geot60-505 geohazards and large, geographically distributed systems

O’Rourke, T. D. (2010). Geotechnique 60, No. 7, 505–543 [doi: 10.1680/geot.2010.60.7.505]

505

Geohazards and large, geographically distributed systems

T. D. O’ROURKE�

A general classification for scale in geotechnical engineer-ing is used to explore the modelling of large, geographi-cally distributed systems and their response togeohazards. Both component and network performanceare reviewed. With respect to components, prototype-scale experiments of underground pipeline response toabrupt ground deformation are described, including con-trol of soil properties, soil–pipeline interaction, and per-formance of high-density polyethylene pipelines. Directshear (DS) apparatus size is shown to have a significanteffect on DS strength, and the most reliable DS device isidentified from comparative tests with different equip-ment. Mohr–Coulomb strength parameters for partiallysaturated sand are developed from DS test data andapplied in finite element simulations of soil–pipelineinteraction that show excellent agreement with prototype-scale experimental results. Apparent cohesion measuredduring shear failure of partially saturated sand is causedby suction-induced dilatancy. With respect to networks,the modelling of liquefaction effects on the San Franciscowater supply is described, and a case history of itssuccessful application during the Loma Prieta earthquakeis presented. The systematic analysis of pipeline repairrecords after the Northridge earthquake is used to iden-tify zones of potential ground failure, and correlate pipe-line damage rates with strong ground motion. Hydraulicnetwork analyses are described for the seismic perform-ance of the Los Angeles water supply, with practicalapplications for emergency response. The effects of Hur-ricane Katrina are reviewed with respect to the NewOrleans hurricane protection system, Gulf of Mexico oiland gas production, and interaction between electricpower and liquid fuel delivery systems. The sustainabilityof the Mississippi delta is discussed with regard to floodcontrol, maintenance of wetlands and barrier islands,and catastrophic change in the course of the MississippiRiver.

KEYWORDS: buried structures; case history; earthquakes;full-scale tests; laboratory equipment; laboratory tests; liquefac-tion; numerical modelling; partial saturation; sands; seismicity;shear strength; soil/structure interaction; suction

On utilise une classification generale pour l’echelle eningenierie geotechnique afin d’explorer la modelisation degrands systemes a distribution geographique, et leurreaction en presence de risques geologiques. On sepenche sur les performances des composants et du re-seau. En ce qui concerne les composants, des experiencessur prototype de reaction de conduites d’hydrocarburessouterraines a des deformations soudaines du sol sontdecrites, y compris le controle de proprietes du sol, lesinteractions sol - conduites d’hydrocarbures, et les perfor-mances de conduites en polyethylene a haute densite. Ondemontre que les dimensions des appareils a cisaillementdirect ont un effet significatif sur la resistance au cisaille-ment direct, et on identifie le dispositif de cisaillementdirect le plus fiable au moyen d’essais compares avecdifferents equipements. On developpe des parametres deresistance de Mohr-Coulomb pour du sable partiellementsature, sur la base de donnees d’essai a cisaillementdirect, que l’on applique a des simulations aux elementsfinis d’interactions sol – conduite d’hydrocarbures pre-sentant une excellente correspondance avec des resultatsexperimentaux sur prototypes. La cohesion apparente,mesuree au cours d’une rupture par cisaillement du sablepartiellement sature, est causee par une dilatance induitepar l’aspiration. En ce qui concerne les reseaux, on decritla modelisation des effets de la liquefaction sur les four-nitures d’eau de San Francisco, et on presente une etudede cas d’une application reussie au cours du tremblementde terre de Loma Prieta. On fait usage de l’analysesystematique des dossiers de reparation des conduitesd’hydrocarbures a la suite du tremblement de terre deNorthridge pour identifier des zones potentielles d’acci-dents de terrain, et determiner l’endommagement poten-tiels de conduites d’hydrocarbures en presence demouvement prononces du terrain. On decrit des analysesdu reseau hydraulique pour les performances sismiquesdu reseau de canalisations de fourniture d’eau de LosAngeles, avec des applications pratiques pour des inter-ventions en cas d’urgence. On examine les effets del’ouragan Katrina en fonction du systeme de protectionde la Nouvelle Orleans contre les ouragans, des installa-tions de production de petrole et de gaz dans le Golfe duMexique, et de l’interaction entre les reseaux de four-nitures d’electricite et les systemes de fourniture decombustibles liquides. On discute de la viabilite du deltadu Mississipi en ce qui concerne la lutte contre lesinondations, l’entretien des zones humides et des ıles-barriere, et la mutation catastrophique du cours duMississipi.

INTRODUCTIONGeotechnical engineers play a critical role in managing theperformance of large, geographically distributed systems thatare affected by geohazards such as earthquakes, floods,hurricanes and landslides. Systems, such as water supplies,levees and gas and liquid fuel supply networks, may coverthousands of square kilometres and be subject to many

Manuscript received 2 February 2010; revised manuscript accepted15 April 2010.Discussion on this paper closes on 1 December 2010, for furtherdetails see p. ii.� School of Civil and Environmental Engineering, Cornell Uni-versity, USA.

different ground response and geotechnical failure mechan-isms. The geotechnical factors affecting system behaviourhave broad implications for life safety and regional econom-ic stability.

The paper begins with a discussion of geotechnical scales,and how scale affects both problem definition and solutions.It then examines how soil–structure interaction of compo-nents is modelled for extreme loading conditions, by de-scribing laboratory tests of underground pipeline response toground rupture. The tests are used to illustrate how experi-ments performed at prototype scale under controlled condi-tions in the laboratory improve our understanding of soil–structure interaction, and lead to improvements in modellingof soil behaviour.

The geotechnical factors affecting regional system re-sponse to geohazards are examined with reference to earth-quake effects on the San Francisco and Los Angeles waterdistribution networks, as well as hurricane effects on theNew Orleans hurricane protection system. The paper de-scribes the modelling procedures for simulating water supplyperformance during earthquakes, and explores how model-ling results are used to address liquefaction and groundfailure hazards, develop relationships between earthquakemotion and underground pipeline damage, and make risk-based decisions about the emergency operation of watersupplies. The effects of Hurricane Katrina on New Orleansand Gulf of Mexico oil and gas infrastructure are described,and the implications of mega-scale problems associated withhurricane and river flooding, coastal restoration and cata-strophic change in river location are reviewed with respectto critical infrastructure performance.

SCALE IN GEOTECHNICAL ENGINEERINGThe physical dimensions involved in geotechnical engi-

neering span 15 orders of magnitude when compared on ametric scale, as shown in Table 1. The table is organised inaccordance with a similar scaling procedure proposed byChong & Davis (1999) for conceptualising research in civiland mechanical engineering. It provides a generalised modelin which there are six divisions of scale, ranging from nanoscale (10–9 m) at the molecular level to mega scale (106 m)at the level of geographic regions. For each division of scale,examples of natural features and engineering applicationsare given. For example, clay surfaces and double-layereffects at the nano scale influence the strength, volumechange and hydraulic characteristics of soils (e.g. Mitchell &Soga, 2005), and examples of engineering applications thatexploit nano-scale characteristics include electro-osmosis aswell as chemical and bio-remediation. Most geotechnicaltesting and characterisation occur at the milli and macroscales, which conform essentially to human dimensions and

distances within normal line of sight. At the mega scale,geotechnical engineering is applied, for example, to river,coastal and offshore systems, where flood protection struc-tures, transportation networks, and energy production anddelivery systems are constructed and managed.

There are many scales explicitly or implicitly embodied inmodels that represent large, geographically distributed sys-tems. For example, pipeline system response to an earth-quake is addressed at a minimum of two levels (O’Rourke etal., 2008), involving: (a) component performance, for whichsoil–structure interaction under earthquake loading is evalu-ated; and (b) system performance, for which the integratedbehaviour of the network is assessed. Such multi-scalemodels involve trade-offs between the detail required foraccuracy and the simplification needed for computationalefficiency and practical applications. Models of large, geo-graphically distributed systems generally focus on kilo- andmega-scale operations that properly account for network flowlaws and the spatial variability of both the system infrastruc-ture and geohazard effects. At the component level, rel-atively simple models may be based on expert opinion orempirical correlations between capacity and demand. Morecomplex component models are based on finite elementsimulations or fragility curves that express the probability offailure as a function of loading intensity (e.g. Shinozuka etal., 2000; Kafali & Grigoriu, 2007). The most importantcriteria for modelling large, geographically distributed sys-tems are model validation with respect to actual field per-formance, and critical review and acceptance by systemoperators. These validation and acceptance criteria establishthe level of practical worthiness necessary for credibleresults.

This paper explores nine orders of magnitude in themodelling of large, geographically distributed systems, fromthe behaviour of partially saturated sand to the performanceof the water supply system in Los Angeles and the floodprotection system in the Mississippi River delta. The conceptof scale is used as a unifying theme to show how the shearstrength of sand at the milli-scale level is linked throughsoil–structure interaction of pipeline components at themacro-scale level to evaluate the earthquake response of aregional water supply system at the mega-scale level. Theeffects of Hurricane Katrina are used to show how geo-technical problems are shaped by fluvial, deltaic and cyc-lonic processes that are distributed over many thousands ofkilometres, with enormous impact on large, geographicallydistributed systems.

GEOHAZARDSTable 2 provides a summary of major geohazards, with

information about their spatial characteristics as well as their

Table 1. Summary of scales in geotechnical engineering

Nano10–9 m

Micro10–6 m

Milli10–3 m

Macro10þ0 m

Kilo10þ3 m

Mega10þ6 m

Molecular Micrometres Millimetres Metres Kilometres Regions

Clay surfacesDouble-layer effects

ClaySilt

SandGravelCobbles

SlopesSoil deposits

AquifersReservoirsRiver crossings

River, coastal and offshore systems

Cation exchangeElectro-osmosisChemical and bio-remediation

Suspension groutsReinforcing stripsPolymer grids

FoundationsRetaining wallsPipe/conduits

BridgesDamsTunnelsLevees

Water suppliesTransportation networksElectric power and fuel systems

506 O’ROURKE

direct and indirect effects. Geohazards have attracted sub-stantial public and policy-making attention, with nearly90 000 dead and missing, and substantial infrastructure de-struction in China from the 2008 Sichuan earthquake (EERI,2008; Stone, 2009); over 228 000 people killed by the 2004Indian Ocean tsunami (Iwan, 2006; Cosgrave, 2007); andover $100 billion in direct losses (Jordan & Paulius, 2006)and nearly 2000 dead and missing (IPET, 2008a) fromHurricane Katrina in 2005. At the time of submitting thispaper, the dead and missing after the 2010 Haiti earthquakeare estimated at approximately 200 000, with nearly com-plete destruction of the physical and governmental infra-structure of its capital city, Port-au-Prince. Geohazards havea strong influence on professional practice and research ingeotechnical and civil engineering. Moreover, they increasethe scope and extent of geotechnical interests from site-specific practices to regional concerns and interdisciplinarypolicies.

Geohazards have generated substantial interest in lifelinesystems. Lifelines include electric power, gas and liquidfuels, telecommunications, transportation, waste disposal,and water supply (O’Rourke et al., 2001). This grouping caneasily be extended to include flood protection and leveesystems for coastal and river communities. Taken individu-ally, or in the aggregate, such systems are intimately linkedwith the economic well-being, security and social fabric ofthe communities they serve. Thinking about lifelines helps

clarify the resources and services that are essential for com-munity resilience, and encourages an exploration of theinterdependences among critical infrastructure systems(O’Rourke, 2007).

The geohazards in Table 2 include earthquake, tsunami,hurricane and typhoon, flood, volcano, drought, subsidenceand landslide. All represent a life safety threat, but onlyselect effects on the built and natural environments are listedin the table. Most geohazards affect broad areas, primarilyin the range of 102 to 105 km2. Coverage may be restrictedto zones as small as 10 m2 owing to local subsidence andslope failure, or may involve a region exceeding 106 km2 inthe case of drought.

Special attention is given to earthquakes and hurricanes inthis paper. The widespread effects of earthquakes are espe-cially well documented. The 1906 San Francisco earthquake,for example, involved significant ground shaking throughoutan area of about 48 000 km2 (Lawson et al., 1908; Ellsworth,1990). Seismic ground waves generated by the 1811–12New Madrid earthquakes caused damage in an area exceed-ing 500 000 km2 (Johnston and Schweig, 1996). Earthquakeeffects on transportation, liquid fuel, electric power andwater supply systems are covered extensively in the technicalliterature (e.g. O’Rourke et al., 2004). The economic impactof earthquakes is well illustrated by the 1995 Kobe earth-quake, with direct losses of $100–150 billion (Hamada etal., 1995).

Table 2. Geohazard spatial characteristics and effects

Geohazard Spatial characteristics Direct effects Indirect effects

Earthquake Area of 102 to 105 km2 withsignificant shaking and groundfailures

Strong ground shakingFaulting, landslides, liquefaction,and consolidation of loose soils

Threat to all civil infrastructure, including buildings,waterfront facilities, transportation and other lifelinesystems

Tsunami Travel distances of 1 to 104 km,affecting 10 to 102 km ofcoastline and waterfront

FloodingImpact from flood debrisErosion and removal of soil,vegetation, dwellings, andinfrastructure

Threat especially to coastal and waterfront facilities. Largerun-up and inundation of coastal buildings and infrastructure

Hurricaneand typhoon

Eye of storm covers 10 to102 km2. Area of high windsand storm surge may cover 102

to 106 km2

Flooding and high windsImpact from airborne andwaterborne debrisErosion and destruction ofwetlands, soils, and waterfrontfacilities

Threat especially to coastal and waterfront facilities. Windand flood damage to buildings, flood protection systems, andlifelines. Transmission towers and electrical substationsvulnerable

Flood 10 to 103 km along main rivercourses and from 10 to 106 km2

within floodplains

InundationFlood loads and impact of flooddebrisSoil erosion, scour, andundermining of structures

Threat especially to river crossings and floodplaininfrastructure, such as bridges, levees, port facilities,dwellings, and lifeline systems

Volcano Area affected by volcanicactivity varies from 10 to105 km2

Lava and pyroclastic flowsVolcanic ash and projectilesLahars and floodsNoxious gases

Infrastructure damage from high heat, fire, flooding, debrisflows, projectile impact and ash falls. Saturated ash loadingand obstruction of infrastructure

Drought Large regional effects of 105 to.106 km2

Loss of water and vegetationLoss of cropsIncreased fire hazardSoil erosion and shrinkage

Loss of reservoirs, watersheds, food supply and industry.Wildfires and urban fires. Slope failure due tovegetation loss

Subsidence Local to regional settlementaffecting 10 m2 to 103 km2

Settlement, change of grade, andlateral ground movementFloodingReduced aquifer storageGrowth faults with abrupt verticaland horizontal grounddisplacement

Adverse effects on stormwater drainage, canals andtransportation systems. Differential settlement and lateraldeformation of buildings and lifelines. Flooding due tosubsidence is a threat to all civil infrastructure

Landslide Volume varies from , 10 m3

to . 107 m3

Largest historic volumes of2–3 km3

Lateral ground movement,settlement, and change in grade.Movements may be rapid ordevelop gradually.Landslide dams

Catastrophic damage to buildings and infrastructure whenmovements are large and rapid. Landslide dams threatendownstream infrastructure

GEOHAZARDS AND LARGE, GEOGRAPHICALLY DISTRIBUTED SYSTEMS 507

A hurricane or typhoon is a tropical cyclone with maxi-mum sustained wind speed of at least 119 km/h (e.g.Emanuel, 2005; Rauber et al., 2005). Its cloud shielddiameter is typically on the order of 800 km, althoughdiameters of the largest storms can exceed twice this value(Fitzpatrick, 2006). The worst effects are associated withstorm surge flooding, which historically accounts for 90% ofcyclone-related deaths (Longshore, 1998). In 2005 Hurri-canes Katrina and Rita were responsible for disrupting 100%and 94% of all Gulf of Mexico offshore oil and gasproduction respectively (MMS, 2006). Hurricane Katrinacaused enormous damage in the New Orleans area, resultingin the greatest natural disaster loss in US history, as well ascontroversy about the performance and failure mechanismsof the regional hurricane protection system (e.g. ASCEHurricane Katrina External Review Panel, 2007; Christian,2007; Seed et al., 2008a).

EXTREME LOADING CONDITIONSExtreme loading associated with geohazards often occurs

with large plastic, irrecoverable soil deformation. Such de-formation may involve geometric changes in the soil mass,such as localised shear rupture, heave, and void formation.Soil–structure interaction under these conditions is fre-quently characterised by a peak, or maximum, interactionforce. The extreme loading of soils occurs collaterally withthe extreme loading of structures. Examples include soil–structure interaction associated with pipelines subjected tofault rupture, piles affected by landslides, and undergroundfacilities exposed to lateral soil movement, flooding andundermining. Structural response often entails large plastic,irrecoverable deformation with material and geometric non-linearities. Hence analytical and experimental modelling forsoil–structure interaction under extreme conditions requiresthe coupled post-yield simulation of both soil and structure.Such behaviour imposes significant demands on modelling,thus requiring large-scale experimental validation of analyt-ical models to improve the simulation process.

Extreme loading, especially in conjunction with geohaz-ards, often affects large lifeline systems. Consider, forexample, Fig. 1, which is a photograph of the corner of Walland Williams Streets in New York City during the 1920s.The photo illustrates at least three important features of theurban environment. First, much of the critical infrastructureis located underground, and its fate is intimately related tothat of the surrounding ground. Second, underground infra-

structure is removed from direct observation unless uncov-ered, and its state of repair and its proximity to otherstructures are often unknown. Third, congestion increasesrisk due to proximity. Damage to one facility, such as a castiron water main, can cascade rapidly into damage in sur-rounding facilities, such as electric and telecommunicationcables and gas mains, with system-wide consequences. Soilsurrounding critical underground infrastructure is frequentlyboth the perpetrator and the mediator of loading that canaffect the systemic performance of an entire city.

LARGE-SCALE EXPERIMENTSA significant trend in geotechnical engineering has been

the implementation of large-scale testing facilities for soil–structure interaction, such as those at the Japanese NationalResearch Institute for Earth Science and Disaster Preventionthat have been used to characterise soil–pile interactionduring liquefaction (Tokimatsu & Suzuki, 2004) and thelarge-scale split box experiments at the George E. Brown, JrNetwork for Earthquake Engineering Simulation (NEES)equipment site at Cornell University (e.g. Palmer et al.,2006; O’Rourke & Bonneau, 2007). The large-scale facilitiesallow for testing at the macro-scale level (see Table 1), sothat conditions in the field can be simulated reliably underlaboratory control.

It is not possible to model with accuracy the soil displace-ment patterns at all potentially vulnerable locations. It ispossible, however, to set an upper bound on deformation bysimplifying spatially distributed ground deformation asabrupt soil movement. Detailed studies of fault deformationdisclose that abrupt soil rupture and offsets are indeedrecurrent patterns at active faults (Bray et al., 1994). Theythus establish a baseline to evaluate soil–pipeline interactionrepresentative of upper-bound ground movements for earth-quakes, landslides, subsidence and flood undermining.

As shown in Fig. 2, split-box testing has the capability ofimposing abrupt soil displacements on buried pipelines con-sistent with those at fault crossings and the margins oflateral spreads and landslides. Relative displacement is gen-erated along a movable interface between two test basins, orboxes, containing soil and the buried pipeline. The pipelineis buried in soil that is placed, compacted and tested accord-ing to field construction practice. The dimensions of theexperimental boxes are selected on the basis of computa-tional modelling and previous test experience to minimisethe influence of the test facility boundary effects.

Figure 3 shows the current generation of split-box testbasin that utilises between 90 and 100 metric tonnes ofpartially saturated sand per test and 1.2 m of strike–slipdisplacement. In the figure, the displacement was providedby two hydraulic structural actuators with load capacities of445 kN tension/650 kN compression and a one-way stroke of1.28 m. Other tests have been performed with four hydraulicactuators positioned to pull the movable section of the testbasin with a combined capacity of 1.5 MN in tension.

As explained by O’Rourke et al. (2008), a special con-veyor system was fabricated to move large quantities of soiland facilitate soil placement in about 4 to 5 days. Soil for atypical experiment was placed in seven 200 mm lifts andone top lift of 100 mm. Each lift was compacted with twopasses of a gasoline-powered plate tamper. The soil was aglacio-fluvial sand, referred to as RMS graded sand, pro-duced in accordance with New York State specifications forconcrete sand, and representative of sand used for backfilland engineered construction. Between 240 and 320 measure-ments of dry unit weight and moisture content were recordedfor each set-up. Dry unit weight, ªdry, was measured in situusing a nuclear density gauge according to ASTM D6938-

Fig. 1. Underground infrastructure at Wall and WilliamsStreets in New York City, 1920s (photograph provided by theConsolidated Edison Company of New York, Inc., New York,NY)

508 O’ROURKE

07b (ASTM, 2008a), and moisture content was measured insitu according to ASTM D2216-98 (ASTM, 2008b). Typi-cally, from five to seven dry unit weight determinations weremade per cubic metre of soil.

Figure 4 presents the grain size distribution curve for theRMS graded sand, which is angular to sub-angular withnon-plastic fines. Examination assisted by microscopeshowed approximately 71% by volume of siltstone, fine-

grained sandstone, shale and limestone fragments, and 29%quartz grains. The mean grain size D50 for the soil is0.67 mm, and the coefficient of uniformity Cu is 2.83.

The peak friction angle �ds-p associated with ªdry meas-ured by the nuclear density gauge was determined fromregressions that had been developed between ªdry and �ds-p

using direct shear (DS) tests for partially saturated sand(O’Rourke et al., 2008). The value of �ds-p is reported interms of total stress, because suction in the moist sand (andthus effective stress conditions) is not measured directly. Fordry sand, �ds-p ¼ �9ds-p.

