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.

Page 2

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

Page 3

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

Page 4

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.

Page 5

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

Page 6

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.

Page 7

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.

Page 8

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

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.

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.

Page 11

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.

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.

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.

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.

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