canterbury, Christchurch, NZ
(2) Structex Ltd, Christchurch, NZ
(3) Formerly Structex Ltd, Christchurch, NZ
Abstract
The 22nd
Feb 2011 Christchurch earthquake highlighted the mismatch between the expectations of
building occupants and owners over the reality of engineered buildings‟ seismic performance.
Ductile plastic hinging behaviour of conventional reinforced concrete (RC) structures in large
seismic event such as those of 22nd
Feb are expected by structural engineers. However, the reality of
months of downtime and loss of occupancy of the building is not expected or desired by building
users and owners.
The innovative PRESSS-technology, utilising un-bonded post-tensioning precast concrete elements
to achieve re-centering behaviour, is a new approach to achieve low damage seismic performance,
in which building functional downtime and required structural repair are minimised. The Southern
Cross Hospital Endoscopy building is the first application of the innovative PRESSS design
technology in the South Island of New Zealand. The structure consists of four post-tensioned
precast concrete frames in the North-South elevation and two sets of post-tensioned coupled precast
concrete walls in the East-West elevation. Structex Ltd, in collaboration with the University of
Canterbury and Fletcher Construction, delivers the five-storey Warren & Mahoney-designed
building with a lower construction cost, reduced construction period and significantly improved
seismic performance.
The seismic performance of the Southern Cross Hospital Endoscopy building during the 22nd
Feb
Christchurch earthquake suggests that the expectation of clients can be met by innovative structural
solution. In addition to reporting on the details of the structural design and its performance during
the 22nd
Feb event, the paper demonstrates the use of non-linear numerical model as a design
verification and post-earthquake assessment tool.
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1 INTRODUCTION
The 22nd
Feb 2011 Christchurch earthquake highlighted the mismatch between the expectations of
building occupants and owners over the reality of engineered buildings‟ seismic performance.
Ductile plastic hinging behaviour of conventional reinforced concrete (RC) structure (Figure 1b) in
large seismic event such as those of 22nd
Feb are expected by structural engineers. The ultimate
limit state (ULS) and maximum credible earthquake (MCE) design level specified in modern
seismic codes, such as NZS1170:5, generally implies severe structural damage (to induce the
required level of ductility) in large earthquakes (Figure 1a). However, the reality of months of
downtime and loss of occupancy of the building is not expected and desired by building users and
owners. Similar mismatch of expectation and reality after the Northridge earthquake in 1994 led to
the concept of performance-based earthquake engineering [5].
Figure 1: a) SEAOC (1999) Performance design and performance matrix [5], b) Ductile beam flexural hinging in a multi-storey RC buildings in Christchurch.
The innovative PRESSS construction technology, utilising un-bonded post-tensioning precast
concrete elements to achieve re-centering behaviour, is a new approach to achieve low damage
seismic performance, in which building functional downtime and required structural repair are
minimised. The Southern Cross Hospital Endoscopy (SCHE) building (Figure 2) is the first
application of the innovative PRESSS design technology in the South Island of New Zealand. The
architect is Warren and Mahoney with Structex acting as the structural engineer. Fletcher
Construction is the principal contractor and Fulton Hogan is the post-tensioning sub-contractor.
Figure 2: Architectural rendition of the Southern Cross Hospital Endoscopy building. (Architect: Warren and Mahoney)
Modern seismic design for ULS and
MCE: Ductile plastic hinges
Longitudinal (East-West) Transverse
(North-South)
North
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This paper will briefly describe the structural details, seismic design and construction of the SCHE
building, highlighting the technology-transfer cooperation between research and industry. Then, in
addition to reporting on its performance during the Canterbury earthquakes (4th
Sept and 22nd
Feb
events), the paper demonstrates the use of non-linear numerical model as a design verification and
post-earthquake assessment tool.
2 DESCRIPTION OF STRUCTURAL SYSTEM
The structure is a three-storey building with a carpark level on ground floor. The building includes
several operation theatres at Level 2 and office space at Level 2 and 3. The roof level included an
additional plant room. The building footprint is 19m x 28m. The building is founded piles
foundation on 9m deep soft soil.