Figure 5 shows plots presented by O’Rourke et al. (2008)of the mean dry unit weight ªdry, gravimetric water contentw and peak friction angle �ds-p of the sand for each lift withrespect to depth for five full-scale split-box pipeline experi-ments. As shown in the figure, a remarkable degree ofcontrol was achieved over the experimental soil propertiesfor many tests with large volumes of material. The prepon-derance of the measurement data show placement of thesand in the range of ªdry ¼ 15.5–15.8 kN/m3, w ¼ 3.5–4.5%and �ds-p ¼ 39–408. Statistical control procedures describedby Trautmann et al. (1985) were applied to determineconfidence intervals for the sand properties based on meancharacteristics of the entire soil mass. For example, the 95%confidence intervals on the mean ªdry and �ds-p for anygiven test were �0.02–0.05 kN/m3 and �0.10–0.198 respec-tively. Fig. 4 illustrates the spread in soil properties lift bylift among several tests, whereas the confidence intervalspertain to overall average values for each individual test.

Partially saturated sand is different from saturated or drysand in that it is affected by matric suction, um. Thepresence of um promotes apparent cohesion in sand, andincreases its dilatancy relative to dry and saturated sand atthe same ªdry. To help evaluate the properties of the partiallysaturated test sand, soil water retention curves (SWRCs)were obtained with Tempe cells, and tensiometer measure-ments were taken during large-scale tests to generate rela-tionships between w and um. The Tempe cell, described inASTM D6836-02 (ASTM, 2008c), was used to developSWRCs by desorption (drying), while tensiometers, de-scribed by Dane & Topp (2002), were used to measure um

in situ during large-scale tests.

Pipetrenchcross-section

Special trench andbackfill at faultcrossing

(a)

(b)

Buried pipeline Elbow

Permanent grounddeformation (PGD)

Compacted sand

Welded steelpipeline

Fixed box

ElbowCompacted sand

Welded steelpipeline

Fixed box

(c)

Straight pipe

Pipe with elbow

Fig. 2. Simulation of ground rupture effects on pipelines by split-box tests: (a) PGD effect on buriedpipelines; (b) PGD effect on buried pipelines with elbows; (c) experimental concepts

Actuators

1·2 m

6·6 m

3·2 m

2·3 m

Fig. 3. Large-scale split-box test basin at the Cornell UniversityNEES equipment site

10 1 0·1 0·01

Particle diameter: mm

0

50

100

% P

assi

ng

Angular to sub-angular29% quartz71% sedimentaryrock fragments

Fig. 4. Grain size distribution curve for RMS graded sand usedin the large-scale tests


Figure 6 shows a plot of w against um, comparing Tempecell and tensiometer measurements for partially saturatedRMS graded sand at ªdry ¼ 15.7 kN/m3 with equations pro-posed by Pham et al. (2005) for the main drying curve(MDC) and main wetting curve (MWC). Pham et al. devel-oped equations to represent the MDC and MWC for sandbased on a known SWRC and key soil properties. Therelationships presented by Pham et al. are for volumetricwater content, but are adapted here for w. The Tempe celldesorption and tensiometer data plot closely with respect tothe MDC and MWC respectively. The tensiometer measuresa vacuum pressure, generated as water is transported acrossthe tensiometer porous stone into the surrounding soil. Thismechanism of measurement is analogous to generating anSWRC by adsorption (wetting), and thus the tensiometerdata follow the MWC. The RMS graded sand naturallydrains in the large-scale test basin to w � 4%, which isabout equal to the residual water content, as shown in thefigure.

SHEAR STRENGTH TESTING OF SANDOver 700 DS tests were performed to quantify the DS

behaviour of sand for the soil–pipeline interaction tests.Such testing corresponds to the milli-scale level (see Table1), and was intended to develop DS testing procedures forthe accurate determination of strength. Of primary interestwere the effects of DS box dimensions. It has been observed

(e.g. Dietz, 2000; Cerato & Lutenegger, 2006) that smallratios of box dimension to particle ratios lead to boundaryconstraints on DS behaviour and overestimation of strengthand dilation parameters, while large ratios lead to progres-sive failure and underestimation of the same parameters.Moreover, to minimise errors, several investigators havemodified the DS equipment to restrain upper frame and loadpad rotation (e.g. Jewell & Wroth, 1987; Shibuya et al.,1997; Lings & Dietz, 2004).

Tests were performed with DS testing devices, havingnominal interior horizontal dimensions of (a) 60 mm 3 60mm, (b) 100 mm 3 100 mm and (c) 300 mm 3 300 mm, forwhich the key horizontal dimension L is 60 mm, 100 mm,and 300 mm respectively. Photographs of the 60 mm,100 mm and 300 mm test boxes are shown in Fig. 7. Thetest programme is described in detail by Olson (2009), andonly the salient features are summarised here.

The 60 mm DS tests were performed using a conventionalWykeham Farrance DS testing apparatus, which conforms tothe specifications of ASTM D 3080-04 (ASTM, 2008d) andBSI 1377-7 (BSI, 1990). The device had a conventionalbrass lower box and an upper box machined from maplewood. The wooden upper frame weighed 0.9 N in compari-son with the 13.0 kN brass upper frame, thus minimising thevertical load transferred from the device to the soil failureplane.

The 100 mm and 300 mm DS tests were performed usinga custom-fabricated DS testing apparatus, designed in accor-dance with the characteristics of the 100 mm test devicedescribed by Dietz (2000) and Lings & Dietz (2004). Loadis applied in this device to a pair of yokes at the midpointof the upper frame sidewalls, such that the application ofshear force is coincident with the central horizontal plane ofthe soil specimen. Great care was taken to prepare the soilwith proper ªdry, as described by Olson (2009). Strict soilplacement and tamping procedures were employed to pro-mote uniform conditions with respect to soil fabric and insitu stress.

Numerous tests were performed to evaluate the influenceof gap size (separation between upper and lower DS testframes) and edging (rubber skirts to prevent extrusion of soilthrough the gap). A 1 mm gap without edging was adoptedfor the tests. It was found to be sufficiently small that soilloosening and deterioration did not occur in the separationbetween the upper and lower boxes, yet sufficiently largethat measurements of peak shear strength and dilation were

φds-p: degrees15·2 15·6 16·0 16·4 0·02 0·03 0·04 0·05 0·06 0·07 38·0 39·0 40·0 41·0 42·00

0·4

0·8

1·2

1·6

2·0

Dep

th: m

Test 1

Test 2

39·0–40·0°

γd3: kN/m w

Test 3

Test 4

Test 5

15·5–15·8 kN/m3 3·5–4·5%

Fig. 5. Mean dry unit weight, water content and peak friction angle by lift against depth for five large-scale tests (O’Rourke et al., 2008)

0·01 0·1 1 10 100Matric suction, : kPaum

0·04

0·08

0·12

0·16

0·20

0·24 MDC (Pham , 2005)et al.MWCTempe cellTensiometer

Main dryingcurve, MDC

Main wettingcurve, MWC

γd315·7 kN/m�

Residual water content

0

Gra

vim

etric

wa

ter

cont

ent,

w

Fig. 6. Soil water retention curve for RMS graded sand used inthe large-scale tests

510 O’ROURKE

not amplified by constraints imposed by the testing appara-tus.

For dry sand tests the peak friction angle �9ds-p wasobtained from the peak shear stress �9p for a given normalstress � 9N as

tan�9ds-p ¼ �9p� 9N

(1)

The peak angle of dilatancy, łp, was obtained from themaximum change of vertical soil displacement, v y, relativeto the horizontal displacement, vx, between the upper andlower frames of the DS test apparatus, using

tanłp ¼ dv y

dvx

(2)

Figure 8 shows �9ds-p against ªdry for RMS graded sandtested at � 9N ¼ 22 kPa. Tests with sand having different grainsize characteristics were performed (Olson, 2009), but only

the test results for RMS graded sand are summarised herein.Sand placed at low ªdry < 15.6 kN/m3 was contractivethroughout the entire test. The data trend for sand at lowªdry differs markedly from the linear trend at higher unitweights, and they are not included in the linear regressionfits to the data.

In general, the regression line for the 60 mm box is 4–68higher than that for the 100 mm box, and 6–78 higher thanthat for the 300 mm box. The 100 mm box trend line has acoefficient of determination r2 of 0.98, higher than thevalues of 0.90 and 0.89 for the 60 mm and 300 mm boxesrespectively, providing an indication of the improved consis-tency of data obtained with the 100 mm box. Fig. 9 showsłp against ªdry for RMS graded sand. The 60 mm box lineplots 3–78 higher than those for the 100 mm and 300 mmboxes, which are statistically indistinguishable from eachother. The trend lines all have relatively high r2 values, withthe 100 mm box showing the highest r 2 ¼ 0.94.

L 60 mm�L 100 mm�

L 300 mm�

(a) (b)

(c)

Fig. 7. Photographs of DS test devices: (a) 60 mm test box; (b) 100 mm test box; (c) 300 mm test box


The test results show clearly that the size of the DSapparatus has a significant effect on the determination of�9ds-p. The results from the 60 mm apparatus significantlyoverestimate �9ds-p and łp, compared with measurementswith the larger devices, despite the fact that it complies withASTM D3080-04 (ASTM, 2008d). The small device is alsoconsistent with BSI 1377-7 (BSI, 1990), except that the ratioof box height to maximum particle size, H/Dmax, is about 9compared with the recommended value of 10. Since most ofthe discrepancy between values of �9ds-p is related to differ-ences in łp, it appears that the 60 mm device tends toamplify dilatancy in the RMS graded sand.

Given the systematically higher values of �9ds-p measuredwith the 60 mm conventional test apparatus, use of this

device was discontinued. Tests were performed with the100 mm device, fabricated in lightweight aluminium accord-ing to the design of Lings & Dietz (2004), and shearstrength values obtained with this device are presented inthe remainder of this paper.

SHEAR STRENGTH OF SANDThe DS test does not provide a peak stress state at

maximum obliquity with respect to the Mohr circle: that is,it does not represent a point on the Mohr circle of stress thatis tangent to the Mohr–Coulomb failure surface. Constitutivelaws that utilise the Mohr–Coulomb failure surface are read-ily available in software, such as ABAQUS and FLAC. Theyrequire specification of the maximum obliquity effectivestress friction angle �9ps (also referred to as the plane-strainfriction angle) for plane-strain problems of soil–structureinteraction.

At peak DS strength for dry granular media, researchershave shown coaxiality of stress and incremental strain,meaning that the principal axes of stress and incrementalstrain coincide (e.g. Lings & Dietz, 2004; Bolton, 1986).Figs 10(a) and 10(b) are Mohr’s circles of incremental strainand stress respectively. Assuming coaxiality, and recognisingthat the horizontal axis is a direction of zero linear incre-mental strain (zero extension), one can use the dilation angleand Mohr’s circle of stress to develop a relationship betweenthe DS values of �9p and � 9N, and the plane-strain values ofshear, �9ps, and normal stress, � 9pN. Davis (1968) first derivedthe equation linking the parameters shown in Fig. 10(b) as

tan�9ds ¼cosł sin�9ps

1 � sinł sin�9ps

(3)

which at the critical state (ł ¼ 0) becomes

tan�9ds ¼ sin�9crit (4)

in which �9crit is the critical-state friction angle. Bolton(1986) and Jewell & Wroth (1987) provide similar deriva-tions.

It is assumed that the coaxiality of incremental strain andstress applies for partially saturated RMS graded sand.Given the low values of suction measured for the test sand,it is likely that the relative orientations of stress and incre-mental strain are not materially different from those for drysand. Moreover, it is shown later that strength and volumechange characteristics evaluated on the premise of coaxialitysupport a rational mechanics-based explanation of observedbehaviour and strength parameters that leads to a veryfavourable comparison between analytical and experimentalresults.

Figure 11(a) shows � against � for a partially saturatedsand with c and �. From this figure it can be seen that thecentre and top of the Mohr circle of stress is related to(�p/�N) through łp, such that

�1 þ �3

2¼ �N þ �p tanłp (5)

�1 � �3

2¼ �p

cosłp

(6)

in which �1 and �3 are the major and minor principalstresses respectively.

If one develops a linear regression for (�1 + �3)/2 (centreof the Mohr circle) and (�1 � �3)/2 (radius of the Mohrcircle) from the DS data, one obtains a plot with intercept cand slope tanÆ, as shown in Fig. 11(b). The plane-strainfailure envelope for maximum obliquity is also illustrated inFig. 11(b). It can be shown by trigonometry that

15 16 17 18 19γdry

3: kN/m

20

25

30

35

40

45

50

55: d

egre

esφ

� ds-p

L 60 mm�

L 100 mm�

L 300 mm�

Linear fit: 60 mmL �



Linear fit: 60 mm( 15·6 kN/m )

5·57 51·5926, 0·90


6·59 73·2111, 0·98

L

y xn r

L

n r

��

� ��

��

��

γ

γ

dry3

2

dry3

2y x �

Linear fit: 300 mm(all )

5·87 62·399, 0·89

L

y xn r

�

� �

� �

γdry

2

Fig. 8. �9ds-p against ªdry for three different DS test device sizeswith RMS graded sand

15 16 17 18 19

γd y3: kN/mr

0

5

10

15

20

25

30

ψp:

deg

rees

L 60 mm�

L 100 mm�

L 300 mm�





6·98 109·4811, 0·94

Linear fit: 300 mm(All )

7·30 114·619, 0·91

L

y xn r

L

y xn r

�

�

� �

� �

�

�

� �

γdry3

2

dry

2

γ�


8·63 132·75260·91

L

y xnr

��

� ��

γdry3

2

Fig. 9. łp against ªdry for three different DS test device sizeswith RMS graded sand

512 O’ROURKE

�1 � �3

2¼ c þ �1 þ �3

2tanÆ (7)

�1 � �3

2¼ cps cos�ps-p þ

�1 þ �3

2sin�ps-p (8)

in which �ps-p and cps are the plane-strain friction angle atpeak DS strength and plane-strain cohesive intercept respec-tively.

The radius of the Mohr circle, (�1 � �3)/2, is simulta-neously satisfied by equations (7) and (8) if and only if

cps ¼c

cos�ps-p

(9)

�ps-p ¼ sin�1 tanÆð Þ (10)

Hence equations (9) and (10) give the parameters thatrepresent the plane-strain failure envelope from data plottedin accordance with equations (5) and (6).

Figure 12 shows plots of (�1 � �3)/2 against (�1 + �3)/2

for partially saturated RMS graded sand with ªdry ¼15.8 kN/m3 at w ¼ 4–5% and for �N from 2.1 to 98.9 kPa.A linear regression fit to the data gives a line with a slopeof Æ ¼ 34.38 and an intercept of c ¼ 0.9 kPa. Using equa-tions (9) and (10), the plane-strain envelope at maximumobliquity can be calculated from c and Æ, and is also plottedin Fig. 12. This line has �ps-p ¼ 43.08 and cps ¼ 1.2 kPa. Ascan be seen in the figure, the Æ–c envelope is a good fit tothe ((�1 � �3)/2, (�1 + �3)/2) data, and the �ps-p and cps

τ�

Pole forplanes

σ�

(0, d /2)xyγ

(d , d /2)ε γyy yx

Pole fordirections

dε

d /2γ

CO

Ψ

Ψ

(a)

(b)

( , )σ�ps τ�ps

( , )σ�xx τ�xy

( , )σ�yy τ�yx

φ�ps

φ�ds

( , )� � �σ τN

Fig. 10. Mohr circles for: (a) incremental strain; (b) stress(after Lings & Dietz, 2004)

σ3

σ3

σ1

σNps

σps σN σ σ1 3

2

�

σ σ1 3

2

�

σ σ1 3

2

�

σ σ1 3

2

�

σ

φps p�

φps p�

φps p�

φps p�

φds p�

φps p�

Ψp

R � �τ

Ψp

pcos

σ1

τ

σp

N( )

ds

(a)

(b)

c

cpscotφps p�

σ

α

τ

τ

τp

τps

τps

cdscps

cps

Fig. 11. (a) DS and plane-strain strength parameters and theirrelationship with the top of the Mohr circle; (b) c–Æ and cps –�psp relationships

( - )/2 against ( )/21 3 1 3�σ σ σ σLinear fit: ( )/2 against ( )/21 1 3� �σ σ σ σ3

Plane-strain envelope at max. obliquity

0 20 40 60 80 100 120σN: kPa

0

20

40

60

80

100

τ: k

Pa

α 34·3°�

Plane-strain envelope at maximum obliquity43·0°

1·2 kPa

Linear fit: ( )/2against ( )/2

34·3°0·9 kPa12

φ

σ

α

ps p

ps

1 3

1 3

� ��

��

��

c

cn

σσ σ

φps p� 43·0°�

Fig. 12. � against �N for partially saturated RMS graded sandat peak strength, ªdry 15.8 kN/m3


envelope is a good approximation for the plane-strain envel-ope at maximum obliquity, as it passes closely to the tangentpoint of each Mohr’s circle. As with �ds-p and cds, �ps-p andcps are assumed to be constant over �N ¼ 2.1–98.9 kPa.

COHESION IN PARTIALLY SATURATED SANDTaylor’s flow rule for the DS strength of dry sand can be

expressed as

�9p� 9N

� �dry

¼ �p

�N

� �dry

¼ tan�9ds-p

¼ sin�9crit þ tan łpð Þdry

(11)

and for the DS strength of partially saturated sand as

�p

�N

� �p:sat:

¼ tan�ds-p þcds

�N

¼ sin�crit þ tan łpð Þp:sat:

(12)

Data were obtained by Olson (2009) from DS tests on dryand partially saturated RMS graded sand with the same ªdry

for �N from 2.1 to 98.9 kPa. Moreover, all tests wereprepared using the same method with the same appliedenergy. The partially saturated sand was prepared within arelatively narrow range of w ¼ 3.9–5.0%. Within the boundsof test measurement error, it was shown that there is nosignificant difference between tan�9ds-p � sin�9crit for drysand and tan�ds-p � sin�crit for partially saturated RMSgraded sand. Assuming these differences to be equal andcombining equations (11) and (12) results in

cds

�N

¼ tan łpð Þp:sat� tan łpð Þdry

(13)

which shows that the increase in dilatancy between dry andpartially saturated sand is equal to the normalised cohesion,cds/�N.

Figure 13 shows tanłp against �N for dry and partiallysaturated RMS graded sand at ªdry ¼ 15.8 kN/m3. Whenthere was more than one test at a given �N for either the dryor partially saturated sand, the average łp of all tests at that

�N was plotted in the figure. Linear regressions are shownfor both datasets.

Values of tan(łp)p:sat: � tan(łp)dry from Fig. 13 areplotted in Fig. 14 against �N. The differences in tanłp areplotted as both the differences between the data points inFig. 13 and the differences in the regression equations. Thedata point differences are fitted with a power regression inFig. 14, which has r2 ¼ 0.81. Also plotted in the figure iscds/�N against �N for cds ¼ 2.1 kPa, corresponding to thevalue measured for partially saturated RMS graded sand atªdry ¼ 15.8 kN/m3 (see Fig. 12). Because of the form of theequations plotted in Fig. 14, it appears that cohesion in-creases without bound as �N approaches zero. This is onlyapparent; the cohesion cannot exceed a maximum valuerelated to the surface tension mobilised between soil parti-cles.

The three curves in Fig. 14 are statistically indistinguish-able, showing that the difference in tanłp for dry andpartially saturated RMS sand is well represented by cds/�N.The apparent cohesion generated in partially saturated RMSgraded sand is therefore related to increased dilatancy. Itappears that suction increases the interference among thesand particles, thus elevating the shear resistance of partiallysaturated sand compared with dry sand. This mechanism ofcohesion related to dilatancy is consistent with critical-stateconcepts for the DS strength of sand described by Schofield(2005).

MODELLING SOIL–PIPELINE INTERACTIONAs mentioned previously, to utilise the Mohr–Coulomb

constitutive models in geotechnical software, it is necessaryto know the plane-strain friction angle and cohesion for two-dimensional (2D) analyses. The 100 mm DS test box fabri-cated in accordance with the design of Lings & Dietz (2004)was used to determine the appropriate strength parameters.The DS values of �ds-d and cds were then converted toplane-strain values of �ps-p and cps, following the proceduresdescribed above.

To validate this approach, a series of finite element (FE)simulations were performed with the software ABAQUS(2006) using �ps-p and cps determined for partially saturatedRMS sand, and compared with the results of full-scale 2Dtests performed at the Cornell NEES site. Fig. 15 presents aschematic diagram of the large-scale 2D test basin, which wasfilled with partially saturated RMS graded sand compacted in200 mm lifts. Over 100 nuclear density gauge measurementsof soil unit weight and w were made for each test.

The basin was designed to measure the lateral forceagainst displacement of pipelines through the application of

Partially saturated RMS graded sandLinear fit: partially saturated RMS graded sandDry RMS graded sandLinear fit: dry RMS graded sand

1 10 100σN and : kPaσ�N

0·01

0·1

1·0

tan(

)Ψ

p.sa

tp

0·01

0·1

1

tan(

)Ψ

dy

pr

Partially saturated RMS graded sandPower fit: 15·8 kN/mln 0·58 ln 0·12

0·95, 7

γdry3

2

�� y x

r n� �

Dry RMS graded sandPower fit: 15·8 kN/mln 0·19 ln 2·34

0·75, = 7

γdry3

2

��

�y x

r n

Fig. 13. tan(łp)p:sat against �N and tan(łp)dry against �9N forRMS graded sand at ªdry 15.8 kN/m3

0 20 40 60 80 100σN and kPaσ�N:

0

0·1

0·2

0·3

0·4

0·5

tan(

)ta

n()

Ψp

p.sa

t.p

dry

�Ψ

0

0·1

0·2

0·3

0·4

0·5

c/

dsNσ

Difference in data point pairs

c cds N ds/ , 2·1 kPaσ �

Power fit: difference in data point pairs

ln 0·95 ln 0·35y x� � �n r7, 0·81� �2

Difference in regression equations

Note: 15·8 kN/mfor all data and regressions

γdry3�

Fig. 14. Values of tan(łp)p:sat 2 tan(łp)dry and c/�N against �Nand �9N for RMS graded sand at ªdry 15.8 kN/m3

514 O’ROURKE

horizontal force with the two long-stroke (1.2 m in onedirection) hydraulic actuators as shown in the figure. Hor-izontal force was measured on each side of the box with aload cell, and lateral movement was measured with Tempo-sonic displacement transducers. The loading arm was de-signed so that the test pipe could rise without verticalrestraint as it was displaced laterally through the soil. Therate of pipe displacement was 2.4 mm/s. The test basin andloading conditions were similar to those used in previousfull-scale tests (e.g. Trautmann & O’Rourke, 1985:O’Rourke et al., 2004), with the main exception being size.The internal dimensions of the test basin were 2.44 m 32.44 m in plan and 1.82 m in depth. The end effects of wallfriction were minimised by the relatively large width of thetest basin, and by lining the interior of the box withFormica

1

and glass, both of which provide for relatively lowangles of interface friction.

Tests were performed on pipelines 124 mm and 172 mmin external diameter, buried at a pipe centreline depth todiameter ratio Hc/D between 3.5 and 7.5. The test pipes hada 2.5 mm thick high-density polyethylene (HDPE) externalcoating, which is a typical coating for pipelines in the field.

Figure 16 shows typical geometric conditions incorporatedin the FE model. The meshes were composed mostly of 700to 800 eight-node plane-strain elements. A refined mesh wasused within a distance of approximately two pipe diametersfrom the test pipe. Key dimensions are given in Fig. 16. Thepipe steel had properties similar to those of ASTM GradeA-36 steel, and was assumed to be linear elastic. The testbasin was fixed in the x and y directions. Interface elementswere used between the soil and both the interior surface of

the test basin and pipe circumference. The interface frictionangle between the soil and pipe with HDPE external coatingand vertical basin wall (Formica

1

and glass) was taken as0:6�9ds-p on the basis of DS tests performed by Trautmann(1983) and Olson (2009). The interface friction angle be-tween the soil and plywood base of the test basin wasestimated as 0.8�ds-p.

A Mohr–Coulomb model was used in which �ps-p ¼ 43.08and cps ¼ 1.2 kPa, corresponding to ªdry ¼ 15.8 kN/m3 forpartially saturated RMS graded sand. To represent strain-softening, the model proposed by Anastasopoulos et al.(2007) was used to diminish linearly both �ps-p and łp toresidual values at �crit and 0 respectively, from the plasticstrain at �ps-p to the plastic strain at �crit, using the resultsof DS testing as

ªpf ¼ dxp � dxy

Hþ dxf � dxp

dFE

(14)

in which ªpf is the shear strain increment beyond yield at

which there is no dilation; H is the thickness of the DSspecimen; dFE is the FE element size; and dxy, dxp, dxf arethe DS test horizontal displacements at yield, peak strengthand �crit at which full softening occurs. In addition, cps wasreduced from its maximum at �9ps-p to a small residual valueof 0.1 kPa at ªp

f . This reduction is consistent with measure-ments by Olson (2009) of little to no cohesion in DS testsfor partially saturated RMS graded sand at large displace-ments. A FORTRAN subroutine developed by Robert &Soga (personal communication, 2009y) to apply theAnastasopoulos et al. (2007) model in the Mohr–Coulombmodel in ABAQUS was used in the 2D FE method simula-tions.

Following the work of Yimsiri et al. (2004), laboratorytest data were used to estimate the soil modulus for the 2Dsimulations. Multiple linear regression (MLR) analyses wereperformed on Young’s modulus E, vertical stress at pipecentre �vc, and ªdry from the results of 25 full-scale testsconducted by Trautmann & O’Rourke (1985), Turner (2004)and Olson (2009) to obtain an expression for E as a functionof �vc and ªdry with highest statistical significance. Forpartially saturated RMS graded sand an empirical equationwas developed as

E ¼ 2 3 10�13:97 ªdry�0:0378vc

� �13:7(15)

in which ªdry and �vc are expressed in units of kN/m3

and kPa respectively. Equation (15) provides an equivalent

(a)

(b)

Reactionbeam Actuator

Level of soil backfill

Direction ofpipe movement

0·4 m

Load cell

Floor level

Actuator

Actuator

Direction of pipemovement

Buried pipe

Rea

ctio

n be

am

2·4 m

2·4 m

1·2 m

Fig. 15. (a) Plan and (b) cross-sectional elevation of full-scale2D tests to measure horizontal force against displacement

3·5 / 7·5124 mm or 174 mm

H DD

c ��

�

Hc

0·48 m D

Hbk

2·2 / 5·2Interface elements

� �H Dbk

2·44 m

Fig. 16. Typical finite-element mesh for simulation of 2D large-scale test results

† Personal communication of FORTRAN subroutine for analysis ofsoil–pipe interaction.


modulus that is empirically calibrated and suitable for directuse in the elasto-plastic Mohr–Coulomb model available inABAQUS.

Figure 17 shows the relationships of dimensionless forceF9 against dimensionless displacement Y9 from the numericalsimulations. The dimensionless force F9 ¼ F/ªtHcDL, whereF is the horizontal force, ªt is the total soil unit weight, andL is pipe length. The dimensionless displacement Y9 ¼ Y/D,where Y is the lateral displacement of the pipe with respectto the soil. The numerical simulations for partially saturatedRMS graded sand show excellent agreement with the experi-mental data for all partially saturated tests. The numericalsimulations compare favourably with the experimental resultsin terms of pre-peak, peak force, and post-peak behaviour.

The FE and 2D test results for partially saturated RMSgraded soil are compared in Fig. 18, in which dimensionlessforce Nq is plotted as a function of dimensionless depthHc/D. There is excellent agreement among the analytical and

test data for all Hc/D. In general, the FE results overpredictthe measured dimensionless peak forces by a small marginof 1–8%.

The favourable agreement between analytical and experi-mental results provides supporting evidence for obtaining�ps-p and cps from DS data according to the approachdescribed previously. It also marks a transition from char-acterising engineering behaviour through experimental re-sults to characterising that behaviour on the basis ofcomputational simulation. Over the past 25 years the experi-mental results of Trautmann & O’Rourke (1985) for drysand have provided guidance in estimating lateral forcesgenerated by soil–pipeline interaction under large grounddeformation (ASCE, 1984). All the experimental results ofTrautmann & O’Rourke can now be replicated with fidelityby FE analyses. Moreover, the analytical methods are beingextended to partially saturated sand, which is relevant for amuch broader range of field conditions. Soil–structure inter-

Hc Hc

HcHc

D D

D D

0 0·2 0·4 0·6

Dimensionless displacement,(a)

Y/D

Dimensionless displacement,(c)

Y/D

Dimensionless displacement,(b)

Y/D

Dimensionless displacement,(d)

Y/D

0

4

8

12

Dim

ensi

onle

ss fo

rce,

/F

HD

Lγ t

cD

imen

sion

less

forc

e,/

FH

DL

γ tc

Dim

ensi

onle

ss fo

rce,

/F

HD

Lγ t

cD

imen

sion

less

forc

e,/

FH

DL

γ tc

Olson 2D moist 3

Numerical analysis

γt316·7 kN/m

Pipe diameter 174 mm�

�

γt316·5 kN/m


�

γt316·7 kN/m


�

γt316·5 kN/m


�

0 0·2 0·4 0·6

0

4

8

12

Olson 2D moist 1

Numerical analysis

0 0·2 0·4 0·6

0

4

8

12

16

Olson 2D moist 6

Numerical analysis

0 0·2 0·4 0·6

0

4

8

12

16

Olson 2D moist 7

Numerical analysis

Fig. 17. Analytical and experimental dimensionless force–displacement relationships for partially saturated RMS graded sand:(a) Hc/D 3.5; (b) Hc/D 5.29; (c) Hc/D 6.5; (d) Hc/D 7.5

516 O’ROURKE

action in partially saturated soils involves complex soilbehaviour and modelling demands, and provides a futureresearch area with great potential.

HIGHLY DUCTILE PIPELINESHigh-density polyethylene (HDPE) pipelines represent a

class of pipe with the capacity to accommodate a largeamount of ground deformation associated with geohazards,such as earthquakes, flooding, subsidence and landslides. Toinvestigate the capacity of HDPE pipelines for abrupt groundmovement, and develop analytical capabilities for evaluatingsoil–structure interaction under such conditions, large-scaletests were conducted on HDPE pipelines subjected to 1.22 mof strike-slip displacement at a crossing angle of 658 withrespect to ground rupture, as illustrated in Fig. 19. Eachpipeline was instrumented with between 80 and 140 straingauges, many with the capability of measuring strains ashigh as 20%. All tests were conducted with partially satu-rated sand, with the same ªdry, w, �ds-p and cps describedpreviously. The experimental pipelines consisted of HDPEpipes, manufactured by the Chevron Phillips Chemical Com-pany under the commercial name DRISCOPLEX. Descrip-tions of the full-scale testing, measurement systems, andtests results are provided elsewhere (O’Rourke et al., 2008),and only select features of the test results are providedherein.

The measured axial and bending strains along a nominal400 mm diameter HDPE pipeline with wall thickness of24 mm are plotted in Fig. 20 at a strike-slip displacement of

1.22 m. An inset diagram shows a schematic representationof the deformed shape of the pipe within the test basin.Another inset diagram shows the crown, invert and spring-line locations. Axial strains are the average of the pipecrown and invert strains, and bending strains are expressedas one half the difference between the springline strains. Thebending strain so calculated is the incremental strain causedby pipeline flexure relative to the axial strain. The maximumaxial strains coincide with the ground rupture, and decreasewith increasing distance from this location. Flexural strainsare zero at the ground rupture, consistent with the point ofcounterflexure for axisymmetric pipeline deformation.

There is a remarkable degree of consistency in the meas-ured bending strains. There are some differences in axialstrains, caused in part by Teflon

1

wrapping to protect specialsensors, as shown in the figure. In essence, the Teflon

1

isolates the pipeline from soil shear forces, resulting inhigher axial strains at the south end of the pipe. The maxi-mum measured strain was 8%, representing the combinedaxial and bending strains, and was located at a distance ofapproximately 1.0 m from the ground rupture either side ofthe fault. Moreover, the axial pipeline load decreased by40% within 2 h after ground rupture. Because HDPE isviscoelastic, it has the beneficial effect of reducing load withtime at anchorages outside the ground rupture zone. The testresults demonstrate the benefits of HDPE pipeline ductilityand viscoelastic response in accommodating permanentground deformation. The maximum measured strains for1.22 m of strike-slip displacement were far below strainlevels associated with rupture or creep instability of the pipewall.

The tests, however, showed a 12% increase in verticaldiameter at the location of ground rupture (O’Rourke &Bonneau, 2007), which was accompanied by 6% loss of thepipe internal cross-sectional area. The results therefore in-dicate that squeeze-off due to bending and locally highlateral soil forces is an important, potential failure mechan-ism for an HDPE pipeline with larger diameter-to-thicknessratio, D/t, affected by severe ground deformation.

As illustrated in the next section, highly ductile pipelinescan play a critically important role in the performance ofwater supplies vulnerable to earthquake-related ground de-formation. The use of tests results at the macro scaleprovides information indispensable for improving systemperformance at substantially larger scales. Through study ofthe San Francisco water supply it is seen that system per-formance may be controlled by a limited number of criticalpipelines. Improvements focused on strategically locatedcomponents can therefore be used for effective system-widerisk reduction.

2 4 6 8

Dimensionless depth, /H Dc

0

4

8

12

16

20

24

28M

axim

um d

imen

sion

less

forc

e,/(

)N

FH

DL

qc

�γ

Large-scale 2D partially saturated RMS graded sand tests

RMS FEM simulation

Experimental data, 0·75, = 4r n2 �

Trend of numerical simulation

D

Hc

Fig. 18. Dimensionless force against dimensionless depth formaximum lateral pipe force in partially saturated sand

65°

1·22 m

3·2 m

10·6 m

Initial pipe position

Initial box position

Final box position

Final pipe position

Fig. 19. Plan view of large-scale pipeline test, with key dimen-sions and geometry

�5 �4 �3 �2 �1 0 1 2 3 4 5Position: mS( ), N( )� �

�2

0

2

4

6

8

10

Str

ain:

%

Test 1: Axial

Test 1: Bending

Test 3: Axial

Test 3: Bending

Test 2: Axial

Test 2: Bending

Location oftactile force

sensors

Splitbox

Pipe

N

Crown

Springline

Invert

Fig. 20. Comparison of strains in 400 mm diameter pipelinesduring large-scale tests


SAN FRANCISCOThe mainland city of San Francisco, CA, is located in an

area of approximately 120 km2, where its buildings andlifeline systems are a living laboratory for infrastructureresponse to geohazards at the kilo-scale level (see Table 1).With respect to earthquakes and infrastructure, San Francis-co is distinctive for at least two reasons. First, it is the placewhere modern, geographically distributed systems were firstaffected by a major earthquake. Second, it is a place wheretwo severe earthquakes have occurred, generating liquefac-tion at the same locations and disrupting critical watersupplies in similar ways.

Fire following earthquake is regarded by many as thegreatest single threat to San Francisco. Approximately 96%of the city buildings are timber structures (Scawthorn et al.,2006b), built in close proximity to each other. Fire in onesuch building will spread next door in about 10 min. With a15 km/h wind, fire unchecked will consume an entire blockwithin 1 h, and spread to six adjacent blocks at conflagrationproportions in 2 h (Blackburn, personal communication,2009{). San Francisco is a location where infrastructure andnatural hazards converge. The presence of frequent highwinds, fragile infrastructure and soils susceptible to liquefac-tion provides lessons from the 1906 San Francisco and 1989Loma Prieta earthquakes important for infrastructure andgeohazards worldwide.

1906 San Francisco earthquakeThe fire following the 1906 San Francisco earthquake is

the largest single fire loss in US history (Scawthorn et al.,2006a), with approximately 490 city blocks burned to theground, an additional 32 blocks severely damaged, and28 000 buildings destroyed. The earthquake had a dramaticeffect on civil infrastructure, including damage to buildings

and lifelines caused by permanent ground deformation(PGD). Investigations by Lawson et al. (1908), for example,identified areas of fill and ‘made’ ground as zones of PGDand concentrated earthquake damage. The locations of waterdistribution pipeline damage were combined with observa-tions of large ground deformation to delineate zones of‘infirm ground’ (Schussler, 1906; Manson, 1908) as a basisfor planning future improvements in the water supply sys-tem.

It is now known that much of the observed grounddeformation was caused by soil liquefaction. Fig. 21(a)shows the zones of soil liquefaction in San Franciscomapped by Youd & Hoose (1978) from their study ofhistoric ground failures in Northern California. The air photoused by Youd & Hoose has been modified to show theMarina, which was not shown in the original work. TheMarina district was sparsely populated in 1906. Althoughpermanent ground movements were observed there (e.g.Lawson et al., 1908), their effects on infrastructure wererelatively minor. Fig. 21(b) shows a rendering of the zonesof liquefaction-induced ground deformation effects devel-oped with a geographical information system (GIS).

Figure 22 is a map of the 1906 water supply within thecity limits, developed from the oldest extant maps of theSan Francisco water distribution system, reservoirs, andpressure districts (O’Rourke et al., 1992, 2006). There werenine reservoirs and storage tanks, with 92% of the totalwater storage contained in the Lake Honda, College Hill andUniversity Mound reservoirs. The remaining reservoirs werelow-volume units, which were drained within hours of theearthquake through ruptured and leaking water mains.

All trunk pipelines, 400 mm or larger in diameter, areplotted relative to the zones of liquefaction-induced grounddeformation delineated by Youd & Hoose (1978) andO’Rourke et al. (1992). Approximately 78% and 22% of the

‡ Personal communication with former Deputy Fire Chief of San Francisco Fire Department in command of firefighters during 1989 LomaPrieta earthquake.

0 1 km N

Zones of soil liquefaction Zones of soil liquefaction

Marina

Embarcadero

Approx. boundary offoot of Market Zone

James Lick Skyway

Approx. boundary ofSouth of Market Zone

Old Mission Bay

U. S. Post Office

Market St Dore St

Valencia St

18th St Mission St

Approx. boundary ofMission Creek Zone

Marina

(a) (b)

Fig. 21. Zones of historic liquefaction in San Francisco: (a) liquefaction zones (after Youd & Hoose, 1978);(b) liquefaction zones in GIS

518 O’ROURKE

trunk lines were composed of cast and wrought iron pipe,respectively. Breaks in the trunk lines crossing the liquefac-tion zones are plotted from the records of Schussler (1906)and Manson (1908). Multiple ruptures of pipelines from theCollege Hill and University Mound reservoirs occurred inthe zones of liquefaction-induced ground deformation, cut-ting off nearly 60% of stored water to the city centralbusiness district (CBD).

The ground deformation on Valencia St (see map locationlabelled ‘13 breaks’) was perhaps the single most devastatingevent of the 1906 earthquake. It was responsible for thecollapse of the Valencia St Hotel, killing 37 people, whichwas the second highest death count at an individual locationduring the earthquake (Fradkin, 2005). It also ruptured twocast iron water trunk lines, 400 and 550 mm in diameter.These broken pipelines emptied the College Hill Reservoirof 53 million litres, thereby depriving firefighters of waterfor the burning CBD of San Francisco. The ground deforma-tion at this location also destroyed a brick sewer, gas mains,electric and telephone conduits, and cable car tracks.

Figure 23 is a map of the area that burned after theearthquake, relative to the water supply depicted in Fig. 22.The trunk lines of the College Hill and University Moundreservoirs north of the pipeline breaks are removed from thisfigure to demonstrate the lack of hydraulic capacity into thecentral city. The southernmost extent of the fire was 20thStreet, where water was still available from the intact pipe-line network south of the liquefaction zone. The maximumwestern extent of the fire coincides approximately with Van

Ness Avenue. Using this wide street as a firebreak, anddrawing on water still available from Lake Honda, fire-fighters were able to stop the western advance of theconflagration.

The computer program EPANET was used to model theflow in the undamaged trunk lines from Lake Honda.EPANET is a Windows-based software program distributedby the Environmental Protection Agency (EPA, 2008) for thesimulation and analysis of flow and pressures in hydraulicdistribution networks. In essence, the program solves a seriesof non-linear equations for continuity and energy conserva-tion of incompressible flow in pressurised pipeline networks(Armando, 1987). Various parameters required to model theflow, including diameter, pipeline length, friction coefficient(C-value), and changes in elevation within the system weretaken from existing maps and historic sources of information(e.g. Schussler, 1906; Manson, 1908), as well as digitalelevation models for San Francisco.

Figure 24 presents a bar chart showing the progressiveloss of water in the Lake Honda Reservoir over four daysfrom 7 am on the day of the main shock. The simulatedreduction in reservoir level compares favourably with thesuccessively lower levels actually measured and reported bySchussler (1906). The hydraulic network analyses show thattotal water flow from Lake Honda along Van Ness Avenuewas approximately 20 000–30 000 l/min, which is sufficientfor effective action at a large firebreak such as Van NessAvenue. The analyses also show that flows from Lake Hondawere inadequate for fire suppression within the burning

Transmission pipelines

Trunk pipelines( 400 mm dia.)�

Zone of liquefaction-induced deformation

Pipeline break

Golden Gate

PresidioMilitaryReservation

PresidioHeightsReservoir

Golden Gate Park

N

Lake HondaReservoir

Scale

0 1000 m

Pac

ific

Oce

an

Lake Merced

Pila

rcito

sC

ondu

it

San

Andr

eas

Con

duit

Francisco St Reservoir

Lombard Reservoir

Clay StTank

Mar

ket S

t

13 breaks

ClarendonHeightsReservoir

Val

enci

aS

t

College HillReservoir

University MoundReservoir

Cry

stal

Spr

ings

Con

duit

8 Breaks

9 breaks

PotreroHeightsReservoir

Bay

ofS

anFrancisco

Fig. 22. Reservoirs, key pipelines and zones of liquefaction-induced ground deformation during 1906 SanFrancisco earthquake (O’Rourke et al., 1992)


CBD, especially given the loss of supply from the CollegeHill and University Mound reservoirs.

The 1906 earthquake underscores the critical importanceof liquefaction hazards in San Francisco, the interactionbetween liquefaction-induced ground deformation and waterdistribution system performance, and the effects of waterlosses on catastrophic fire and the destruction of the builtenvironment. There is perhaps no better illustration ofgeohazard effects on critical infrastructure, nor a betterdemonstration of the importance of geotechnical factors forthe design and management of large, geographically distrib-uted systems.

As described previously, highly ductile HDPE pipelineshave been tested and evaluated for ground deformation atlevels comparable to those sustained at Valencia Street andelsewhere in the 1906 zones of liquefaction-induced soilmovement. Today, many cast iron pipelines are located inzones of potentially damaging liquefaction in cities at riskfrom earthquakes. In San Francisco, many of these zones arewell known from previous earthquakes, and have beenmapped to a high degree of accuracy with respect to location(e.g. O’Rourke & Pease, 1997; O’Rourke et al., 2006).Research and case history analysis show that the replacementof existing brittle pipelines (especially trunk lines) with well-constructed HDPE or steel pipelines can substantially im-prove water supply performance during an earthquake. Theopportunity for system improvement is especially keen inSan Francisco, where critical water supply pipelines can beidentified at known liquefaction hazards, and replacementsprioritised for significant impact on system performance.