The lateral-load resisting system consisted of four limited ductile post-tensioned precast concrete
frames in the transverse (North-South) elevation and two sets of nominally ductile post-tensioned
coupled precast concrete walls in the longitudinal (East-West) elevation. Two perimeter cast-in-situ
frames running parallel to the structural walls served as drag ties as well as secondary resistance,
especially at levels 3 and 4. The structural layout is shown in the plan view presented in Figure 3.
Figure 3: Plan view of the building.
The frames incorporated post-tensioned tendons and top mild-steel reinforcements (see Figure 9).
The post-tensioned concrete walls had a combination of un-bonded mild steel reinforcements at the
base and U-shaped flexural plates (UFPs) in between the walls for energy-dissipation and damping
for the systems. Figure 4 shows the details of the North-side coupled walls with UFP elements. The
lateral loads are transferred to the frames and walls by the 90mm thick concrete topping. The level 3
and 4 are torsionally-sensitive in the longitudinal direction with the full-height walls acting only on
the south side. The transverse post-tensioned frames are designed for the torsion-induced demand.
The roof structure consists of large-span steel rafters (in the transverse direction) and purlins. The
floors are typically 200mm prestressed hollowcore units with 90mm topping. The precast
hollowcore, spanning in the longitudinal direction, were supported on the post-tensioned frames.
Post-
tensioned
moment-
resisting
frames
North
Full height post-
tensioned coupled
walls (South)
Half height post-
tensioned coupled
walls (North) Hollowcore
span
Cast-insitu
RC gravity
frame
Cast-insitu
RC gravity
frame
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The beam-elongation effect on the floor diaphragm from the post-tensioned frames was mitigated
by the use of cast-in-situ band beam-slab at the single-hinging beam-column rocking interface
(Figure 15d).
Figure 4: top-left) Elevation of the north-face half-height coupled post-tensioned walls; top-right) Detail of the U-shaped flexural plates (UFP) coupling elements; bottom) Cross-section
of south-face full-height walls (coupling detailed not shown).
3 DESIGN AND CONSTRUCTION
The PRESSS rocking systems relies on concentrated energy dissipation elements while the high-
strength steel post-tensioning tendons and precast concrete units remain generally elastic (or
confined crushing for concrete). The post-tensioning tendons also provide re-centering capacity to
the system, minimising residual deformation post-earthquake. The design and sectional analysis of
the PRESSS-technology have been thoroughly covered by the NZCS‟s PRESSS Design Handbook
[2] and the Concrete Standard NZS3101:2006 Appendix B [4]. Herein, some interesting aspects of
the design decision and construction phases in relation to the use of the PRESSS solution are
described.
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3.1 Decision process to use the PRESSS-system
As the PRESSS-technology had not been used in any Christchurch or South Island construction,
there was an augmented decision process to use the PRESSS-technology for this particular building.
In general, several key aspects in a relative chronological order:
1. Discussion with client, quantity surveyor and architect to outline the concept and the advantages, including possible price savings (in capital construction cost).
2. Complete construction and costing analysis review with Fletcher Construction.
3. Peer review and external consultation (with the University of Canterbury team) for design confidence and Building Consent application.
Pricing analysis and post-construction review confirmed that construction savings were achieved
(when compared to a monolithic precast concrete solution). Beam depth was reduced from 700mm
to 600mm, allowing services to be accommodated without raising the building height (a critical
issue in a built-up residential area).
3.2 Construction process
Precast elements and limited on-site casting can accelerate the construction process significantly.
Construction process (in an approximate sequence with some overlapping):
1. Sheet piling and excavation
2. Screw piles foundation
3. Foundation beams in-situ construction
4. Installation and grouting of ground floor precast concrete columns
5. Installation and propping of 1st floor beam and hollowcore floor.
6. Installation of upper floor columns and beams (in sequence).
Figure 5: The building under construction: a) Installation and erection of the precast columns and walls (North ends); b) Precast elements erection and propping from the North-East
elevation view.