Transmission pipelines

Trunk pipelines( 400 mm dia.)�

Zone of liquefaction-induced deformation

Golden Gate

PresidioMilitaryReservation

PresidioHeightsReservoir

Golden Gate Park

N

Lake HondaReservoir

Scale

0 1000 m

Pac

ific

Oce

an

Lake Merced

Pila

rcito

sC

ondu

it

San

Andr

eas

Con

duit

Francisco St Reservoir

Lombard Reservoir

Clay StTank

ClarendonHeightsReservoir

College HillReservoir

University MoundReservoir

Cry

stal

Spr

ings

Con

duit

PotreroHeightsReservoir

Bay

ofS

anFrancisco20th St

Burnt sectionsof San Francisco

Fig. 23. Reservoirs and key pipelines in relation to the fire-damaged area of San Francisco in 1906 (O’Rourke etal., 1992)

Date and time

0

20

40

60

80

100

120

Mill

ion

litre

s of

wa

ter

Observation

Simulation

7 am18 Apr1906

7 am19 Apr1906

7 am20 Apr1906

7 am21 Apr1906

7 am22 Apr1906

Fig. 24. Comparison of simulated and observed water levels inthe Lake Honda reservoir after the 1906 San Franciscoearthquake

520 O’ROURKE

1989 Loma Prieta earthquakeAfter the earthquake and fire of 1906 a supplementary

water distribution network, known as the Auxiliary WaterSupply System (AWSS), was constructed in San Francisco toprovide emergency fire protection. The AWSS has no build-ing connections or service lines; only fire hydrants can drawfrom the network. It is intended to augment the city’spotable water distribution system, referred to as the Munici-pal Water Supply System (MWSS), by providing a systemthat works independently of, but in parallel with, the munici-pal supply.

The AWSS is described elsewhere (e.g. O’Rourke &Pease, 1992; Scawthorn et al., 2006a), and only key featuresof the system are treated herein. At the time of the LomaPrieta earthquake it comprised 200 km of cast and ductileiron pipelines, with nominal diameters of 250–500 mm.About 80% of the system was composed of cast ironpipelines.

Earthquake damage was relatively low throughout theMWSS, with the exception of the Marina, where there were123 repairs, caused principally by liquefaction-inducedground deformation (O’Rourke & Pease, 1992). Damage tothe AWSS was light, but had serious consequences. Waterflow through a 300 mm pipeline rupture at a location of soilliquefaction, supplemented by losses at broken hydrants,emptied a critical pressure-stabilising reservoir known as theJones Street Tank. Loss of this reservoir led to the loss ofwater and pressure throughout the entire CBD of the city. Inother words, the critical auxiliary supply that had been builtand maintained after the 1906 earthquake for seismic protec-tion was unable to supply water throughout large portions ofSan Francisco.

Computer simulations of the AWSS network were per-formed with a special program. GISALLE (Graphical Inter-active Serviceability Analysis for LifeLine Engineering),which had been developed as a first-generation hydraulicnetwork model to evaluate earthquake effects on large,geographically distributed water supplies (Trautmann et al.,1986; Khater et al., 1989). The computer model was builtaround a hydraulic pipeline network program that wasmodified to allow for simulation of post-earthquake damagestates. A special code was developed and incorporated in theprogram to model the hydraulic performance of damagedpipeline systems with many breaks. The program accountedfor the effects of ground deformation by modelling pipelinebreaks in zones where soils are susceptible to liquefaction.Moreover, it was checked successfully against special fireflow tests performed in the field by the San Francisco FireDepartment.

Figure 25 shows a schematic plan view of the system thatwas simulated to reproduce the conditions on the night ofthe Loma Prieta earthquake. The AWSS is divided into alower zone below an elevation of about 30 m and an upperzone above this level. The zones generally are isolated fromeach other by closed valves to prevent damage in the lowerzone from depleting water resources at the higher elevations.The Twin Peaks Reservoir and Ashbury Tank supply theupper zone, and the Jones Street Tank supplies the lowerzone. During a severe earthquake the Jones Street Tank isintended to provide elevated pressure and some flow for thelower zone, while one or both pump stations, shown in thefigure, supply water from San Francisco Bay into the lowerzone pipelines. Pump stations 1 and 2 were not activatedimmediately after the earthquake because of operator con-cerns and delay in communication, and are not included inthe simulation. The locations of damage in the AWSS areindicated in the figure and identified in the legend.

Figure 26 shows the results of the analysis in schematicformat. Open arrows denote water egress either from the

Jones Street Tank or from damaged components. The solidarrows denote internal flow. Zones of potential soil liquefac-tion are denoted in the figure as the South of Market andFoot of Market areas. These zones are the same as thoseshown at similar locations in Fig. 21. The South of Marketarea was recognised as a zone of potentially unstable ground,called ‘infirm ground’ when building the system (Manson,1908), and was isolated from adjoining portions of thenetwork by closed gate valves.

Only one open gate valve was provided for the South ofMarket area at the intersection of Market and 6th Streets, asshown by the open circle in Fig. 26. This gate valve wasdesigned to be operated remotely with utility-supplied elec-tric power. Because of electric power loss at the time ofearthquake, the valve could not be closed remotely. Conse-quently, water flowed through this gate valve until the JonesStreet Tank ran dry.

The total flow rate from the Jones Street Tank wasapproximately 78 000 l/min. Given that the normal operating

Upper zone

Lower zoneL LeakH Hydrant break

Pipe breakFB Fire boat manifold

Pump station no. 2

Pumpstation no. 1

FB FBFB

Jones StreetTank

LMarinaleak

FBH

H H

H

LFolsomleak

AshburyTank

FBTwin PeaksReservoir

Fig. 25. Schematic representation of the Auxiliary Water SupplySystem in San Francisco, and damage related to the 1989 LomaPrieta earthquake (O’Rourke & Pease, 1992)

Jones StreetTank

78000

Marina leak:12000

2900010000

1300014000

7000

South ofMarket

Foot ofMarket

Isolation valve6th and Market

38000

15000

19000

4000

10000

Folsom leak: 8000

PipelineHydrant break

Pipeline break

Zone of potential soil liquefaction

Outflow: litres/minInternal flow: litres/min

Fig. 26. Hydraulic network analytical results for the AuxiliaryWater Supply System during the 1989 Loma Prieta earthquake(O’Rourke & Pease, 1992)


capacity of the Jones Street Tank is approximately 2.72million litres, the time required to empty the Jones StreetTank would have been about 35 min. The estimated time toloss of tank agrees with observations during the earthquake.

The computer simulation package also was used in plan-ning studies before 1989 to predict correctly that the AWSSwould not be reliable in a future severe earthquake. Theresults of coupled water supply and fire simulations werepresented to the San Francisco mayor and other city offi-cials, and a bond measure was developed in 1987 for $46million, which passed with 89% voter approval, to providefunding for rehabilitation of the AWSS and other fire-relatedinfrastructure (Scawthorn et al., 2006a).

The bond approval provided the stimulus for developing athird system, called the Portable Water Supply System(PWSS), which consists of special vehicles, called hosetenders, carrying about 1.5 km of 125 mm diameter hosesand above-ground hydrants. The most serious fire after theLoma Prieta earthquake broke out in the Marina district,where liquefaction-induced ground deformation severelydamaged the MWSS, cutting off alternate sources of water.Although the Marina fire could not be controlled with theconventional components of the water supply, it was broughtunder control by the PWSS. Three hose tenders weredispatched to the Marina, where they were able to containand extinguish the fire using water pumped from SanFrancisco Bay by the city fireboat.

The 1989 Loma Prieta earthquake once again shows theimportance of liquefaction and permanent ground deforma-tion for the fate of underground lifeline systems. Because ofliquefaction, San Francisco came dangerously close to asecond earthquake-related conflagration. Fortunately, the ef-fects of liquefaction-induced ground movement on systemperformance were anticipated. They were studied by meansof computer simulations that led to a bond measure toimprove the fire department infrastructure, and a portablewater supply that is able to draw from the inexhaustiblesource of San Francisco Bay.

The earthquake demonstrates the interdependencies be-tween lifeline systems. Failure of the electric power supplycontributed to failure of the water supply because of inter-operational links. The inability of the remotely controlledisolation valve of the AWSS to be activated by utility-supplied electricity emphasises the importance of identifyingcritical interfaces between systems and providing suitableback-up for emergency conditions. The isolation valve isnow connected to a battery pack that can be activated byradio transmission.

The earthquake provides a good illustration of the benefitsof computer simulations of large, geographically distributedsystems affected by geohazards. The AWSS computer simu-lations showed the spatial consequences of damage, provid-ing the fire department and city officials with the ability tovisualise system-wide behaviour and take steps to improveperformance before the next earthquake.

LOS ANGELESThe City of Los Angeles water distribution system, which

is operated by the Los Angeles Department of Water andPower (LADWP), is located in an area of approximately1200 km2. This area is roughly ten times that of mainlandSan Francisco at a size that qualifies it at the mega-scalelevel (see Table 1). The LADWP system allows for examina-tion of a large, geographically distributed system that isinfluenced by large variations of geotechnical conditions,seismic response characteristics, and types and states ofrepair of pipelines and associated facilities.

1994 Northridge earthquakeIn 1994 Los Angeles was struck by the Northridge earth-

quake, with direct losses to the built environment exceeding$40 billion (Eguchi et al., 1996). Its impact on the waterdistribution system was substantial. Over three-quarters ofthe water supply for the City of Los Angeles was disrupted.Los Angeles Department of Water and Power (LADWP) andMetropolitan Water District (MWD) trunk lines (nominalpipe diameter > 600 mm) were damaged at 74 locations, andthe LADWP distribution pipeline (nominal pipe diam-eter , 600 mm) system was repaired at 1013 locations (e.g.Jeon & O’Rourke, 2005). Next to the 1906 San Franciscoearthquake, this event caused the most extensive earthquakedamage to a US water supply system.

Given the large earthquake-affected area, which coincideswith the underground water distribution system and thewidespread damage to that system, it was decided to treatthe pipeline network as a giant strain gauge. Water distribu-tion pipelines are constructed in a broad and relativelydense, rectilinear pattern. As a minimum, each pipeline actsas a binary extensometer, which provides site-specific infor-mation as to whether it was or was not repaired. Theobjective was to collect and normalise repair data so thatseismic intensity could be visualised and relationships be-tween damage and spatial variables quantified. The earth-quake-induced damage to water pipelines and the databasedeveloped to characterise Northridge earthquake damage aredescribed elsewhere (e.g. Toprak et al., 1999; Jeon &O’Rourke, 2005; O’Rourke & Bonneau, 2007), and only keyfeatures are presented here.

The portion of the LADWP water distribution systemmost seriously affected by the Northridge earthquake isshown in Fig. 27. GIS databases for repair locations, charac-teristics of damaged pipe, and lengths of pipelines accordingto composition and size were assembled with ARC/INFOsoftware. Nearly 11 000 km of distribution lines and over1000 km of trunk lines were digitised and plotted with ageospatial precision of �10 m throughout the San FernandoValley, Santa Monica Mountains, and Los Angeles Basin.This work was accomplished before a GIS system wasdeveloped by LADWP, and all digitisation was performedfrom analogue maps of the system.

A properly calibrated strain gauge at mega scale requiresa grid of repetitive structures having reasonably consistentproperties and a damage threshold sensitive to the displace-ments being measured. Charts depicting the relative lengthsof LADWP and MWD trunk and distribution lines accordingto pipe composition are presented in Fig. 28, where it isclear that cast iron (CI) pipelines were the most pervasive,repetitive structure in the Los Angeles distribution system atthe time of the earthquake. The 7800 km of CI pipelines hadthe broadest geographic coverage, with sufficient density inall areas to qualify as an appropriate measurement grid.Moreover, CI is a brittle material subject to increased ratesof damage at tensile strains of the order of 250–500 micro-strain, which are levels of ground strain attainable duringstrong seismic ground motion. The CI pipelines are thereforesufficiently sensitive for monitoring variations in seismicdisturbance.

Figure 29 is a map of distribution pipeline repair locationsand repair rate contours for CI pipeline damage. The repairrate contours, where repair rate is defined as the number ofrepairs per kilometre of pipeline, were developed by dividingthe map into 2 km 3 2 km areas, determining the number ofCI pipeline repairs in each area, and dividing the repairs bythe distance of CI mains in that area. Contours were thendrawn from the spatial distribution of repair rates, each ofwhich was centred on its tributary area. A variety of gridswere evaluated, and the 2 km by 2 km grid was found to

522 O’ROURKE

provide a good representation of damage patterns for themap scale of the figure (Toprak et al., 1999). By choosingthe contour interval as the mean repair rate over the entirearea, only those contours exceeding the system average areshown.

The seismic intensity, calibrated against pipeline perform-ance, correlates with the number of contour lines, with the

zones of highest seismic intensity identified as the areas ofconcentrated contours. In each instance, areas of concen-trated contours correspond to zones where the geotechnicalconditions are prone either to ground failure or to amplifica-tion of strong motion. Each zone of concentrated damage islabelled in Fig. 30 according to its principal geotechnicalcharacteristics. In effect, therefore, Fig. 29 is a seismic

Foothill feeder LA aqueduct 1 and 2

5 0 5 km

Interstate highways

Main streets

Trunk lines

Distribution linesSanta Monica Bay

N

I-118

Devonshire St

Bal

boa

Blv

d

I-5

Win

netk

a A

v.

Roscoe Blvd

405

Topa

nga

Can

yon

Blv

d

Mulholland Dr.San

taM

onica

Blvd

Venice Blvd

I-10

State Hwy 134

I-210

Fig. 27. Map of Los Angeles water supply system affected by Northridge earthquake (O’Rourke & Toprak,1997)

(a)

(b)(c)

Steel56%

Concrete18%

Riveted steel14%

Ductile iron1%

Cast iron11%

100

1000

10000

Leng

th: k

m

LADWP LADWP MWD

ConcreteRiveted steelCast ironSteelAsbestos cementDuctile iron

Trunk lines : 1014 kmDistribution lines : 10750 km

Steel11%

Ductile iron4%

Cast iron76%

Asbestos9%

Fig. 28. Statistics for water trunk and distribution pipelines in the City of Los Angeles during the 1994Northridge earthquake (O’Rourke & Toprak, 1997)


hazard map for the Los Angeles region, calibrated accordingto pipeline damage during the Northridge earthquake.

Of special interest is the location of concentrated repairrate contours in the west central part of San Fernando Valley(designated in Fig. 30 as the area of soft clay deposits).Subsurface investigations reported by Holzer et al. (1999)found local deposits of soft, normally consolidated clay.Field vane shear tests disclosed clay with uncorrected vaneshear undrained strength Suvst ¼ 20–25 kPa at a depth of5 m, just below the water table. Newmark sliding blockanalyses reported by O’Rourke (1998) provide evidence thatnear-source pulses of high acceleration were responsible forsliding and lurching on the soft, normally consolidated claydeposit. The principal cause of ground displacement andpipeline damage in the West Valley area is therefore sheardeformation in the soft clay deposits. This phenomenon ofpermanent ground movement in the absence of liquefiablesoils is often referred to as lurching.

As explained by Jeon & O’Rourke (2005), records from164 strong-motion stations (screened from a larger dataset of240 records) throughout the earthquake-affected area werecollected and analysed with respect to various seismicparameters. Spatial distributions of the different seismicparameters were estimated by interpolation and superim-posed on the pipeline network and spatially distributeddatabase of pipeline damage. Using GIS software, the repairrate was calculated for areas influenced by specific seismicparameters. Correlations were then developed through re-gression procedures to obtain the most statistically signifi-cant relationships among repair rate, values of differentseismic parameters and pipeline characteristics. Fig. 31shows the CI pipeline repair rate contours superimposed onpeak ground velocity (PGV) zones, which were developedby interpolating the maximum horizontal velocities recordedat the strong motion stations. Using the GIS database, a

pipeline repair rate was calculated for each PGV zone, andcorrelations were made between the repair rate and averagePGV for each zone. As explained by O’Rourke (1998),similar correlations were investigated for pipeline damagerelative to spatially distributed peak ground acceleration,spectral acceleration and velocity, Arias intensity, modifiedMercalli intensity (MMI), and other indices of seismicresponse. By correlating damage with various seismic para-meters, regressions were developed between repair rate andmeasures of seismic intensity.

The most statistically significant correlations for bothdistribution and trunk line repair rates were found for PGV.Such correlations are important for loss estimation analyses,which are employed to assess the potential damage duringfuture earthquakes and develop corrective measures andemergency response procedures to reduce the projectedlosses (e.g. Whitman et al., 1997).

Figure 32(a) presents the linear regression that was devel-oped between CI pipeline repair rates and PGV on the basisof data from the Northridge and other US earthquakes. Fig.32(b) shows repair rate correlations for steel, CI, ductile iron(DI) and asbestos cement (AC) distribution lines. The regres-sions indicate that the highest rate of damage for a givenPGV was experienced by steel pipelines. This result seemssurprising at first, because steel pipelines are often substan-tially more ductile than CI or AC pipelines. Steel distribu-tion pipelines in Los Angeles, however, carry the highestwater pressures, and contain a subset of pipelines that areespecially vulnerable to corrosion and slip at joints withgasket connections.

Pipeline system performance and decision supportAs described by O’Rourke et al. (2008), system perform-

ance and modelling are important, for at least three main

I-5

Santa Monica Bay

Cast iron repair rate

Distribution pipe repairsCast iron

SteelOthers

Interstate highwaysMain streets

5 0 5 10 km

N

1·3

Devonshire St

I-40

5

Topa

nga

Can

yon

Blv

d

Contour interval 0·1 repairs/km�

Bal

boa

Blv

d

Roscoe Blvd

0·6Mullholland Drive

State Hwy 134

0·5I-10

Venice BlvdSan

taM

onica

Blvd

0·6

Win

netk

a A

v.

Fig. 29. Cast iron pipeline repair rate contours for Northridge earthquake (O’Rourke & Toprak, 1997)

524 O’ROURKE

I-210

I-5

1·3

Devonshire St

Topa

nga

Can

yon

Blv

d

Bal

boa

Blv

d

Roscoe Blvd


0·5I-10

Venice Blvd

Santa

Mon

icablv

d

0·6

Win

netk

a A

v.

Soft clay deposits susceptible to lateralmovement and lurching

Sands and interbedded clay/siltssusceptible to liquefaction

Steep slopes with soils and fills susceptibleto slumping and landslides

Saturated sand and clay depositssusceptible to site amplification

Contour interval 0·1 repairs/km� 5 0 5 10 km

0·4

State Hwy 134N

Fig. 30. Geotechnical characteristics of the areas of concentrated pipeline damage after the Northridgeearthquake

I-5

1·3

Devonshire St

I-40

5

Topa

nga

Can

yon

Blv

d

Bal

boa

Blv

d

Roscoe Blvd


0·5

1·5

0·4

I-10

Venice Blvd

Santa

Mon

icaBlvd

0·6

Win

netk

a A

v.

St

NSanta Monica BayCast iron repair rateInterstate highways

Main streetsPeak ground velocity: cm/s

0–1010–2020–3030–4040–5050–6060–7070–8080–9090–100100–170

Contour interval 0·1 repairs/km�

5 0 5 10 km

Fig. 31. Pipeline repair rate contours relative to Northridge earthquake peak ground velocity (O’Rourke &Toprak, 1997)


reasons. First, system performance provides the basis forplanning and engineering at a scale commensurate withearthquake or other hazards that have large, geographicallydistributed effects. Second, system performance is the logicalextension of component and/or individual pipeline response.It entails the outcome of integrated component behaviour,and for a pipeline network represents the ultimate expressionin terms of service and the consequences of soil–structureinteraction. Third, system performance provides the onlyway by which managers and engineers can gauge the scaleand regional impact of an earthquake or similar naturalhazard. System performance sets the stage for quantifyingthe regional economic consequences and community impactof an earthquake, as well as planning for emergency re-sponse and system restoration.

The current generation of hydraulic network models forlarge, geographically distributed water supplies has evolvedsufficiently that engineers and managers can use them toplan and design for complex performance under the highlyvariable and uncertain conditions associated with geohaz-ards. Simulations are run for a suite of different scenariosthat allow system personnel to visualise a wide range ofresponses for an entire system or a specific part of thatsystem. By running multiple scenarios, with and withoutmodifications of the system, engineers and managers canidentify recurrent patterns of response and develop an over-view of potential performance, helping them plan for manyeventualities and improving their ability to improvise andinnovate during an extreme event. The plan that emergesfrom any particular suite of scenarios, however, is not asimportant as the planning process itself, because as soon asa disaster unfolds, the reality of the event will diverge fromthe features of the most meticulously designed scenario.With good planning, however, emergency managers and life-line operators can improvise, and skilled improvisation en-ables emergency responders to adapt to field conditions.

A decision support system is a computer-based informa-tion and modelling system that works interactively withusers to address unstructured problems for strategic plan-

ning, management and operations (Turban, 1995). Such asystem was developed for water supplies by using LADWPas a test bed. The system is intended to plan operations,emergency response, and new system facilities and config-urations to optimise water supply performance during andafter earthquakes (O’Rourke et al., 2008). It is generic, andthe architecture of its computer programs is adaptable to anywater supply. The system works in conjunction with thepreviously described hydraulic network model, EPANET,which is available on-line from the US EnvironmentalProtection Agency (EPA, 2008), as well as a special programfor damaged network flow modelling, known as GraphicalIterative Response Analysis for Flow Following Earthquakes(GIRAFFE). Detailed information about the developmentand evaluation of GIRAFFE is provided by Bonneau &O’Rourke (2009).

The decision support system uses a hydraulic networkmodel that accounts for all 11 700 km of water trunk anddistribution pipelines and related facilities (e.g. tanks, reser-voirs, pressure regulation stations) in the LADWP system.The system also accounts for the aggregated seismic hazardin Los Angeles through an ensemble of 59 scenario earth-quakes. The 59 scenario earthquakes provide a library ofseismic scenarios, from which engineers can select specificscenarios or combinations of scenarios to assess systemperformance. The decision support system works with riskand reliability assessment tools to provide metrics of systemperformance. The computer simulations account for theinteraction of the water and electric power supplies, andmodel output can be used to evaluate the regional economicand community impacts of water losses. All system inputand output can be visualised through GIS with advancedquery logic and web-based features. The simulations aredynamic in time, and can account for loss of service astanks and local reservoirs lose water over time through leaksand breaks in pipelines.

As explained by O’Rourke et al. (2008), the LADWPdecision support system was validated through comparisonof simulation results showing system response to the 1994

2 3 4 5 6 7 8 9

10 100PGV: cm/s

(a)

PGV: cm/s(b)

2

3

4

5

6

7

89

2

3

4

0·01

0·10

Fit equation:log 1·55 log 8·15y x� �

R2 0·85�

1994 Northridge

1989 Loma Prieta

1987 Whittier Narrows

1971 San Fernando (south)

Rep

air

rate

: num

ber

ofre

pairs

/km

2 3 4 5 6 7 8 910 100

2

3

4

56789

2

3

4

56789

2

3

4

0·001

0·010

0·100

CI

DI

AC

Fit equation (DI):

log logy x �� 1·83 9·39

R2 � 0·73

Fit equation (AC):

log logy x �� 2·26 11·02

R2 � 0·71

Fit equation (CI):

log logy x �� 1·21 6·78

R2 � 0·84

Steel

Fit equation (steel):log 0·88 log 5·28y x� �

R2 0·90�

Note: log ln, natural log�

Fig. 32. Pipeline repair rate correlation with PGV for: (a) cast iron water distribution lines; (b) steel, castiron (CI), ductile iron (DI) and asbestos cement (AC) water distribution lines

526 O’ROURKE

Northridge earthquake with the actual areas of lost waterservice, as well as pre- and post-earthquake measurementsof flow documented by LADWP after the Northridge earth-quake. There is very good agreement between model resultsand LADWP records with respect to system-wide service-ability, geographic distribution of lost service, and pre- andpost-earthquake flows over time at key locations.

Pipeline damage caused by permanent ground deformation(PGD) is accounted for explicitly by locating the area of thesystem subject to large ground movements and estimatingthe damage by various methods, including expert judgement,simplified or site-specific FE models for soil–pipeline inter-action, and fragility curves for pipeline response. Pipelinedamage caused by transient ground deformation (TGD), orseismic waves, is estimated by means of regressions devel-oped from previous earthquake records. Regressions devel-oped by Jeon & O’Rourke (2005) and Wang & O’Rourke(2008), for water distribution and trunk lines respectively,are used. Because it is not possible a priori to know theexact damage locations, multiple system response analysesknown as Monte Carlo simulations are run, and the statisticsof the simulated performances are summarised. As explainedby Bonneau & O’Rourke (2009), Monte Carlo simulationsare run according to an algorithm that uses a Poissonprocess to simulate the occurrence of pipeline damage. Theuser can specify the number of Monte Carlo simulations, orallow the program to determine the number of simulationswhen convergence criteria have been met. Typically, 15

simulations were required to meet the convergence require-ments for the simulations discussed in this paper.

In addition to modelling pipeline damage, it is importantto account for the vulnerability of other facilities. Forexample, tank damage was modelled by fragility curvesdeveloped for different types of tank used by LADWP. Thesimulations incorporate fragility curves proposed byO’Rourke & So (2000) for steel tanks as well as fragilitycurves for concrete tanks used in the loss estimation pro-gram HAZUS (FEMA, 2006).

Scenario earthquakesSystem simulations were performed for the Los Angeles

water supply response to a 7.8 MW earthquake on the south-ern San Andreas fault (SAF), known as the ShakeOutscenario earthquake (USGS, 2008). This scenario was usedas part of an earthquake preparedness exercise referred asthe Great Southern California ShakeOut, which was thelargest earthquake preparedness drill in US history, with anestimated 5.5 million people participating.

Figure 33 shows the strong motions predicted by theShakeOut scenario throughout the LADWP system. Fig.33(a) shows a regional map of the Los Angeles area inwhich the LADWP system is located, relative to the SAF.Fig. 33(b) presents the spatial distribution of PGA generatedby the earthquake. The San Fernando Valley and the south-ern part of the service area are the locations of maximum

PGA

PGA: g0·00–0·100·10–0·150·15–0·200·20–0·250·25–0·300·30–0·35

PGV and fracture

PGV: cm/s

0–2020–4040–6060–8080–100100–120120–140140–160160–180180–200200–500

(a)

SAF N

N

NSAF

(c)(b)

00 10 km10 km

LADWPsystem 0 10 km

Fig. 33. (a) Location of San Andreas fault (SAF) and LADWP system; (b) PGA contours; (c) PGV contours andSAF (Romero et al., 2010)


PGA (approximately 0.3g). Fig. 33(c) presents the SAFrupture and distribution of PGV. The locations of maximumPGV correspond to deep sediment basins, which amplify theincoming ground waves to produce locally high PGV. Thereare two main locations of wave amplification in the northernpart of the San Fernando Valley and the southeast part ofthe region, where PGVs approach 200 cm/s.

Application of decision support systemThe LADWP is undertaking an extensive capital improve-

ment programme to meet the requirements of the USEnvironmental Protection Agency and California State De-partment of Health Services with respect to surface watertreatment and disinfection by-products. Significant watersystem changes are necessary to meet the requirements.System changes include the removal of Encino, Hollywoodand Lower Stone Canyon reservoirs from normal operatingservice, which places a much greater importance on the LosAngeles Reservoir and Van Norman Complex for reliablewater distribution. Fig. 34(a) shows the locations and ap-proximate water storage capacities of the Los Angeles,Encino, Lower Stone Canyon and Hollywood reservoirs. Theremoval of these reservoirs represents a loss of about 34million m3 of water from immediate use in the system.

The system response was evaluated for 15 water serviceareas, shown in Fig. 34(b). Water service areas are geo-graphic groupings of pipelines, pumps, valves, tanks, reser-voirs and demands that can be analysed individually. Fromnorth to south the water service areas are: Granada Hills(GH): Foothills (FH); Sunland-Tujunga (ST); Valley Floor A,B and C (VF A, VF B, VF C); Encino Hills (EH); SantaMonica (SM); Hollywood Hills (HH); Mount Washington(MW); Highland Park (HP); Santa Ynez (SY); Westside(WS); Central City (CC); and Harbor (H). By showing theresults for the 15 water service areas, it is possible tounderstand the spatial variability of the system performanceas expressed in terms of serviceability index (SI), which isthe percentage of post-earthquake flows relative to pre-earth-quake flows at all demand nodes within a water service area.

The system serviceability index (SSI) is the same percentagefor the entire system.

All simulations discussed here were run for the average24 h summer daily water demand. Permanent ground move-ments predicted with simplified models in the ShakeOutscenario (USGS, 2008) were used to evaluate PGD damageto water supply pipelines, using a decision process developedwith LADWP that accounts for pipe type and the predictedmagnitude of movement (Romero et al., 2010).

Figure 35 shows the flow conditions in the trunk linesystem at 0 and 24 h after the earthquake for a singlenetwork analysis representing the median results of theMonte Carlo simulations. This figure provides informationabout the spatial distribution of flows, and the way theydiminish with time. The deterioration in performance isgenerated by losses from leaking pipelines that draw downtanks and reservoirs, causing some sections of the system tolose all local sources of water. Following such a large event,it will take a considerable time to isolate and repair leakingpipelines. A period of 24 h was chosen in consultation withLADWP personnel as a representative interval for showingtime-dependent losses before significant repair and restora-tion can be initiated. The decrease in pipelines with reliablewater flow and the increase in unsatisfied demand nodes areclearly shown by comparison of the 0 and 24 h conditions.The mean SSI declines from 76% to 34% over 24 h, whichindicates that 66% of the normal water demand cannot bemet one day after the main shock.

Figure 36 provides histograms of outcomes for theLAWDP system response to the ShakeOut earthquake sce-nario with and without the disconnected reservoirs. Becauseit is not possible to know a priori where pipelines will bedamaged by ground wave effects, the simulated earthquakeperformance outcomes must be sampled until a statisticallysignificant number is acquired, as previously explained. TheSSI statistics for 24 h after the earthquake are summarisedin the figure, where the number of Monte Carlo simulationsthat contribute to a particular SSI was divided by the totalnumber of simulations to provide an approximate probabilityindex. The histograms of ‘probability’ allow one to compareperformance outcomes when the Encino, Lower Stone Can-

13·6 M m3

EncinoReservoir

14·8 M m3

N N

0 010 km 10 km

HollywoodReservoirs

5·4 M m3

Lower StoneCanyon Reservoir

13·6 M m3

GH FHST

VFCVFB

EHVFA

HHHP

MW

SM

SY WS

CC

H

LA reservoir

(a) (b)

Fig. 34. Maps of Los Angeles water supply showing: (a) reservoirs; (b) water service areas

528 O’ROURKE

yon and Hollywood reservoirs are closed and open. Themedian SSI increases by about 6% when the reservoirs areopen as opposed to closed. The shift in the probabilitydistributions can be seen clearly in the figure.

Figure 37 provides a similar display in which service areaSIs associated with the most populated areas of Los Angelesare represented. These areas include water service areas WS,CC, HP and MW, which are geographically close to theEncino, Lower Stone Canyon and Hollywood reservoirs.

Because of their large populations, these areas of LosAngeles are likely to have the greatest need for water tofight post-earthquake fires.

The probability distributions for the most populated areasare substantially different from those for the entire system.The median SI decreases from 38% with reservoirs open to21% with reservoirs closed. Perhaps the most importantfinding is that the worst-case outcomes with reservoirsclosed vary from SI of 5% to 15%. Such low levels of water

NN

(a) (b)

0 010 km 10 km

Flows and demands

Pipes

Damaged

Flow

Non-functional

Unsatisfied demands

Fig. 35. Simulated water flows and water demands, (a) immediately after and (b) 24 h later for the MW 7.8 ShakeOutearthquake scenario (Romero et al., 2010)

0 0·05 0·10 0·15 0·20 0·25 0·30 0·35 0·40 0·45 0·50 0·55 0·60Serviceability index

0

0·1

0·2

0·3

0·4

Pro

babi

lity

Reservoirs openReservoirs closed

Reservoirs closed

Reservoirs open

Mean: 0·343Median: 0·347St dev: 0·027Min: 0·280Max: 0·371

Mean: 0·417Median: 0·396St dev: 0·044Min: 0·365Max: 0·509

Fig. 36. Histograms of system serviceability indices with andwithout disconnected reservoirs for the MW 7.8 ShakeOutearthquake scenario

0 0·05 0·10 0·15 0·20 0·25 0·30 0·35 0·40 0·45 0·50 0·55 0·60Serviceability index

0

0·05

0·10

0·15

0·20

0·25

Pro

babi

lity

Reservoirs openReservoirs closed

Reservoirs closed

Reservoirs open

Mean: 0·213Median: 0·214StDev: 0·062Min: 0·062Max: 0·356

Mean: 0·383Median: 0·385StDev: 0·089Min: 0·249Max: 0·581

Fig. 37. Histograms of serviceability indices for the mostpopulated water service areas with and without disconnectedreservoirs for the MW 7.8 ShakeOut earthquake scenario


service would expose the most populated areas of LosAngeles to exceptionally high risk. With reservoirs open, theprobability distributions shift markedly to the right, such thatthe worst-case scenarios have SIs higher than the mean SIfor reservoirs closed. In some cases water is available atnearly 50% of the demand nodes after 24 h.

The results of the scenario earthquake simulations showthat opening the disconnected reservoirs immediately after aserious earthquake improves serviceability significantly inthe locations of highest population, and is thus an effectivemeans of emergency response, even though such action willrequire tap water safety notices to be issued for the entiresystem. Work, such as this, that links emergency responsewith public health decisions provides a good example ofhow system modelling integrates geotechnical effects on aregional scale with the complex performance of a criticallifeline network to make difficult decisions involving publichealth and emergency response.

Permanent versus transient ground deformation effectsThe hydraulic network model for the LADWP system

provides an opportunity to evaluate how the Los Angeleswater supply was affected by PGD and TGD during the1994 Northridge earthquake. Because Northridge earthquakedamage to the LADWP pipelines has been carefully docu-mented with respect to PGD and TGD effects (Bonneau &O’Rourke, 2009), it is possible using the LADWP hydraulicnetwork model to simulate system response to both theindependent and combined effects of PGD and TGD.

Figure 38(a) presents the results of the simulations per-formed with only PGD-induced damage 24 h after theNorthridge earthquake. This figure shows the locations ofPGD-induced damage, pipelines with and without flow, andpipelines not normally connected but available for emer-gency operations. The figure also shows PGV contoursinterpolated from 164 strong ground motion records, whichare the same as those illustrated in Fig. 31. The GIS layercontaining the ocean and coastal boundary is not shown sothat the pipeline network and strong motion patterns aremore clearly visualised. The SSI immediately after the earth-quake is about 98%, falling to nearly 89% after 24 h.

Figure 38(b) shows the results 24 h after the Northridgeearthquake for TGD-induced pipeline damage. The figurepresents the same information as in Fig. 38(a), except thatonly TGD-induced damage is presented. The SSI immedi-ately after the earthquake is 100%, falling to 75% after 24 h.The TGD-induced damage reduces the flow in the northernand southern parts of the San Fernando Valley relative to theflow for PGD-induced damage.

Figure 38(c) shows the system performance 24 h after theNorthridge earthquake for combined PGD- and TGD-induced pipeline damage. The SSI immediately after theearthquake is about 93%, falling to 74.6% after 24 h.Combining the PGD and TGD damage produces a 24 h SSIthat is only 0.4% lower than that associated with TGDdamage only. A careful examination of the simulation resultsdiscloses that the TGD damage is widespread, and affectsimportant trunk lines that are old and vulnerable to seismicwave effects. In contrast, the PGD-induced damage is loca-lised, being concentrated at a few specific areas of liquefac-tion, and lurching in the northern part of San FernandoValley. When the PGD-induced damage is added to theTGD-induced damage, the hydraulic capabilities of thedamaged system are not significantly changed. For a com-plex system such as the Los Angeles water supply thenetwork performance is highly non-linear. Adding PGD-induced damage, which produces an 11% decline in service-ability over 24 h, to TGD-induced damage does not result ina proportionate drop in system performance.

The results show a markedly different system response inLos Angeles from the system response to earthquakes in SanFrancisco. The San Francisco system is smaller, and lackssufficient dispersion and redundancy to compensate forliquefaction-induced damage to key pipelines. The fate ofthe system is truly the fate of the ground, primarily withrespect to large deformation and failures that can incapaci-tate the network. In contrast, the Los Angeles water distribu-tion system is much larger, with alternative paths for flowshould localised PGD disrupt a limited number of pipelinesand associated facilities. Moreover, there is vulnerability inLos Angeles to TGD effects, where the widespread passageof ground waves seeks out the most vulnerable parts of thesystem.

(a) (b) (c)

N N N

PGV: cm/s PGV: cm/s PGV: cm/s0 0 0160 160 160

0 0 010 km 10 km 10 km

PGD damageTGD damage TGD damage

PGD damage

Pipes with flow

Pipes with flow

Pipes with flow

Pipes with no flow Pipes with no flow Pipes with no flowPipes not connectedPipes not connected Pipes not connected

Fig. 38. Simulation results for Los Angeles water supply damage at 24 h after the Northridge earthquake: (a) PGDonly; (b) TGD only; (c) PGD and TGD

530 O’ROURKE

Scale must be remembered, of course. Localised groundfailure in Los Angeles can reduce local serviceability with-out a dramatic reduction in system serviceability. Because ofits size, however, local reduction of serviceability in LosAngeles can affect areas of the system comparable in size tothat covered by the entire water distribution network in SanFrancisco.

HURRICANE KATRINAHurricane Katrina is generally recognised as the greatest

US natural disaster. As previously indicated, its direct cost isover $100 billion (Jordan & Paulius, 2006), of which nearly$25 billion are direct property losses in New Orleans (IPET,2008a). Over 80% of New Orleans was flooded, in someplaces to depths of 5 m. It took 53 days to dewater the city.Residential property losses were approximately 78% (IPET,2008a), resulting in the displacement of hundreds of thou-sands of people, many of whom have not returned. In 2006and 2008 (one and three years after Katrina) the populationof New Orleans was only 43.5% and 64.3%, respectively, ofits pre-hurricane level (US Census Bureau, 2009). There arenearly 2000 dead and missing attributed to Hurricane Katrina.

Hurricane Katrina also had serious consequences for Gulfof Mexico oil and natural gas production. After HurricaneKatrina, 100% of oil and 94% of gas production in the Gulfof Mexico were suspended (MMS, 2006). Restoration ofproduction was interrupted 26 days later in response toHurricane Rita, which in turn resulted in the suspension of100% of oil and 80% of gas production. Hurricanes Katrinaand Rita destroyed 111 offshore platforms, with seriousdamage to 52 others, and severely damaged 169 submarinepipelines (Wisch & Ward, 2007). Predicting mudslide dam-age to submarine pipelines is an ongoing concern. Riskanalyses link mudslides with hurricane-induced wave forces,and show the geographical distribution of potential pipelineand offshore platform damage based on mudslide recurrenceintervals (Nodine et al., 2007).

Much has been written about Hurricane Katrina and itsimplications with respect to geohazards and geotechnicalengineering. The most comprehensive study of the hurricaneeffects on the New Orleans and southeast Louisiana HPSwas prepared by the Interagency Performance EvaluationTask Force (IPET, 2008a). It comprises over 7500 pages,covering geodetic vertical and water level measurements,HPS characteristics and performance during Hurricane Katri-na, characteristics of the hurricane, consequences of thedisaster, and risk and reliability analyses. A special issue ofthe ASCE Journal of Geotechnical and GeoenvironmentalEngineering (Christian, 2008) was published, with 15 paperscovering most geotechnical aspects of the hurricane. Othernotable reports include, but are not confined to, those by theNational Academy of Engineering and National ResearchCouncil (NAE/NRC, 2009), the ASCE Hurricane KatrinaExternal Review Panel (2007), and the Independent LeveeInvestigation Team (2006).

It is the intention of this paper to provide a brief summaryof Hurricane Katrina’s effects and the performance of theHPS. Only select features are covered. Reference should bemade to the publications identified above for more detailedinformation and analyses.

Hurricane protection system (HPS)As shown in Fig. 39, the HPS consists of approximately

565 km of protective structures, including earthen levees andfloodwalls, surrounding primarily five counties, or parishes,adjacent to the Mississippi River for a total protected area of5400 km2. The system is composed of three hurricane

protection projects, identified in the figure as the LakePontchartrain and vicinity, West Bank and vicinity, and NewOrleans to Venice Projects. The levees and floodwalls pro-vided life safety protection for approximately 1.1 millionpre-Katrina residents of metropolitan New Orleans, as wellas security for one of the most important commercial andindustrial hubs in the Gulf of Mexico region. By anyinterpretation of the term the HPS is a lifeline, crucial forthe economic and societal well-being of a major port andurban centre.

The HPS was authorised by the US Congress under the1965 Flood Control Act after Hurricane Betsy caused cata-strophic flooding in New Orleans. Twenty-seven years later,after multiple design changes, environmental challenges inthe courts, and continuous interaction with local commu-nities, a final resolution was reached on how to build thesystem. When Hurricane Katrina struck New Orleans theHPS was incomplete, with no parish having the full level ofprotection authorised in 1965. As emphasised by IPET(2008a) and the NAE/NRC review (NAE/NRC, 2009), theHPS ‘did not perform as a system.’ It was constructed in apiecemeal fashion over many years that represented a historyof ‘continuous incompleteness’ (IPET, 2008a).

Figure 40 is an aerial view of the main Orleans East BankMetropolitan Basin, which is the most populous area of NewOrleans, containing the main downtown area and historicFrench Quarter. It is surrounded by a portion of the HPS,and is drained by three outfall canals (17th Street, Orleansand London Avenue canals), which convey water pumpedfrom the basin to Lake Pontchartrain, immediately north ofthe city. As part of the resolution reached with localcommunities in 1992, a parallel protection system of leveeand I-wall structures was built along the outfall canals as thesole measure to contain hurricane surge from Lake Pontchar-train. An I-wall uses a steel sheetpile driven into an existinglevee to increase the height of the flood protection structure,and thus increase its capacity to resist a higher level ofstorm surge. Fig. 40 shows the Inner Harbor NavigationalCanal (IHNC), which connects Lake Pontchartrain, the Mis-sissippi River, and a canal, which at the time of thehurricane served as a channel for both the Gulf IntercoastalWaterway and the Mississippi River Gulf Outlet. As de-scribed in detail elsewhere (e.g. IPET, 2008a; Seed et al.,2008b), that canal acted as a conduit for storm surge,elevating waterway levels and breaching the HPS at severallocations along the IHNC.

A review of the regional geology (Dunbar & Britsch,2008; IPET, 2008a; Rogers et al., 2008) discloses that muchof the soil within 10–20 m of the ground surface iscomposed of organics and normally to lightly consolidated

Lake PontchartrainLake Pontchartrain and vicinity

JeffersonNew Orleans

West Bank and vicinity

MississippiRiver

Federal HPS

State andlocal levees

LakeBorgne

St. BernardPlaquemines

New Orleans to Venice

ChandeleurSound

BaratariaBay

Gulf of Mexico

0 10

km

N

Fig. 39. Hurricane protection system for New Orleans andsouthern Louisiana (after IPET, 2008a)


clays. Subsidence triggered by drainage, atmospheric expo-sure of organics to decomposition, groundwater withdrawaland placement of fills altered the datum of many city bench-marks, with the result that some floodwalls were constructedapproximately 0.6 m below the levels intended by design.