7. Installation of shear walls
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8. Casting of slab toppings
9. Threading and post-tensioning of tendons in frames and walls.
10. Complete in-situ end beams, gutters etc.
11. Installation of prefabricated steel roof structure.
3.3 Construction challenges and PRESSS-solution
1. Limited cranage on site with basement limits the use of large precast concrete elements (e.g. full height frames). Therefore, post-tensioning of smaller precast elements (e.g PRESSS-
technology) is a desirable solution to the cranage problem.
2. Full height southern wall and partial height northern wall induced torsional demands onto the transverse direction frames. The post-tensioned frames, while supposedly „limited
ductility‟ in design, possessed substantial lateral strength and ductility to account for the
increased demand due to torsion amplification.
3. Screw piles required to reduce noise during construction (building in an established neighbourhood).
4. Need to keep all components simple and easy to be constructed, as well as “as conventional” as possible. There was also a need to avoid structural components that may be perceived as
expensive.
5. Post-tensioning anchorage blocks are large and required specific design to accommodate the steel work, spiral and transverse reinforcing. The 450mm wide column and 275mm thick
walls dimensions were driven by the need for the anchorage blocks (see Figure 6).
6. Design of construction sequence of the installation of the precast elements (beams, columns and floors) and post-tensioning work can accelerate the construction time significantly. As
with any new building system, it may be worthwhile for engineers to specify/clarify the
construction sequence with the contractors.
Figure 6: Detailling of the post-tensioned frames: a) Anchorage block and spiral reinforcing within the precast column; b) Precast columns with ducts for post-tensioning tendons and
mild-steel reinforcements; c) On-site post-tensioning.
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3.4 Further advantages of the PRESSS systems
1. Full length precast beam and columns – eliminating significant amount of insitu concreting. This led to more rapid construction time as full advantage of precast concrete was utilised.
2. Limited „structural damage‟ in the rocking plastic hinge zone. Self-centering capacity limits residual lean and displacement following an earthquake.
3. Analysis indicates lower level of floor acceleration compared to monolithic limited-ductility or nominally-ductile systems.
4. The new PRESSS-technology uses conventional building components and elements (e.g. precast beams, walls, drossbach ducts etc).
5. Reduced wall reinforcements (as post-tensioning tendons supplemented 40-50% of the required tension reinforcements.
4 INELASTIC MODELLING AND VERIFICATION
4.1 Inelastic 2D models
As part of the original peer-review process, inelastic 2D models of the SCHE building were
developed for non-linear push-over and time-history assessment. The modelling of the frame and
wall systems has been carried out using a lumped plasticity approach, following the procedure
described in NZCS‟s PRESSS Design Handbook [2]. Inelastic rotational springs are used in parallel
at the rocking connections (beam-to-column, column-to-foundation and wall-to-foundation)
connection to represent the self-centering contribution of the post-tensioned tendons (Non Linear
Elastic Hysteresis loop) and the dissipative contribution from the mild steel (Elasto-Plastic with
hardening, or bilinear). Frame, column and wall elements, away from the interface section are
modelled as elastic elements. Lumped mass and plasticity 2D model is implemented in the finite
element code Ruaumoko [1]. Figure 7 and Figure 8 illustrate the 2D models of the post-tensioned
frames and post-tensioned coupled walls respectively.
Figure 7: 2D elevation view of the post-tensioned frames model.
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Figure 8: 2D elevation view of the post-tensioned coupled walls model: a) Full height south walls; b) Half height north walls.
4.2 Section analysis of the rocking connections
The moment rotation curves of the jointed ductile connections were derived using the procedures
described in NZCS‟s PRESSS Design Handbook [2] while precast elements were analysed using
typical moment-curvature. Figure 9 shows the typical moment-rotation analysis result for the beam-
to-column rocking connections. Due to the discontinued bottom mild-steel reinforcements, the
negative beam moment capacity is higher than the positive beam moment capacity.