The HPS was designed for a standard project hurricanethat was developed from meteorological data acquired from1900 to 1956. This design approach underestimated thestorm surge generated by Hurricane Katrina, which wasmeasured as 5–6 m on the eastern side of New Orleanscompared with the 3.7–4.4 m design levels associated withthis area. The HPS had not been designed for protectionagainst overtopping, with no armouring of the back sides ofthe levees. Moreover, some levees were constructed witherodible soils that were dredged from local sources, thusleading to substantial erosion and breaches in the system.

According to IPET (2008a) there were 50 major breachesin the HPS, four of which are attributed to foundation-induced failures. Three of these failures occurred along theoutfall canals to Lake Pontchartrain (one at the 17th StreetCanal and two at the London Avenue Canal), with theremaining one located on the east side of the IHNC near itsintersection with the east–west shipping channel. There isdisagreement between IPET (2008a) and the IndependentLevee Investigation Team (2006), which claims that a fifthbreach, further south along the east side of the IHNC, wasmost likely caused by loss of soil strength due to under-seepage and consequent lateral instability (Seed et al.,2008b). Both IPET and the Independent Levee InvestigationTeam note that many of the breaches occurred at transitions,where weaknesses existed at connections between flood wallsand earthen levees, as well as at penetrations of the HPS bypipelines, rail lines, roads etc. (IPET, 2008a; Seed et al.,2008b).

The HPS along the outfall canals to Lake Pontchartrain

was designed for a safety factor of 1.3, and the soil strengthsassumed in design were higher than warranted (IPET,2008a). Attention has been drawn to the inadequacy of sucha low safety factor, as well as its inconsistency with thehigher safety factors used routinely in the design of earthdams affecting populated areas (e.g. ASCE HurricaneKatrina External Review Panel, 2007; NAE/NRC, 2009).

Of critical concern are the reasons for the foundation-induced failures of the HPS. Centrifuge studies were per-formed that demonstrate the loading mechanism involved inthe failures (Sasanakul et al., 2008; Ubilla et al., 2008). Thephotograph in Fig. 41(a) shows the results of a centrifugetest simulating the conditions associated with the 17th StreetCanal failure. The hydrostatic force associated with risingwater causes a gap to open between the sheetpile and soilon the outboard side of the I-wall. Water flows into the gap,generating hydrostatic pressure that extends to the base ofthe sheetpile, and a sliding surface develops from the toe ofthe pile into the landward side of the I-wall. Fig. 41(b)provides an illustration of this loading mechanism (IPET,2008a).

There is general agreement within the engineering com-munity that the loading mechanism described above contrib-uted to the I-wall failures at the 17th Street, London Avenueand Inner Harbor navigational canals. There is disagreement,however, with respect to soil failure on the inboard side ofthe 17th Street canal I-wall. Investigations and analyses bythe IPET team (Brandon et al., 2008; Dunbar and Britsch,2008; Duncan et al., 2008; IPET, 2008a) indicate that themost likely mechanism for soil failure was sliding in lacus-trine clay, just beneath its interface with overlying swamp/marsh deposits that were immediately below the 17th Streetcanal I-wall. Investigations and analyses by the IndependentLevee Investigation Team (2006) indicate that the mostlikely mechanism for soil failure was sliding along a thin

North London Avenue canal breach

17th Street canal breach

South London Avenuecanal breach

Orleans Canal

Inner Harbour navigationcanal breaches

Lake Pontchartrain

NewOrleansCBD

N0 10

km

Area of interest

LakeBorgne

N0 2

kmMississippi River

GulfIntercoastalWaterway

Inner HarborNavigationalCanal

Lake Pontchartrain

Fig. 40. Aerial view of main Orleans East Bank metropolitan basin (after Sasanakul et al., 2008)

532 O’ROURKE

(20–30 mm) layer of sensitive organic clay at a higherelevation in the soil profile (Seed et al., 2008c). Thesensitive clay layer was located between the swamp/marshdeposits and an underlying layer of soft clay and silt lensesthat has been interpreted as soil derived from the intermix-ing of salt and freshwater environments (Rogers et al.,2008).

The lack of consensus on the mechanism for soil failureat the 17th Street canal represents a significant gap in thelessons derived for geotechnical engineering from HurricaneKatrina. Steps have been taken to limit problems with I-walls in future hurricanes, and thus the urgency for resolvingthe mechanism of soil failure has been reduced. I-wallsalong the outfall canals have been isolated from storm surgeeffects by the construction of flood gates and pump stationsat the canal junctions with Lake Pontchartrain. Elsewhere,many I-walls have been replaced by stronger flood walls.Nonetheless, uncertainty regarding the 17th Street canalfoundation failure should be resolved. The NAE/NRC review(NAE/NRC, 2009) concluded that the IPET explanation forthe failure mechanism at the 17th Street canal ‘whileplausible, is not fully convincing, and alternative failuremechanisms should be more rigorously assessed.’ The NAE/NRC review further encouraged that consideration in thedesign of levees and floodwalls be extended to all reason-able, possible failure modes.

Repairs and strengthening of the HPS are being imple-mented in a massive construction programme led by the USArmy Corps of Engineers (USACE) to create a new hurri-cane and storm damage risk reduction system (HSDRRS) by2011. It will provide protection against a hurricane with a100 year return interval (1% annual probability of failure).The standard of a 100 year hurricane is derived from the USNational Flood Insurance Program, which sets flood insur-ance requirements for buildings below the 100 year flood

elevation (e.g. Burby, 2001). Building the HPS for a100 year flood excludes protected property from designationas a flood hazard area according to federal insurancerequirements, and is used frequently as a de facto designstandard in US floodplains.

The 100 year flood level will result in an HPS signifi-cantly more resilient than the pre-Katrina HPS. As of 2008,reconstruction had produced an HSDRRS that performed ata level substantially better than what would have beenexpected from the pre-Katrina HPS during Hurricane Gustav,which was a Category 2 hurricane generating storm surge inthe IHNC less than 1 m below Hurricane Katrina levels(Wooten et al., 2009).

Nevertheless, questions have been raised about a 100 yearflood or hurricane standard for heavily populated areas (e.g.ASCE Hurricane Katrina External Review Panel, 2007;NAE/NRC, 2009). Such a standard is low compared withrecurrence intervals adopted, for example, in the Netherlandsfor flood protection systems, which use 4000–10 000 yearreturn intervals (Hoeksema, 2006). It has been pointed outthat the use of a 100 year event is inadequate for urbanisedareas (NAE/NRC, 2009). Risk and reliability analyses devel-oped by IPET (2008a) show that a 100 year design for NewOrleans is not likely to prevent major flooding for a stormwith characteristics comparable to those of HurricaneKatrina. Hurricane return intervals of 400–1000 years havebeen recommended by an independent National Academiesreview panel as the design basis for protection against stormsurges in New Orleans (NRC, 2009).

The systemic failure of the New Orleans HPS demon-strates the need for independent peer review during planningand design. Independent, external peer review for large,complex projects has been advocated by National Academiesreview panels (NRC, 2002; NAE/NRC, 2009). Not only doesindependent peer review provide fresh thinking about inter-disciplinary issues, it also allows for expert opinion aboutproblems that may be politically sensitive for local planningand design teams (NAE/NRC, 2009). It is one of the mosteffective means of identifying and managing risk, especiallyfor projects involving large, geographically distributed sys-tems.

Risk and reliability analysesOne of the most significant accomplishments of the IPET

work is the development of a risk and reliability approachfor hurricane effects in the New Orleans metropolitan area.Details are provided by IPET (2008a), and only the salientfeatures are described here.

In essence, the methodology involves the selection of asuite of scenario hurricanes from a joint frequency distribu-tion, based on historical data, which cover intensity, size,speed, track, landfall location and shape. A suite of 76scenario hurricanes, each with an annual frequency ofoccurrence, was chosen to represent the aggregate hurricanehazard in New Orleans, and physics-based computer simula-tions were performed to determine the storm surge and waveheights associated with each one. The computer simulationprocess was validated by comparison with measurements andobservations from Hurricanes Katrina and Rita.

The HPS was divided into multiple reaches, correspondingto lengths of levees and floodwalls with reasonably uniformelevation, soil and groundwater conditions, and hazard, plusfeatures such as pump stations, flood gates and transitions.A fragility curve showing the probability of failure plottedagainst water elevation was developed for each reach, transi-tion and feature. The fragility curves were defined for themass stability of the levee or floodwall and foundation soils,as well as erosion after overtopping.

Deflection and pressure

Failure and movement

Clay

Clay

Organics

Water penetrationbehind wall

Sheetpile wall

Clay

Sliding in clay

Organics

(a)

CL

(b)

CL

Fig. 41. (a) Centrifuge test cross-section and (b) loadingmechanism for New Orleans I-walls (IPET, 2008a; Sasanakulet al., 2008)


Water volumes entering various areas behind the HPS,referred to as basins and sub-basins, were calculated andlinked to the frequency of occurrence of each hurricane.Using event tree analysis, the probabilities of different levelsof flooding were estimated and weighted over the 76 scen-ario hurricanes to assess the annual probabilities of floodingthroughout the New Orleans metropolitan area. Floodheights were then related to direct property losses and lossof life.

As an example of the results of the risk and reliabilityanalyses, Fig. 42 shows flood-depth frequency maps for the50 year, 100 year and 500 year hurricanes in the mainOrleans East Bank metropolitan basin. The results are shownfor the reconstructed HPS as of 2007. Shades of greydisplay the estimated depths of flooding in 0.6 m increments.The results assume that pumps, which are available toremove water from the basin, are not working, as was thecase during Hurricane Katrina.

Figure 43 shows the direct property losses resulting fromthe predicted flooding for a 100 year hurricane for both thepre-Katrina and 2007 HPS, assuming no pumping. Theproperty losses for each sub-basin are shade coded, withblack and grey denoting the most severe losses. The 2007improvements for no pumping have a relatively minor effecton property losses, amounting to about a 7% reduction inlosses compared with those associated with the pre-KatrinaHPS. The 2007 HPS, however, is shown to provide addi-tional protection when all available pumps are working.Reductions in property losses for the 2007 HPS are approxi-mately 17% lower than those for the pre-Katrina HPS withall pumps working. The reliability and risk analysis shows asubstantial reduction in property losses of 35% when all asopposed to no pumps are operating for both the pre-Katrinaand 2007 HPS.

Interactive maps have been provided by IPET (2008b) at awebsite that can be accessed by the public to make risk-informed decisions about hurricane flooding potentialthroughout the New Orleans metropolitan area. Maps areprovided for each sub-basin that show flood depths in 0.6 mincrements associated with 50 year, 100 year and 500 yearhurricanes. A person can view the flood depths in any sub-

basin for any of the three recurrence intervals, and comparethe level of protection for the pre-Katrina, 2007 and 100 yearHPS.

Lifeline interdependencesFigure 44 shows the restoration of electric power custo-

mers in the New Orleans and southern Louisiana area as a

2007, 50 year

2007, 100 year

2007, 500 year

Feet of flooding(1 m 3·28 ft)�

�8

6–8

4–6

2–4

0–2

Fig. 42. Examples of flood depth frequency maps for 50 year, 100 year and 500 year hurricanes in New Orleansmain basin with no pumping (IPET, 2008a)

10%10–30%

30–50%50–70%

70–90%� 90%

(a)

(b)

Fig. 43. Property loss risk maps for 100 year hurricane and nopumping: (a) pre-Katrina; (b) 2007 (IPET, 2008a).

534 O’ROURKE

function of time after both Hurricanes Katrina and Rita.Because of extensive wind damage to electric power linesand flooding of substations, the restoration of electricityproceeded slowly. Twenty-six days after Hurricane Katrinalandfall, approximately 20% of customers were not restored,primarily because of property damage and loss of customerswho were waiting for reconstruction of their homes orbusinesses. When Hurricane Rita struck, there was againserious damage and prolonged power outages.

After Hurricane Katrina, the supply of crude oil andrefined petroleum products was interrupted because of a lossof electric power at the pump stations for three majortransmission pipelines: the Colonial, Plantation and Caplinepipelines. Fig. 45 shows a map of the Capline, whichtransports oil from the Louisiana offshore oil port to Illinoisfor refining and distribution throughout the US Midwest.The Capline is a nominal 1050 mm diameter high-pressurecrude oil line, and is one of the largest outside the Trans-Alaska Pipeline System. As a result of Katrina, major linesof crude oil and refined products were not available fordelivery to southern and eastern states, and gasoline anddiesel production in the Midwest was seriously affected bylack of supply. About 1.4 million barrels per day of thecrude oil supply were lost, accounting for 90% of theproduction in the Gulf of Mexico, and nearly 160 millionlitres per day of gasoline production was lost, accounting forapproximately 10% of the US supply (O’Rourke, 2007). Thethree major pipelines were not fully restored until 17 daysafter Katrina made landfall.

Hurricane Katrina caused the loss of a national liquid fueldelivery system because of electric power failure. This lossis similar to the inability to supply water for firefightingbecause of electric power failure during the 1989 LomaPrieta earthquake, discussed previously. Both these instancesillustrate the importance of interdependences among lifelinesystems, with electric power being central for inter-opera-tional control of water distribution in San Francisco andliquid fuel transmission in the central US. Especially note-worthy is the scale of fuel supply affected by HurricaneKatrina, which covers thousands of kilometres of pipelines.One of the primary lessons learned from Hurricane Katrinais the vulnerability of critical energy delivery systems as aconsequence of hurricanes. At the time of Hurricane Katrina,Gulf of Mexico infrastructure provided over 60% of portfacilities for US oil imports, nearly 47% of US refining

capacity, and approximately 19% and 29% of US natural gasand oil production respectively (US Energy InformationAdministration, 2005).

MISSISSIPPI RIVER AND GULF OF MEXICOGEOSYSTEMS

Figure 46 shows a cross-section through New Orleansnear its central business district. The cross-section is shownat exaggerated scale to emphasise the basin characteristics ofthe city. The approximate pre-Katrina flood design water

2 4 6 8 10 12 14 16 18 20 22 24 26 28 30 32 34 36 38 40 42 44 46 48

Restoration days

0

200000

400000

600000

800000

1000000

1200000

Cus

tom

ers

with

out s

ervi

ce

Reconstruction customers

Return to service customers

LF

Katrina landfall

Rita landfall

Fig. 44. Restoration of electric power customers in area affected by Hurricanes Katrina andRita (after Entergy Services, Inc., 2006)

N

Louisianaoffshoreoil port

PortFourchon

Capline

Transmissionpipelines

0 500

km

Patokaterminal

Fig. 45. Map of Capline and associated liquid fuel transmissionpipelines


levels in metres according to the National Geodetic VerticalDatum are shown for the Mississippi River and LakePontchartrain. Parts of the city are as much as 5–6 m belowhurricane surge and river flood levels. Moreover, subsidencein the New Orleans metropolitan area varies from 6 to17 mm/year, with about 10 mm/year in Orleans Parish nearLake Pontchartrain (IPET, 2008a), further compounding thedifferential between high water and the basin ground sur-face.

New Orleans is located between a proverbial rock and ahard place, with the rock being the river and the hard placebeing the sea. It is therefore instructive to examine moreclosely the Mississippi River and Gulf of Mexico geosys-tems, which set the ultimate boundary conditions for NewOrleans and the entire Mississippi delta.

Mississippi River and deltaThe Mississippi River system (including the Missouri

River and its tributaries) drains 41% of the US and parts ofCanada. As illustrated in Fig. 47, the total drainage areacovers 3 227 000 km2 (Kaufman, 1978). The official peakflow measured by the USACE in 1927 was 72 000 m3/s,although maximum flow measured by others is reported tobe as high as 85 000 m3/s (Barry, 1998). To put this inperspective, peak flow of the Mississippi River in oneminute exceeds the maximum daily New York City watersupply of 4 billion litres. The river carries not only waterbut also sediment, crucial for sustaining the delta andmaintaining the wetlands that act as a buffer against hurri-cane surge. The sediment load in the Mississippi River justnorth of the Atchafalaya River is approximately 228million t/year (Keown et al., 1981). This volume of sedimentis enough to cover 150 km2 to a depth of about 1 m.

The Mississippi River delta covers an area of about25 000 km2 (Day et al., 2007), referred to as the Mississippideltaic plain (MDP). The MDP is in dynamic equilibriumwith the Gulf of Mexico through the complex interaction ofsedimentation, vegetation, erosion, and human development.

The MDP formed about 6000–7000 years ago, and consistsof a series overlapping deltaic lobes that were deposited inresponse to changes in the main river channel location. Fig.48 presents a satellite image of the MDP developed by Dayet al. (2007). Each lobe is numbered in the figure captionwith its name and approximate time of formation beforepresent (BP). The figure shows that the Mississippi Riverhas changed course about every 700 to 1800 years. The rivercourse alterations are the natural means of restoring andmaintaining the MDP. They are punctuated, however, bycatastrophic changes in main channel location, which mayinvolve distances of several hundred km.

Figure 49 is a map of southeastern Louisiana, showing theAtchafalaya and Mississippi Rivers. The Atchafalaya Riveris the fifth largest river in the US, as measured by discharge(Reuss, 2004). Its average gradient is over two times steeperthan that of the Mississippi, and if unchecked would erodeinto the main channel of the Mississippi River and provide ahydraulically more efficient path to the sea. Between 1900and 1950 the amount of the Mississippi flow into theAtchafalaya increased from 12% to 30% (Reuss, 2004). By1954 the Atchafalaya River had captured about one third ofthe flow of the Mississippi River. At that time the USCongress authorised the USACE to build a protective struc-ture to regulate flow and sediment transfer between the tworivers.

In 1963 the Old River control structure was commissionedby the USACE. Its approximate location at the junction

LakePontchartrain

LakeBorgne

GentillyRidge

A'

A

(a)

Gentilly Ridge

5·5 m Mississippiflood level

Floodwall alongMississippi River

Hurricane protectionlevee and floodwall

5·3 mHurricanesurge level 3·5 m

7 m

St Louis Cathedral

Interstate 10

Dillard University

LakePontchartrain

Mis

siss

ipi R

iver

p

�6 m

�3 m

0 m

3 m

6 m

9 m

(b)

Fig. 46. New Orleans cross-section: (a) aerial view; (b) cross-section A–A9

Drainage areaof Mississippiand MissouriRivers

Mississippidelta

New Orleans0 500

km

N

Fig. 47. Mississippi River drainage system

NewOrleans

0 25km

N

Mississippi‘bird foot

deltaGulf of Mexico

AtchafalayaBay

Mississippi River

1

2 4

5

3

6

Fig. 48. Mississippi River deltaic plain showing overlappingdepositional lobes caused by different river channels. Times ofdeposition: (1) 4600 years BP; (2) 3500–2800 years BP; (3)2800–1000 years BP; (4) 1000–300 years BP; (5) 750–500 yearsBP; (6) 550 years BP to current times (after Day et al., 2007)

536 O’ROURKE

between the Mississippi and Atchafalaya Rivers is shown inFig. 49, and an expanded view of it is shown in Fig. 50. Asindicated in the figure, the Red River joins the Atchafalayaat the Old River, which is a remnant meander of theMississippi. When originally constructed, the Old Riverstructure consisted of a low sill structure, which acts as aweir to control flow into the Atchafalaya at any stage of theMississippi River. An overbank structure was also con-structed to divert water during flood stages of the Mississip-pi. Water is conveyed to the Atchafalaya through anoverflow channel. The project also involved construction ofa navigation lock, which is shown in the figure as the OldRiver lock.

In the flood of 1973 the Old River structure nearly failed.Severe erosion around the low sill structure destroyed itssouthern wing wall and scoured the soil surrounding thepile-supported main portion of the facility. There is surpris-ingly little written in the technical literature about the per-formance of this facility, although it is graphically describedin journalistic accounts (e.g. McPhee, 1989). Water scoureda 16 m deep hole on the Mississippi side of the low sillstructure, and threatened to enlarge a hole approximating thesize of a football stadium that had developed on theAtchafalaya side during the previous decade (Reuss, 2004).

The foundation system for the lower sill structure wasrestored by dumping riprap into the scour zones, grouting,and changing the maximum difference in head from 11.3 to6.7 m between the Mississippi and Atchafalaya sides of thesill. An additional facility, known as the auxiliary structure,was constructed and commissioned in 1986 to supplementthe control system. This structure is shown in Fig. 50 justsouth of the low sill structure.

Failure of the Old River structure in 1973 would have hadserious economic and social consequences. Some of theseconsequences have been summarised by Kazmann &Johnson (1980): they include catastrophic flooding of theAtchafalaya River basin, failures of bridges and associatedeffects on regional highway and rail transportation, failure ofpipelines and associated effects on energy transmissionsystems, and loss of fresh water supply for drinking andindustrial purposes along the lower Mississippi River. Themost important effect is loss of the main river course toNew Orleans and Baton Rouge and the attendant impact onthe commerce, industry and social fabric of southeasternLouisiana.

Gulf of Mexico and Louisiana coast protectionHistorically, protection against Mississippi floods and hur-

ricane surge has been a vicious circle. The approach tocontrolling the river has been to build increasingly largerlevees, and close outlets to distributaries. This in turn hasrestricted the distribution of sediment, leaving wetlands andbarrier islands without soil to counteract storm and hurricaneerosion. Loss of wetlands and barrier islands has increasedexposure to storm surge and hurricane flooding.

From 1875 to 1927 Mississippi River flood protection wasdominated by a policy of constructing higher levees andclosing flow to distributaries (Barry, 1998). By confiningflow and sediment transport to the main channel of the river,the natural mechanisms for delta restoration by flow throughdistributaries and overbank flooding were blocked, with theresulting loss of most sediment to the deep waters of theGulf of Mexico. In the lower Mississippi Valley alone thereare 2560 km of levees along the Mississippi River(Kaufman, 1978). Sediment transport is still confined pre-dominantly to the main river channel. Moreover, the con-struction of several large dams on the Missouri River andthe construction of the Missouri River Bank Stabilizationand Navigation Project have greatly reduced sediment deliv-eries to the Mississippi (NRC, 2009). From 1964 to 1981the measured sediment load in the lower Mississippi Riverdecreased by nearly 50%, from 388 million to 228 million t/year (Keown et al., 1981).