Figure 9: Moment-rotation evaluation of the rocking beam-column connections: a) Negative moment capacity i.e. tension at the bottom of the beam; b) Positive moment capacity.
4.3 Inelastic push-over analysis versus design values
Figure 10 shows the cyclic push-pull curves of the post-tensioned frames and the coupled walls
(only south walls result is shown). Figure 11 plots the inelastic push-over curves of the frames and
walls systems, superimposed with the seismicity demand (in terms of Acceleration-Displacement
Response Spectrum (ADRS) curves). The demand curves are reduced using computed hysteresis
damping curves from the actual system, and as higher energy dissipation is achieved at large
rotation/displacement, the actualised damping reduction factor increases at higher displacements.
Post-tensioning
Mild-steel
Total/Post-
tensioning
Total
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Page 9
Cyclic Push Pull
-2000
-1500
-1000
-500
0
500
1000
1500
2000
-4 -3 -2 -1 0 1 2 3 4 Drift (%)
Bas
e S
hea
r (k
N)
-544 -408 -272 -136 0 136 272 408 544Top Disp (mm)
Cyclic Push Pull
-2000
-1500
-1000
-500
0
500
1000
1500
2000
-4 -3 -2 -1 0 1 2 3 4 Drift (%)
Bas
e S
hea
r (k
N)
-500 -375 -250 -125 0 125 250 375 500Top Disp (mm)
Figure 10: Cyclic push-pull (displacement-controlled) analysis of a) Post-tensioned frames; b) South Coupled Walls.
0
0.05
0.1
0.15
0.2
0.25
0.3
0.35
0.4
0.45
0.5
0 0.5 1 1.5 2 2.5 3Top of Structure Drift (%)
Sei
smic
co
effi
cien
t (g
)
0
860
1720
2580
3440
4300
5160
6020
6880
7740
8600
Ba
se S
hea
r (k
N)
ULS Demand
ULS Elastic
SLS2 Demand
PushOver Capacity
Nominal Capacity (0.85)
Elastic 5%
1/1000 yrs
Damped (x)
1/1000 yrs
NL Push-over
Capacity
ø = 1.0
ø =0.85
x = 8.2%
Vb= 5455kN
x = 13.2%
Vb= 5892kN
x = 10.5%
Vb= 4766N
2.0%
x = 13.6%
Vb= 5093kN
Damped (x)
1/500 yrs
0
0.05
0.1
0.15
0.2
0.25
0.3
0.35
0.4
0.45
0.5
0 0.5 1 1.5 2 2.5 3 3.5 4
Drift at Effective Height (%)
Sei
smic
coef
fici
ent
(g)
0
800
1600
2400
3200
4000
4800
5600
6400
7200
8000
Base
Sh
ear
(kN
)
Tall Walls
Short Walls
Combined Walls
Combined Walls (phi=0.85)
Elastic Demand
1/1000 Damped Demand
1/500 Damped demand
Tall Walls
Damped (x )
1/1000 yrs
x = 20.2%
Vb= 4566kN
x = 17.0%
Vb= 5198kN
Damped (x )
1/500 yrs
Short Walls
x = 18.5%
Vb= 5293kN
1.8%
5%-damped
1/1000 yrsNominal
Capacity,
ø =0.85
NL Push-over
Capacity, ø = 1.0
Figure 11: Inelastic push-over curves on a base-shear versus drift at effective height domain: a) PT frames in transverse direction; b) PT walls in longitudinal direction.
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Page 10
The SCHE building was designed as an Importance Level 3 building (R=1.3, design seismic hazard
= 1/1000 years) according to NZS1170:5 [3]. A displacement-based design approach was adopted,
as this approach was more suited for the PRESSS system. For design, 8% and 13% equivalent
viscous damping (ξ) were assumed for the frames and walls respectively. The non-linear pushover
analysis of the frames and walls (next section) indicates the ξ to be 13% and 18.5% for the frames
and walls respectively. The design inter-storey drifts at the ULS (1/1000 years seismicity) were
2.0% and 1.8% for the frames and walls respectively. The non-linear pushover analysis suggests the
ULS inter-storey drift to be 1.67% and 1.75% for the frames and walls respectively. The push-over
assessment base-shear values were also comparable to the DDBD design base shear (frames:
5093kN versus 4286kN design; walls: 4656kN versus 3343kN design). The higher than expected
base shear for the walls system can be attributed to the high post-yield stiffness of the coupled
walls.
4.4 Inelastic time-history analysis results
Seven strong-ground motion records (listed in Table 1) were selected and scaled according to the
NZS1170:5 guidelines. The records are selected as representatives of the site and seismicity
conditions (Soil class D, 0.176g < PGA=0.22g < 0.33g, source magnitude, Mw of 5-7). Of the seven
records, four records had no forward directivity (near-fault) effects while three records had
distinctive directivity effects.
Table 1: Selected and scaled strong ground motions for time-history analysis verification.
Name Earthquake Event Year Mw StationRclosest
(km)
Soil Type
(NEHRP)
Unscaled
PGA (g)
Unscaled
PGV
(cm/s)
Scaling
Factor
Scaled
PGA (g)
Ground Motion with No Directivity Effects
EQ1 Superstition Hils 1987 6.7 Plaster City 21 D 0.155 20.6 2.44 0.379
EQ2 Northridge 1994 6.7 LA – Hollywood Stor FF 25.5 D 0.231 18.3 2.02 0.467
EQ3 Loma Prieta 1989 6.9 Gilroy Array #7 24.2 D 0.226 16.4 2.32 0.525
EQ4 Landers 1992 7.3 Yemo Fire Station 24.9 D 0.2095 29.7 2.01 0.421
Ground Motion with Directivity Effects ( Near Fault Earthquakes)
EQ5 Northridge 1994 6.7 Newhall Fire st. 5.92 D 0.59 97.20 0.49 0.288
EQ6 Imperial Valley 1979 6.6 El Centro Array #5 3.95 D 0.38 90.5 1.13 0.431
EQ7 Tabas, Iran 1978 7.35 Tabas 2 D 0.852 121.4 0.96 0.816
The inelastic time-history analysis results of the 2D models are shown in Figure 11 and Figure 14. The average inter-storey drift responses for the frames (in transverse direction) and the walls (in the
longitudinal direction) are shown in Figure 11a and Figure 14a-b respectively. In general, the bottom two floors exhibited a stiffer response for both frames and walls with lower inter-storey
drift. In fact, for the frames and the shorter north walls, the inter-storey drifts were significantly less
than the design level drifts – indicative of a need for a inelastic dynamic model as the assumed
damping might not be realised at all levels. For the whole building, the inter-storey drift responses
in the 2D models were about 1.0% at levels three and four, and 0.3-0.5% at levels one and two.
The floor acceleration was also checked as part of the design and verification. The floor
acceleration time history responses of the frames and walls are presented Figure 11b-d. In general,
the design intention is to limit the floor acceleration, particularly at level two and three to less than
1g, in order to protect the medical equipment. The stronger lateral resistance and significant energy
dissipation capacities (from having the north walls and more beam-column joints) at the lower two
floors managed to achieve the target floor accelerations.
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0
1
2
3
4
0.0% 1.0% 2.0% 3.0%
Interstorey Drift [%]
Lev
el
Mean
-/+ 1Std
Design
0
1
2
3
4
0.0 0.5 1.0 1.5 2.0
Floor acceleration [g]
Lev
elMean
-/+ 1Std
1
2
0.00 0.50 1.00 1.50 2.00
Floor acceleration [g]
Lev
el
Mean
-/+ 1Std
0
1
2
3
4
0.00 0.25 0.50 0.75 1.00
Floor acceleration [g]
Lev
el
Mean
-/+ 1Std
Figure 12: Average response from the time-history analyses of the 7 earthquake records suite: a) Inter-storey drift response for the frames; b-d) Floor acceleration responses for the
frames, south walls and north walls respectively.