On the Gulf side, New Orleans and the Mississippi deltaare at increasing risk from storms and hurricanes. In 2005there were 28 major named storms in the North AtlanticOcean (NAO), the largest number on record. Fourteen ofthese storms were hurricanes, of which seven were majorhurricanes with maximum sustained wind speed of at least178 km/h (e.g. Rauber et al., 2005). The threat of storms iscompounded by a rising sea level. The Intergovernmental

Louisiana Old Rivercontrolstructure

Mississippi

BatonRougeAtchafalaya

River

20 m

Gulf of Mexico

0 50

N

km100 m

NewOrleans

Mississippi River

Fig. 49. Mississippi River, Atchafalaya River and Mississippidelta

LOUISIANA

MISSISSIPPI

LOUISIANA

Red River

Sydney A. Murraypower plant Overbank

structure

Low sillstructure

Overflow channel

Auxiliary structure

Old River

Old River lock

AtchafalayaRiver

MississippiRiver

0 5

km

N

Fig. 50. Old River structure


Panel on Climate Change (IPCC, 2008) reports a globalaverage sea level rise of 1.8 mm/year from 1961 to 2003,and an average rise of 3.8 mm/year from 1993 to 2003.Moreover, subsidence in the Mississippi delta, as describedpreviously for the New Orleans area, increases the differen-tial settlement relative to sea surface. The IPCC also reportsan increase in intense tropical cyclone activity in the NAOsince about 1970. Holland & Webster (2007) show that theincrease in NAO cyclonic activity correlates with an increasein the sea surface temperature (SST), with a 100% increasein tropical cyclone and hurricane numbers corresponding toan increase of 0.78C in SST.

Loss of sediment from the Mississippi River exacerbatesthe problems associated with erosion and loss of wetlandsrelated to storms and increased sea level in the Gulf ofMexico. From 1990 to 2000 wetland loss in the MDP wasapproximately 62 km2/year, and is forecast over the next50 years under current restoration policy to be approximately1300 km2 (Barras et al., 2003; NRC, 2009).

Recognising the integrated effects of flood control, sedi-ment transport and hurricane protection, the USACE afterHurricane Katrina was authorised through public law ‘todevelop and present a full range of flood control, coastalrestoration, and hurricane protection measures to provide . . .protection for a storm surge equivalent to a Category 5hurricane’ (NRC, 2009). In response to the federal mandate,the Corps issued the Louisiana Coastal Protection andRestoration (LACPR) Final Technical Report (USACE,2009), which was developed in coordination with the Stateof Louisiana master plan for coastal protection and restora-tion (CPRA, 2007). The LACPR study developed 111 alter-native risk reduction plans for five areas of the Louisianacoast and delta system.

An independent National Academies review panel (NRC,2009) has raised concerns about the LACPR plan. Amongthe institutional issues of concern are the lack of a unifiedplan and priority projects, the piecemeal approach to projectauthorisation, the lack of a convincing plan to preventdevelopment in high-risk areas, the need for more rigorousmulti-criteria decision-making, and the absence of a compre-hensive systems approach. Among the technical issues ofconcern are the need to quantify uncertainties associatedwith wetland restoration and river diversions, a lack ofclarity about the impact of river diversions on navigation,insufficient treatment of potential structural failures, andquestions about the availability of adequate sediment torevive coastal systems.

A key assumption of the LACPR study is that the currentconfiguration of the Louisiana coastline is sustainable(USACE, 2009). The National Academies review panelraises concerns about the technical feasibility of this as-sumption, especially given the reductions in Mississippisediment load and the increasing loss of wetlands fromstorms and rising sea level. Rather than consuming financialresources on maintaining the current coastal system enmasse, the NRC committee encourages the LACPR team tofocus its plans on high-priority projects.

SummaryA review of Hurricane Katrina’s effects on the New

Orleans HPS and the Gulf of Mexico oil and natural gasproduction and transportation infrastructure reveals problemsof immense scale and complexity. The 565 km long systemof levees and floodwalls, for example, involves geotechnicalproblems of subsidence, erosion, site characterisation, andHPS loading and failure mechanisms under diverse anddifficult soil conditions. It involves storm surge and wavepredictions and their potential frequency at various locations

for risk-based loading assessments. It involves historicaltrends, institutional cultures, community values, and interdis-ciplinary communication that helps set acceptable levels ofrisk and guide the structural and non-structural measures forrisk reduction.

Hurricane Katrina also reveals issues at an even largerscale. The sustainability of New Orleans and the Mississippidelta depends on the control of floods and management ofsediment in a 3 227 000 km2 drainage area, the prevention ofcatastrophic change in the course of the Mississippi Riverthrough capture by the Atchafalaya River, the distribution ofsediment through a 2500 km2 deltaic plain to maintain wet-lands and barrier islands, the construction and managementof hurricane protection facilities that are coordinated witheffective coastal restoration, and preparations that accountfor sea level rise and the potential for increased stormfrequency and intensity in the Gulf of Mexico.

As previously discussed, the solutions to these overarchingproblems are frequently in conflict with each other. Toresolve these conflicts engineers need to focus increasinglyon trade-offs that address the physical, environmental andhuman uncertainties that are inherently part of geosystemsand infrastructure. Planning for geosystems needs to recog-nise the limitations of sustainability en masse, and makechoices about allocating limited resources, restricting high-risk developments, and moving communities from vulnerablelocations. Selective sustainability may be the best policy, inwhich high-priority projects are implemented within theexisting budget of resources to achieve well-defined econom-ic and societal benefits.

CONCLUSIONSTo address the issue of scale in geotechnical engineering

for geohazards and large, geographically distributed systems,a generalised model is proposed in which there are sixdivisions of scale, ranging from nano scale (10–9 m) at themolecular level to mega scale (106 m) at the level ofgeographic regions. The paper explores nine orders ofmagnitude of scale with respect to the modelling of large,geographically distributed systems, from the behaviour ofpartially saturated sand to the earthquake performance of thewater supplies in San Francisco and Los Angeles, and thehurricane performance of the flood protection system inNew Orleans.

A significant trend in geotechnical engineering for compo-nent behaviour has been the implementation of large-scaletesting facilities for soil–structure interaction, and the fol-lowing conclusions can be drawn from the results of large-scale split-box testing of pipeline response to earthquake-induced ground rupture,

(a) For the test sands described in this paper, the size ofthe DS apparatus has a significant effect on thedetermination of the DS peak friction angle, �9ds-p.Values of �9ds-p for the 60 mm DS apparatus were 4–68higher than those for the 100 mm DS apparatus, and 6–78 higher than those for the 300 mm DS apparatus. Thedifference in �9ds-p is related to the increased angle ofdilatancy, łp, for the 60 mm device. The 100 mmdevice, fabricated according to the design of Lings &Dietz (2004), gives the most reliable results, and isrecommended for DS testing.

(b) A procedure is presented for using DS test measure-ments to determine the plane-strain friction angle atpeak DS strength, �ps-p, and the plane-strain cohesiveintercept cps. The cohesion is directly related to theincreased suction-induced dilatancy of partially satu-rated sand relative to that of dry sand.

538 O’ROURKE

(c) There is excellent agreement between FE simulations ofthe horizontal force against displacement of buried pipeusing �ps-p and cps determined from DS tests onpartially saturated sand and the analytical results offull-scale 2D tests for a wide range of pipe depth todiameter ratios. These results provide supportingevidence for the reduction of �ps-p and cps from DSdata as presented in this paper.

(d ) The large-scale test data confirm the substantialductility of HDPE pipe and its beneficial effects inaccommodating permanent ground deformation. Themaximum measured strains of 8% for 1.2 m of strike-slip displacement were far below strain levels asso-ciated with pipe wall rupture. The maximum reductionof pipe diameter due to ovalling, however, was 12%,indicating that loss of pipe cross-sectional area is likelyto govern failure of HDPE pipes with large D/t forground rupture effects.

The evaluation of water supply response to earthquakes inSan Francisco and Los Angeles, and of the flood protectionsystem response to hurricanes in New Orleans, providesmany lessons with respect to complex system performance,and the following conclusions can be drawn with respect tosystems modelling and infrastructure protection:

(e) Computer simulations of water supply performanceduring the 1906 San Francisco earthquake underscorethe critical importance of liquefaction hazards, and theinteractions among liquefaction-induced ground defor-mation, water distribution system performance, andcatastrophic fire and destruction of the built environ-ment. Simulations of the water supply response toearthquake effects before the 1989 Loma Prieta earth-quake were instrumental in developing alternative watersupply sources and emergency procedures that wereactually used to suppress fire and prevent conflagration,even though water flow in underground pipelines hadbeen cut off by liquefaction-induced ground deforma-tion.

( f ) The systematic analysis of pipeline repair records afterthe 1994 Northridge earthquake shows the locations ofimportant seismic and geotechnical hazards. They areused to identify zones of potential ground failure notrecognised in previous explorations, and to developregressions linking damage rates and various levels ofstrong motion that are important for earthquake lossestimation studies.

(g) A hydraulic network model has been developed forearthquake effects on the Los Angeles water supply thataccounts for all 11 700 km of its water trunk anddistribution pipelines and related facilities, an ensembleof 59 scenario earthquakes, and the effects of transientand permanent ground deformation on undergroundpipelines. Simulations performed for the Los Angeleswater supply response to a 7.8 MW earthquake on thesouthern San Andreas fault show that opening reser-voirs that have been disconnected because of waterquality regulations improves serviceability significantlyin the locations of highest population, and is thus aneffective means of emergency response, even thoughsuch action will require tap water safety notices to beissued for the entire system.

(h) The effects of Hurricane Katrina on New Orleans andthe Mississippi delta demonstrate the problems resultingfrom inadequate characterisation of system componentsand the absence of an integrated assessment ofgeohazards. Substantial and important work has beenaccomplished after Hurricane Katrina in evaluating the

HPS and taking corrective measures to develop resilientflood protection. The HPS performance in response toHurricane Katrina and its reconstruction represents anindispensable case history with respect to geohazardsand large, geographically distributed systems. One ofthe most significant accomplishments of the post-Katrina work is the development of a comprehensiverisk and reliability approach for hurricane effects,which is described in this paper.

(i) Flooding of the Mississippi, hurricanes in the Gulf ofMexico, and the sustainability of the Mississippi deltaicplain and Louisiana coastline involve engineering at anunprecedented scale, not only with respect to sheergeographic size, but also with regard to the complex-ities of the physical processes, environmental systems,and human communities embedded in the river, deltaand coastal areas. Harmonising an approach andrational plan under these conditions is a work inprogress. There are several unresolved issues that arepart of this process, including lack of consensus on themechanism for soil failure at the 17th Street canal,selection of appropriate safety factor and hurricanerecurrence interval for HPS design in urbanised areas,and an integrated flood prevention plan for theMississippi River and Gulf of Mexico that involvesrealistic assumptions about sediment distribution, coast-al restoration, and protection against catastrophicchanges in the course of the Mississippi River.

Analysing the response of infrastructure to geohazardsteaches us about the interdependences among critical infra-structure systems, their inherent vulnerabilities, and im-proved methods for modelling and managing them.Evaluating the effects of geohazards on geographically dis-tributed systems teaches us about the characterisation ofnatural hazards, modelling of their spatial and temporalvariability, and methods for making risk-based decisionsunder conditions of high uncertainty. The lessons from thisreciprocal learning process are important for the life safetyand economic viability of communities worldwide.

In the future, large, geographically distributed systemswill be planned and managed with increasingly more com-plex, risk-based models. Geotechnical engineering plays animportant and indispensable role in this modelling andmanagement activity, which is inherently multidisciplinary.The degree to which the geotechnical community learnsfrom geohazards and develops the skills necessary forsolving problems on the scale of geosystems will determinethe level of leadership it will exercise in addressing some ofthe most important challenges of the next century.

ACKNOWLEDGEMENTSI gratefully acknowledge Dr David J. Henkel, the 22nd

Rankine Lecturer, and Dr Melvin I. Esrig for their exampleand encouragement as teachers at Cornell University andtheir lifelong friendship. I also thank Drs Edward J. Cordingand Ralph B. Peck, former professors at the University ofIllinois, for their teaching, exemplary professional conduct,and encouragement.

Part of the work, on which this paper is based, wassupported by the George E. Brown, Jr Network for Earth-quake Engineering Simulation (NEES) Program of theNational Science Foundation (NSF) under Grant No. CMS-0421142 and the Multidisciplinary Center for EarthquakeEngineering Research (MCEER) and NSF under awardEEC-9709471. Any opinions, findings and conclusions orrecommendations expressed in this material are those of theauthor and do not necessarily reflect the views of the NSF.


Thanks are extended to the Cornell NEES equipment siteteam, including Professor Harry Stewart, Dr Mike Palmer,and Messrs Tim Bond and Joe Chipalowsky. I thank inparticular Dr Nate Olson and Jai Jung for their help andcontributions to this paper. The support of the Los AngelesDepartment of Water and Power, including Dr Craig Davis,Mr Glenn Singley, Dr Jianping Hu, and Mr Victor Vargas, isgratefully acknowledged. Frank Blackburn, former DeputyFire Chief of the San Francisco Fire Department, and DrCharles Scawthorn provided advice regarding fire risk in SanFrancisco. Professor Kenichi Soga and Dilan Robert of theUniversity of Cambridge provided help in modelling thestrain-softening of soil for FE analyses of soil–pipelineinteraction. Dr Ed Link helped in obtaining informationabout the New Orleans HPS and details of the IPET studyused in this paper.

NOTATIONCu coefficient of uniformity

c intercept of linear regression of tops versus centres ofMohr circles from DS tests; cohesive intercept

cds total stress DS cohesive interceptcps total stress plane-strain cohesive interceptD pipe external diameter

Dmax maximum particle sizeD50 mean grain sizedFE FE element size

dxy, dx p, dxf DS test horizontal displacements at yield, peakstrength, and �crit at which full softening occurs

d� normal strain incrementdª=2 shear strain increment

E Young’s modulusF horizontal force

F9 dimensionless horizontal forceH box height in DS apparatus; thickness of DS

specimenHc pipe centreline depth

Hbk distance between pipe and closest test basin wallL horizontal dimension of DS testing device; pipe

lengthn number of test specimens

Nq dimensionless lateral force

r2, R2 coefficient of determinationSuvst undrained vane shear strength

t pipe wall thicknessum matric suctionw gravimetric water contentY lateral displacement of pipe with respect to soil

Y 9 dimensionless lateral pipe displacementtanÆ slope of linear regression of tops versus centres of

Mohr circles from DS tests� strain

� yy strain normal to DS rupture planeª shear strain

ªd, ªdry dry soil unit weight

ªpf shear strain increment beyond yield at which there is

no dilationªt total soil unit weight

ªyx, ªxy shear strain parallel and normal to the DS ruptureplane, respectively

ıx horizontal DS soil displacementıy vertical DS soil displacement� total stress� 9 effective stress�N total normal stress� 9N effective normal stress

�ps, �Nps plane-strain total normal stress� 9ps, � 9pN plane-strain effective normal stress

�vc vertical total stress at pipe centre� 9xx effective stress in direction of DS rupture plane� 9yy effective stress normal to DS rupture plane�1 major principal total stress

�3 minor principal total stress� total shear stress�9 effective shear stress�p peak total shear stress�9p peak effective shear stress�ps plane-strain total shear stress�9ps plane-strain effective shear stress

�9yx, �9xy effective shear stress parallel and normal to the DSrupture plane, respectively

�crit total stress friction angle at constant volume duringshear deformation

�9crit critical-state effective stress friction angle�ds�p total stress DS peak friction angle�9ds�p effective stress DS peak friction angle�9ps effective stress plain-strain friction angle

�ps�p total stress plane-strain friction angle at peak DSstrength

ł angle of dilatancyłp peak angle of dilatancy

REFERENCESABAQUS (2006). User’s manual. Providence, RI: ABAQUS, Inc.American Society for Testing and Materials (ASTM) (2008a).

D6938-07b: Standard test method for in-place density and watercontent of soil and soil-aggregate by nuclear methods (shallowdepth). Annual Book of ASTM Standards Vol. 4.09. West Con-shohocken, PA: ASTM International.

American Society for Testing and Materials (ASTM) (2008b).D2216-98: Standard test methods for laboratory determinationof water (moisture) content of soil and rock by mass. AnnualBook of ASTM Standards, Vol. 4.08. West Conshohocken, PA:ASTM International.

American Society for Testing and Materials (ASTM) (2008c).D6836-02: Standard test methods for determination of the soilwater characteristic curve for desorption using a hanging col-umn, pressure extractor, chilled mirror hygrometer, and/or cen-trifuge. Annual Book of ASTM Standards, Vol. 4.09. WestConshohocken, PA: ASTM International.

American Society for Testing and Materials (ASTM) (2008d).D3080-04: Standard test methods for direct shear test of soilsunder consolidated drained conditions. Annual Book of ASTMStandards, Vol. 4.08. West Conshohocken, PA: ASTM Interna-tional.

American Society of Civil Engineers (ASCE) (1984). Guidelines forthe seismic design of oil and gas pipeline systems. Reston, VA:ASCE, Committee on Gas and Liquid Fuel Lifelines.

Anastasopoulos, G., Gazetas, M. F., Bransby, M. C. R., Davies, A.& El Nahas, A. (2007). Fault rupture propagation through sand:finite-element analysis and validation through centrifuge experi-ments. J. Geotech. Geoenviron. Engng ASCE 133, No. 8, 943–958.

Armando, L. (1987). Handbook of hydraulic engineering. NewYork, NY: John Wiley & Sons.

ASCE Hurricane Katrina External Review Panel (2007). The NewOrleans hurricane protection system: What went wrong and why.Reston, VA: ASCE.

Barras, J., Beville, S., Britsch, D., Hartley, S., Hawes, S., Johnson,J., Kemp, P., Kinler, Q., Martucci, A., Porthouse, J., Reed, D.,Roy, K., Sapkota, S. & Suhayda, J. (2003). Historic andpredicted coastal Louisiana land changes: 1978–2050, USGeological Survey Open File Report 03-334. Baton Rouge, LA:US Geological Survey, National Wetlands Research Center.

Barry, J. M. (1998). Rising tide. New York, NY: Touchstone Books,Simon & Schuster.

Bolton, M. D. (1986). The strength and dilatancy of sands. Geo-technique 36, No. 1, 65–78, doi: 10.1680/geot.1986.36.1.65.

Bonneau, A. & O’Rourke, T. D. (2009). Water supply performanceduring earthquakes and extreme events, MCEER-03-0009. Buf-falo, NY: MCEER, University of New York at Buffalo.

Brandon, T. L., Wright, S. G. & Duncan, J. M. (2008). Analysis ofthe stability of I-walls with gaps between the I-wall and thelevee fill. J. Geotech. Geoenviron. Engng ASCE 134, No. 5,692–700.

Bray, J. D., Seed, R. B., Cluff, L. S. & Seed, H. B. (1994).

540 O’ROURKE

Earthquake fault rupture propagation through soil. J. Geotech.Engng ASCE 120, No. 3, 543–561.

British Standards Institution (BSI) (1990). BS 1377: Part 7: 4:Determination of shear strength by direct shear. Milton Keynes:British Standards Institution.

Burby, R. (2001). Flood insurance and floodplain management: theUS experience. Global Environmental Change Part B: Environ-mental Hazards 3, Nos 3–4, 111–122.

Cerato, A. B. & Lutenegger, A. J. (2006). Specimen size and scaleeffects of direct shear box tests on sands. Geotech. Test. J. 29,No. 6, 507–516.

Chong, K. P. & Davis, D. C. (1999). Engineering mechanics andmaterials research in the information technology age. Appl.Math. Mech. 20, No. 12, 1305–1309.

Christian, J. T. (2007). Lessons from Hurricane Katrina. The Bridge,Spring, 5–11.

Christian, J. T. (ed.) (2008). Performance of geo-systems duringHurricane Katrina (Special issue). J. Geotech. Geoenviron.Engng ASCE 134, No. 5.

Coastal Protection and Restoration Authority of Louisiana (CPRA)(2007). Louisiana’s comprehensive plan for a sustainable coast.Coastal Protection and Restoration Authority of Louisiana, Stateof Louisiana. http://www.lacpra.org/

Cosgrave, J. (2007). Synthesis report: Expanded summary. Jointevaluation of the international response to the Indian Oceantsunamis. London: Tsumani Evaluation Coalition.

Dane, J. H. & Topp, G. C. (2002). Methods of soil analysis. Part 4:Physical methods. Madison, WI: Soil Science Society of America.

Davis, E. H. (1968). Theories of plasticity and the failure of soilmasses. In Soil mechanics: selected topics (ed. I. K. Lee), pp.341–380. London: Butterworth.

Day, J. W. Jr, Boesch, D. F., Clairain, E. J., Kemp, G. P., Laska, S.B., Mitsch, W. J., Orth, K., Mashriqui, H., Reed, D. J., Shabman,L., Simenstad, C. A., Streever, B. J., Twilley, R. R., Watson, C.C., Wells, J. T. & Whigham, D. F. (2007). Restoration of theMississippi delta: lessons from Hurricanes Katrina and Rita.Science 315, No. 5819, 1679–1684.

Dietz, M. S. (2000). Developing an holistic understanding of inter-face friction using sand within the direct shear apparatus. PhDthesis, University of Bristol, UK.

Duncan, J. M., Brandon, T. L., Wright, S. G. & Vroman, N. (2008).Stability of I-walls in New Orleans during Hurricane Katrina. J.Geotech. Geoenviron. Engng ASCE 134, No. 5, 668–691.