4.5 3D model results
While not discussed thoroughly in this paper, the 3D model of the building (Figure 13) was
developed to analyses the torsion-induced amplification on the frames (transverse and longitudinal).
As time history analysis results of the 2D and 3D walls models presented in Figure 14 show, the torsional amplification in terms of inter-storey drift responses of the North (Tall) walls were
approximately 30% and 80% at Level 3 and Level 4 respectively.
Figure 13: 3D model of the Southern Cross Hospital Endoscope building.
0
1
2
3
4
0.00% 0.50% 1.00% 1.50% 2.00%
Interstorey Drift
Lev
el
Mean
+/- 1 Stdev
0
1
2
0.0% 0.2% 0.4% 0.6% 0.8% 1.0%Interstorey Drift
Lev
el
Mean
+/- 1 Stdev
0
1
2
3
4
0 0.5 1 1.5 2Inter-storey Drift (%)
Lev
el
Tall Walls
Short Walls
Figure 14: Average inter-storey drifts from the time-history: a) South walls 2D model; b) North walls 2D model; c) 3D model walls results.
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Page 12
5 PERFORMANCE OF THE BUILDING IN CANTERBURY 2010/2011 EARTHQUAKES
5.1 Observed structural and non-structural damage
No observable structural damage was detected in the building after the 4th
Sept 2010 7.1 Mw
Darfield earthquake. SCHE building was almost immediately re-occupiable (after an immediate
structural assessment). In the 22nd
Feb 2011 6.3 Mw Christchurch earthquake, the structure had
signs of significant transient movements (Figure 15a-c), especially in the East-West longitudinal
direction (consistent with the polarity of the Feb earthquake). On the top of the south walls, minor
crushing damage was observed at the interface between the coupled walls. Most of the UFPs had
Lueder yield lines, indicating the building‟s inter-storey drift of at least 0.5%-0.75% (corresponding
to the yield drift of the UFPs).
Figure 15: Observable damage (a-c) and non-damage (d) after the 22nd Feb 2011 earthquake.
Non-structural damage was more significant when compared with the structural damage. A
architectural glass panel on the staircase was cracked in both the earthquakes. In the Feb event, the
non-structural façade‟s connection to the wall spalled (Figure 15c), possibly consequence of the
fixity of the façade at the ground floor. Hospital staff reported one damaged water pipe and several
internal lining cracks as other observed non-structural damage.
In general, the seismic performance of the Southern Cross Hospital Endoscopy building during the
22nd
Feb Christchurch earthquake suggests that the expectation of clients can be met by innovative
structural solution.
5.2 Inelastic analysis using the 22nd
February 2011 recorded strong ground motions
The inelastic 2D model used in the peer review verification was used for a post-earthquake
reassessment of the SCHE building performance. Resthaven (REHS) recording station is 250
metres south-east of the SCHE building site. As both sites have significant soft soil layers (soil class
D), it is reasonable to infer similar strong ground motions at the SCHE site using REHS records.
Figure 16a shows the response spectra of several records from the 22nd February event, when compared to the NZS1170:5 design spectra. For brevity, only the analysis and results of post-
tensioned walls in the principal direction of the REHS records (along the East-West direction -
Figure 16b) will be discussed.
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Page 13
0.0
0.2
0.4
0.6
0.8
1.0
1.2
1.4
1.6
1.8
0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5
Sp
ectr
a A
ccel
erat
ion
/ S
a (g
ms-
2)
.
Period (sec)
NZS1170:5 (2004)
500-year motion
Mean of 4
CBD records
EQ2:CHHC (S89W)
EQ1:CBGS (NS64E)
EQ3:REHS
(S88E)
EQ1:CBGS
EQ4:CCCCEQ2:CHHC
EQ3:REHSPrincipal direction
NZS1170:5 (2004)
2500-year motion
N
EQ4:CCCC (N89W)
-0.6
-0.5
-0.4
-0.3
-0.2
-0.1
0
0.1
0.2
0.3
0.4
0.5
0.6
10 15 20 25 30 35 40
Acc
eler
atio
n (
g m
s-2
)
Time (sec)
Resthaven E-W
Principal Direction
Figure 16: a) 5%-damped response spectra of the records from Christchurch CBD in comparison with the NZS1170:5 design spectra; b) Resthaven station (REHS) recorded time-
history.