Dunbar, J. B. & Britsch, L. D. III (2008). Geology of the NewOrleans area and the canal levee failures. J. Geotech. Geoenvir-on. Engng ASCE 134, No. 5, 566–582.

Earthquake Engineering Research Institute (EERI) (2008). TheWenchuan, Sichuan Province, China earthquake of May 12,2008, EERI Special Earthquake Report, October 2008. Oakland,CA: EERI.

Eguchi, R. T., Goltz, J. D., Taylor, C. E., Chang, S. E., Flores, P. J.,Johnson, L. A., Seligson, H. A. & Blais, N. C. (1996). TheNorthridge earthquake as an economic event: direct capitallosses, analyzing economic impacts and recovery from urbanearthquake: issue for policy makers. Proceedings of the EERIConference, Pasadena, CA, pp. 1–28.

Ellsworth, W. L. (1990). Earthquake history, 1769–1989. In TheSan Andreas Fault System, USGS Professional Paper 1515, pp.153–188. Washington, DC: US Government Printing Office.

Emanuel, K. (2005). Divine wind. New York, NY: Oxford Univer-sity Press.

Environmental Protection Agency (EPA) (2008). EPANET. Wa-shington, DC: Environmental Protection Agency (http://www.e-pa.gov/nrmrl/wswrd/dw/epanet.html).

Entergy Services, Inc. (2006). Electric power customer restorationplots provided by Utility Operations Communications Depart-ment. Little Rock, AR: Energy Services, Inc.

Federal Emergency Management Agency (FEMA) (2006). Multihazard loss estimation methodology earthquake model, HAZUS-MH MR3. Technical Manual. Washington, DC: Department ofHomeland Security, FEMA Mitigation Division (http://www.fe-ma.gov/plan/prevent/hazus/hz_manuals.shtm).

Fitzpatrick, P. J. (2006). Hurricanes. Santa Barbara, CA: ABC-CLIO, Inc.

Fradkin, P. L. (2005). The great earthquake and firestorms of 1906.Berkeley, CA: University of California Press.

Hamada, M., Isoyama, R. & Wakamatsu, K. (1995). The 1995Hyogoken-Nambu (Kobe) earthquake liquefaction, ground displa-cement and soil condition in Hanshin area. Tokyo: Associationfor Development of Earthquake Prediction, Waseda University.

Hoeksema, R. (2006). Designed for dry feet flood protection andland reclamation in the Netherlands. Reston, VA: ASCE.

Holland, G. J. & Webster, P. J. (2007). Heightened tropical cycloneactivity in the North Atlantic: Natural variability or climatetrend? Phil. Trans. R. Soc. London A 365, 2695–2716.

Holzer, T. L., Bennet, M. J., Ponti, D. J. & Tinsley, J. C. III (1999).Liquefaction and soil failure during 1994 Northridge earthquake.J. Geotech. Geoenviron. Engng ASCE 125, No. 6, 438–452.

Independent Levee Investigation Team (2006). Investigation ofthe New Orleans regional flood protection systems during hurri-cane Katrina: Final report (http://www.ce.berkeley.edu/,new_orleans/).

Interagency Performance Evaluation Task Force (IPET) (2008a).Performance evaluation of the New Orleans and southeastLouisiana hurricane protection system: Final report. US ArmyCorps of Engineers (https://ipet.wes.army.mil/).

Interagency Performance Evaluation Task Force (IPET) (2008b).Risk and reliability report. US Army Corps of Engineers (http://nolarisk.usace.army.mil/)

Intergovernmental Panel on Climate Change (IPCC) (2008). Climatechange 2007: Synthesis report. Geneva: Intergovernmental Panelon Climate Change (http://www.ipcc.ch/).

Iwan, W. D. (ed.) (2006). The great Sumatra earthquakes and IndianOcean tsunamis of 26 December 2004 and 28 March 2005.Earthquake Spectra Special Issue III, 22, 1–12.

Jeon, S.-S. & O’Rourke, T. D. (2005). Northridge earthquake effectson pipelines and residential buildings. Bull. Seismol. Soc. Am.95, No. 1, 1–25.

Jewell, R. A. & Wroth, C. P. (1987). Direct shear tests onreinforced sand. Geotechnique 37, No. 1, 53–68.

Johnston, A. C. & Schweig, E. S. (1996). The enigma of the NewMadrid earthquakes of 1811–1812. Ann. Rev. Earth Planet. Sci.24, 339–384.

Jordan, J. W. & Paulius, S. L. (2006). Lessons learned from HurricaneKatrina. Proc. Forensic Engng 2006, Reston, VA, 41–69.

Kafali, C. & Grigoriu, M. (2007). Seismic fragility analysis:application to simple linear and nonlinear systems. EarthquakeEngng Struct. Dynam. 36, No. 13, 1885–1900.

Kaufman, R. I. (1978). Control of 1973 Mississippi River flood. J.Water Res. Plan. Mngmt Div. ASCE 104, No. WR1, 105–121.

Kazmann, G. & Johnson, D. B. (1980). If the old river structurefails, Bulletin 12. Baton Rouge, LA: Louisiana Water ResourcesResearch Institute.

Keown, M. P., Dardeau, E. A. & Causey, E. M. (1981). Character-ization of the suspended-sediment regime and bed-materialgradation of the Mississippi River basin: Final report. Vicks-burg, MS: US Army Engineer Waterways Experiment Station.

Khater, M. M., Grigoriu, M. D. & O’Rourke, T. D. (1989). Service-ability measures and sensitivity factors for estimating seismicperformance of water supply systems. Proc. 9th World Conf.Earthquake Engng, Tokyo 8, 123–128.

Lawson, A. C., Gilbert, G. K., Reid, H. F., Branner, J. C., Fairbanks,H. W., Wood, H. O., Hayford, J. F., Baldwin, A. L., Omori, F.,Leuschner, A. O., Davidson, G., Matthes, F. E., Anderson, R.,Louderback, G. D., Holway, R. S., Eakle, A. S., Crandall, R.,Hoffman, G. F., Warring, G. A., Hughes, E., Rogers, F. J. &Baird, A. (1908). The California earthquake of April 18, 1906:report of the California State Earthquake Investigation Commis-sion, Publication No. 87 (two volumes and atlas). Washington,DC: Carnegie Institute.

Lings, M. L. & Dietz, M. S. (2004). An improved direct shearapparatus for sand. Geotechnique 54, No. 4, 245–256, doi:10.1680/geot.2004.54.4.245.

Longshore, D. (1998). Encyclopedia of hurricanes, typhoons, andcyclones. New York, NY: Facts On File, Inc.

Manson, M. (1908). Reports on auxiliary water supply system forfire protection for San Francisco, California. San Francisco, CA:Board of Public Works.

McPhee, J. (1989). The control of nature. New York, NY: Farrar,Straus, and Giroux.

Minerals Management Service (MMS) (2006). Impact assessment ofoffshore facilities from Hurricanes Katrina and Rita, Release


3418. Washington, DC: Minerals Management Service, USDepartment of Interior.

Mitchell, J. K. & Soga, K. (2005). Fundamentals of soil behavior,3rd edn. Hoboken, NJ: John Wiley & Sons.

National Academy of Engineering and National Research Council(NAE/NRC) (2009). The New Orleans Hurricane ProtectionSystem: Assessing pre-Katrina vulnerability and improving miti-gation and preparedness. Washington, DC: National AcademiesPress.

National Research Council (NRC) (2002). Review procedures forwater resources project planning. Washington, DC:National Academies Press.

National Research Council (NRC) (2009). Final Report from theNRC Committee on the review of the Louisiana Coastal Protec-tion and Restoration (LACPR) Program. Washington, DC: Na-tional Academies Press..

Nodine, M. C., Cheon, J. Y., Wright, S. G. & Gilbert, R. B. (2007).Mudslides during Hurricane Ivan and an assessment for thepotential for future mudslides in the Gulf of Mexico. MMSProject No. 552, Phase II Project Report prepared for MineralsManagement Service by University of Texas at Austin andOffshore Technology Research Center.

Olson, N. A. (2009). Soil performance for large scale soil-pipelinetests. PhD thesis, Cornell University, Ithaca, NY.

O’Rourke, M. J. & So, P. (2000). Seismic fragility curves for on-grade steel tanks. Earthquake Spectra 16, No. 4, 801–815.

O’Rourke, T. D. (1998). An overview of geotechnical and lifelineearthquake engineering, Geotechnical Special Publication No.75, Vol. 2, 1392–1426. Reston, VA: ASCE.

O’Rourke, T. D. (2007). Critical infrastructure, interdependencies,and resilience. The Bridge, Spring, 22–29.

O’Rourke, T. D. & Bonneau, A. (2007). Lifeline performance underextreme loading during earthquakes. In Earthquake geotechnicalengineering (ed. K. D. Pitilakis), pp. 407–432. Dordrecht:Springer.

O’Rourke, T. D. & Pease, J. W. (1992). Large ground deformationsand their effects on lifelines facilities: 1989 Loma Prieta earth-quake, NCEER-92-0002, Vol. 2, pp. 5.1–5.85. Buffalo, NY:MCEER.

O’Rourke, T. D. & Pease, J. W. (1997). Mapping liquefiable layerthickness for seismic hazard assessment. J. Geotech. EngngASCE 123, No. 1, 46–56.

O’Rourke, T. D. & Toprak, S. (1997). GIS assessment of watersupply damage from the Northridge earthquake. In Spatialanalysis in soil dynamics and earthquake engineering (ed. D.Frost), pp. 117–131. Reston, VA: ASCE.

O’Rourke, T. D., Beaujon, P. A. & Scawthorn, C. R. (1992). Largeground deformations and their effects on lifeline facilities: 1906San Francisco earthquake, NCEER-92-0002, Vol. 2, pp. 1.1–1.130. Buffalo, NY: MCEER.

O’Rourke, T. D., Stewart, H. E. & Jeon, S.-S. (2001). Geotechnicalaspects of lifeline engineering. Proc. Instn Civ. Engrs Geotech.Engng 149, No. 1, 13–26.

O’Rourke, T. D., Wang, Y. & Shi, P. (2004). Advances in lifelineearthquake engineering. Proc 13th World Conf. on EarthquakeEngng, Vancouver, paper no. 5003 (CD-ROM).

O’Rourke, T. D., Bonneau, A., Pease, J., Shi, P. & Wang, Y. (2006).Liquefaction ground failures in San Francisco. Earthquake Spec-tra 22, No. 52, 6091–6112.

O’Rourke, T. D., Jezerski, J. M., Olson, N. A., Bonneau, A. L.,Palmer, M. C., Stewart, H. E., O’Rourke, M. J. & Abdoun, T.(2008). Geotechnics of pipeline system response to earthquakes.Proc. Geotechnical Earthquake Engineering and Soil DynamicsIV, Sacramento, CA, 1–38.

Palmer, M. C., O’Rourke, T. D., Stewart, H. E., O’Rourke, M. J. &Symans, M. (2006). Large displacement soil–structure inter-action test facility for lifelines. Proc. 8th US Nat. Conf. onEarthquake Engineering, San Francisco, CA, paper no. 1257(CD-ROM).

Pham, H. Q., Fredlund, D. G. & Barbour, S. L. (2005). A study ofhysteresis models for soil water characteristic curves. Can.Geotech. J. 42, No. 6, 1548–1568.

Rauber, R. M., Walsh, J. E. & Charlevoix, D. J. (2005). Severe andhazardous weather, 2nd edn. Dubuque, IW: Kendall/Hunt Pub-lishing Co.

Reuss, M. (2004). Designing the bayous: The control of water in

the Atchafalaya Basin 1800–1995. College Station, TX: TexasA&M University Press.

Rogers, J. D., Boutwell, G. P., Schmitz, D. W., Karadeniz, D.,Watkins, C. M., Athanasopoulos-Zekkos, A. G. & Cobas-Boa,D. (2008). Geologic conditions underlying the 2005 17th Streetcanal levee failure in New Orleans. J. Geotech. Geoenviron.Engng ASCE 134, No. 5, 583–601.

Romero, N., O’Rourke, T. D., Nozick, L. K. & Davis, C. R. (2010).Seismic hazards and water supply performance. J. EarthquakeEngng (accepted for publication).

Sasanakul, I., Vanadit-Ellis, W., Sharp, M., Abdoun, T., Ubilla, J.,Steedman, S. & Stone, K. (2008). New Orleans levee systemperformance during Katrina: 17th Street Canal and OrleansCanal North. J. Geotech. Geoenviron. Engng ASCE 134, No. 5,657–667.

Scawthorn, C., O’Rourke, T. D. & Blackburn, F. B. (2006a). TheSan Francisco earthquake and fire of 1906: enduring lessons forfire protection and water safety. Earthquake Spectra 22, No. 52,5135–5158.

Scawthorn, C., Kornfield, L., Seligson, H. & Rojahn, C. (2006b).Estimated losses from scenario earthquakes affecting San Fran-cisco: CAPPS: Part 2. Proc. 8th US Conf. Earthquake Engng,San Francisco, CA, paper no. 1595 (CD-ROM).

Schofield, A. (2005). Disturbed soil properties and geotechnicaldesign. London: Thomas Telford.

Schussler, H. (1906). The water supply of San Francisco, Califor-nia. New York, NY: Martin B. Brown Press.

Seed, R. B., Bea, R. G., Abdelmalak, R. I., Athanasopoulos-Zekkos,A., Boutwell, G. P., Briaud, J.-L., Cheung, C., Cobos-Roa, D.,Ehrensing, L., Govindasamy, A. V., Harder, L. F. Jr, Inkabi, K.S., Nicks, J., Pestana, J. M., Porter, J., Rhee, K., Riemer, M. F.,Rogers, J. D., Storesund, R., Vera-Grunauer, X. & Wartman, J.(2008a). New Orleans and Hurricane Katrina I: Introduction,overview and the east flank. J. Geotech. Geoenviron. EngngASCE 134, No. 5, 701–717.

Seed, R. B., Bea, R. G., Athanasopoulos-Zekkos, A., Boutwell, G.P., Bray, J. D., Cheung, C., Cobos-Roa, D., Ehrensing, L.,Harder, L. F. Jr, Pestana, J. M., Riemer, M. F., Rogers, J. D.,Storesund, R., Vera-Grunauer, X. & Wartman, J. (2008b). NewOrleans and Hurricane Katrina II: The Central Region and theLower Ninth Ward. J. Geotech. Geoenviron. Engng ASCE 134,No. 5, 718–739.

Seed, R. B., Bea, R. G., Athanasopoulos-Zekkos, A., Boutwell, G. P.,Bray, J. D., Cheung, C., Cobos-Roa, D., Harder, L. F. Jr, Moss, R.E. S., Pestana, J. M., Riemer, M. F., Rogers, J. D., Storesund, R.,Vera-Grunauer, X. & Wartman, J. (2008c). New Orleans andHurricane Katrina III: The 17th Street Drainage Canal. J. Geo-tech. Geoenviron. Engng ASCE 134, No. 5, 740–761.

Shibuya, S., Mitachi, T. & Tamate, S. (1997). Interpretation ofdirect shear box testing of sands as quasi-simple shear. Geotech-nique 47, No. 4, 769–790, doi: 10.1680/geot.1997.47.4.769.

Shinozuka, M., Feng, M. Q., Lee, J. & Naganuma, T. (2000).Statistical analysis of fragility curves. J. Engng Mech. ASCE126, No. 12, 1224–1231.

Stone, R. (2009). A deeply scarred land. Science 324, 8 May, 713–714.

Tokimatsu, K. & Suzuki, H. (2004). Pore water pressure responsearound pile and its effects on p–y behavior during soil liquefac-tion. Soils Found. 44, No. 6, 101–110.

Toprak, S., O’Rourke, T. D. & Tutnucu, I. (1999). Geographicinformation system (GIS) characterization of spatially-distributedlifeline damage. Proc. 5th US Conf. on Lifeline EarthquakeEngng, Washington, DC, 110–119.

Trautmann, C. H. (1983). Behavior of pipe in dry sand underlateral and uplift loading. PhD thesis, Cornell University, Ithaca,NY.

Trautmann, C. H. & O’Rourke, T. D. (1985). Lateral force–displacement response of buried pipe. J. Geotech. Engng ASCE111, No. 9, 1068–1084.

Trautmann, C. H., Kulhawy, F. H. & O’Rourke, T. D. (1985). Sanddensity measurements for laboratory studies. Geotech. Test. J.ASTM 6, No. 4, 159–165.

Trautmann, C. H., O’Rourke, T. D., Grigoriu, M. D. & Khater, M.M. (1986). Systems model for water supply following earth-quakes. In Lifeline seismic risk analysis: case studies (ed. R.Eguchi), pp. 30–50. Reston, VA: ASCE.

542 O’ROURKE

Turban, E. (1995). Decision support and expert systems: Manage-ment support systems, 4th edn. Upper Saddle River, NJ: PrenticeHall.

Turner, J. E. (2004). Lateral force–displacement behavior of pipesin partially saturated sand. MS thesis, Cornell University,Ithaca, NY.

Ubilla, J., Abdoun, T., Sasanakul, I., Sharp, M., Steedman, S.,Vanadit-Ellis, W. & Zimmie, T. (2008). New Orleans leveesystem performance during Hurricane Katrina: London Avenueand Orleans Canal south. J. Geotech. Geoenviron. Engng ASCE134, No. 5, 668–680.

US Army Corps of Engineers (USACE) (2009). Louisiana coastalprotection and restoration: Final Technical Report. US ArmyCorps of Engineers, New Orleans District (http://lacpr.usace.ar-my.mil/default.aspx?p¼LACPR_Final_Technical_Report).

US Census Bureau (2009). US Census Bureau news release, 1 July(http://www.census.gov/Press-Release/www/releases/archives/population/013960.html).

US Energy Information Administration (2005). Hurricane Katrinaimpacts on the US oil and natural gas markets: Special report,31 August (http://tonto.eia.doe.gov/oog/special/eia1_katrina_083105.html/)

US Geological Survey (USGS) (2008). The ShakeOut scenario,USGS Open File Report 2008-1150, CGS Preliminary Report25. US Geological Survey, California Geological Survey (http://pubs.usgs.gov/of/2008/1150/).

Wang, Y. & O’Rourke, T. D. (2008). Seismic performance evalua-tion of water supply systems, MCEER-09-0008. Buffalo, NY:MCEER, University of New York at Buffalo.

Whitman, R. V., Anagnos, T., Kircher, C. A., Lagorio, H. J.,Lawson, R. S. & Schneider, P. (1997). Development of anational earthquake loss estimation methodology. EarthquakeSpectra 13, No. 4, 643–661.

Wisch, D. J. & Ward, E. G. (2007). The legacy of the terribletriplets and standards of the future. Proc. Offshore Technol.Conf., Houston, TX, paper OTC 18979.

Wooten, R. L., Gilbert, R. B., Marcuson, W. F., Harder, L. F. &Nicolson, P. G. (2009). Reconnaissance of the New Orleanshurricane and storm damage reduction system after HurricaneGustav, Report No. GEER-015. Berkeley, CA: GeoengineeringExtreme Events Reconnaissance Association.

Yimsiri, S., Soga, K., Yoshizaki, K., Dasari, G. & O’Rourke, T. D.(2004). Lateral and upward soil–pipeline interactions in sand fordeep embedment conditions. J. Geotech. Geoenviron. EngngASCE 130, No. 8, 830–842.

Youd, T. L. & Hoose, S. N. (1978). Historic ground failures innorthern California triggered by earthquakes, USGS Profes-sional Paper 993. Washington, DC: US Government PrintingOffice.

VOTE OF THANKSC. J. F. P. JONES, Emeritus Professor of Geotechnical

Engineering, Newcastle University

Tom O’Rourke has been a regular visitor to the UnitedKingdom for some 30 years. During his early visits hisinterests were associated primarily with the effects of tunnel-ling on structures, and the effects of mining subsidence onburied services such as gas pipelines. The detrimental effectsof mining subsidence on buried services, as well as onstrategic structures such as bridges, is well understood in thecoalfields of Britain, and Professor O’Rourke’s studies wereclearly aligned with those interested in these subjects, part-icularly in the North of England. Mining subsidence effectson buried services are an example of system componentsoperating under extreme conditions of soil–structure inter-action. In essence, mining subsidence has been described asbeing similar to a ‘slow earthquake’.

Professor O’Rourke was always a popular lecturer duringhis frequent visits, with a natural ability to capture theaudience’s attention, as he has done this evening. He notablyexplained to engineering students that geotechnical engineer-ing can be described by the three Ds: ‘dirty, difficult anddangerous’. Tom has now moved on from a study ofindividual system components, extending his experience andknowledge to the management of large geographical distri-bution systems that are affected by a range of potentialgeohazards such as earthquakes, floods and hurricanes. Thiswork could be described by the three Is: ‘intelligent; inte-grated and influential’.

Intelligence comes from a fundamental understanding ofcomplex soil–structure interaction, which offers improve-ments in our ability to model soil–structure behaviour.Integration results from the use of advanced geographicalinformation systems and remote sensing, and permits thisunderstanding to be applied to large distributed lifelinesystems. The work is influential in that it widens the rolethat geotechnical engineers have in society, whereby theynow play a critical role in managing the performance oflarge infrastructure networks that are susceptible to naturalhazards.

In his lecture Professor O’Rourke has shared his unri-valled experience of the management systems resulting fromhis studies, illustrating their effectiveness in the managementof the Los Angeles and San Francisco water distributionnetworks, and as a tool in understanding the effects ofhurricanes on New Orleans and the Gulf of Mexico oil andgas networks.

Tom O’Rourke’s lecture looks to the future of geo-technical engineering, and points the way for engineers totake a more proactive and influential role in the managementof the infrastructure systems upon which society depends.

It is with great pleasure that I invite the audience to thankProfessor O’Rourke for an elegant and influential Rankinelecture.


geot60-505 geohazards and large, geographically distributed systems

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