The inelastic time history analysis results is presented in Figure 17 for the south full height walls and the north half-height walls. The taller wall was significantly more flexible and attracted lesser
base-shear when compared with the shorter walls. While this resulted in torsional-induced
movement in the upper floors as well as in the transverse frames, the inelastic time-history results
indicated that the existing capacity of the walls and frames were adequate for the event. The
maximum inter-storey drift was approximately 2.5%, a good seismic performance considering the
REHS event is approximately 40-60% above the 2500-years return period design ground motion.
The immediate results suggested possible non-structural damage such as linings, façade and
services as per observed after the earthquake. Such a „quick result‟ may assist structural engineers
in assessing the structural health of the building rapidly with higher confidence than a visual
inspection alone. Nevertheless, it should be noted that the simplistic analysis do not consider the
secondary and redundant elements (e.g. gravity frames in the longitudinal direction and intrinsic
soil-structural damping).
-2.5
-2.0
-1.5
-1.0
-0.5
0.0
0.5
1.0
1.5
2.0
2.5
3.0
10 15 20 25
Inte
r-st
ore
y D
rift
(%
)
Time (second)
3F-4F
2F-3F
1F-2F
GF-1F
-2.5
-2.0
-1.5
-1.0
-0.5
0.0
0.5
1.0
1.5
2.0
2.5
3.0
10 15 20 25 30
Inte
r-st
ore
y D
rift
(%
)
Time (second)
1F-2F
GF-1F
Figure 17: Inter-storey drift time-history responses of the a) South full-height walls; b) North half-height walls.
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Page 14
6 CONCLUSIONS
The 22nd
Feb 2011 Christchurch earthquake is an unfortunate and tragic reminder of how far
earthquake engineering has progressed in the past 50 years. It also provides an avenue and
opportunity for the seismic engineering community to explore innovative construction technology
in order to deliver higher seismic performance, which matches the expectation of building
occupants and owners.
The Southern Cross Hospital Endoscopy (SCHE) building is a successful demonstration of the
PRESSS-technology for precast concrete multi-storey buildings with competitive cost, reduced
construction period and significantly improved seismic performance. The seismic performance of
the SCHE building during the 22nd
Feb Christchurch earthquake suggests that the expectation of
clients can be met by innovative structural solution.
Non-linear time-history modelling was demonstrated as an useful design verification as well as a
post-earthquake assessment tool.
7 ACKNOWLEDGEMENTS
Acknowledgement to Dr Dion Marriott for the assistance in the modelling phase of the study.
Special thanks to the building owner, Southern Cross Hospital Trust, the architect, Warren and
Mahoney, and the contractor, Fletcher Construction, for their willingness to be part of this
innovative and challenging project.
8 REFERENCES
[1] Carr, A. (2008). "RUAUMOKO2D - The Maori God of Volcanoes and Earthquakes." Uni. of
Canterbury, Christchurch, NZ, Inelastic Analysis Finite Element program.
[2] NZCS (2010). PRESSS Design Handbook, New Zealand Concrete Society (NZCS), Auckland, New
Zealand.
[3] NZS1170 (2004). NZS 1170:2004 Structural design actions, Standards New Zealand, Wellington, NZ.
[4] NZS3101:2006 (2006). "Appendix B: Special provisions for the seismic design of ductile jointed precast
concrete structural systems." NZS3101: 2006, Concrete standards, Standards New Zealand, Wellington, NZ.
[5] SEAOC (1999). Recommended lateral force requirements and commentary, Structural Engineers
Association of California (SEAOC), Sacramento, CA.