Tsunami loading of near-shoreline structures: a primer

Tsunami loading of near-shoreline structures:a primer

Dan Palermo, Ioan Nistor, Younes Nouri, and Andrew Cornett

Abstract: The effects and estimation of tsunami-induced loading on near-shoreline structures located within inundationzones have recently gained significant interest from researchers, engineers, and government agencies. Building codes,namely the 2005 National building code of Canada, do not explicitly consider tsunami loading, as it is understood that in-land structures can be protected by proper site planning. However, recent catastrophic events (Indian Ocean, 2004; Solo-mon Islands, 2007) indicate that tsunami loading should be considered in structural design. Presented herein is a review offorce components that arise from tsunami-induced hydraulic bores running inland, along with proposed loading combina-tions and load cases readily applicable for building codes. Test results from a comprehensive experimental program con-ducted in a large-scale flume are also provided. A comparison of experimental results with force components provided inreadily available design documents is presented, and suggestions for improvements are further discussed.

Key words: tsunami, hydrodynamic forces, surge forces, hydrostatic forces, debris impact, design codes.

Resume : Les effets et l’estimation des charges causees par les tsunamis sur les structures cotieres situees dans des zonesinondables ont recemment souleve l’interet des chercheurs, des ingenieurs et des agences gouvernementales. Les codes dubatiment, particulierement le Code national du batiment du canada 2005, ne considerent pas explicitement la charge de tsu-nami puisqu’il est entendu que les structures a l’interieur des terres peuvent etre protegees par une planification adequatedu site. Toutefois, de recents evenements catastrophiques (Ocean Indien, 2004, Iles Salomon 2007) indiquent que l’on doittenir compte de la charge de tsunami dans la conception structurale. Le present article comporte une revue des composan-tes des forces survenant lorsque des ondes a front raide generees par un tsunami passent sur l’ıle; il propose en plus descombinaisons de charge et des cas de charge qui pourraient etre inscrits dans les codes de batiment. Les resultats d’essaisd’un programme experimental complet realise dans un canal a grande echelle sont egalement fournis. Une comparaisondes resultats experimentaux avec les composantes des forces est presentee dans les documents de conception disponiblesdes maintenant et des suggestions d’amelioration sont discutees en profondeur.

Mots-cles : tsunami, forces hydrodynamiques, forces de poussees, forces hydrostatiques, impact des debris, codes deconception.

[Traduit par la Redaction]

IntroductionSeveral recent events, including the 24 December 2004,

Indian Ocean Tsunami and the 2 April 2007, South Pacific(Solomon Islands) Tsunami, have brought to the fore the im-portance of understanding the effects of tsunami-inducedloading and the need for design guidelines for structures lo-cated near shorelines in tsunami-prone regions. Canada,bounded by oceans at its western and eastern extremities,should give consideration to its infrastructure to protect

against such catastrophic events. Although Canada has beensomewhat spared from a major earthquake off the coast ofBritish Columbia in recent history, seismologists agree thatthe next ‘‘big one’’ could occur at any time. An event simi-lar to the 1700 Cascadia megathrust earthquake has the po-tential to generate a tsunami of devastating proportion. The2005 National building code of Canada (NBCC), however,remains silent, for the most part, on design for tsunami ef-fects (NRCC 2005). Commentary J of the 2005 NBCC,‘‘Design for seismic effects,’’ states that damage to buildingsas a result of an earthquake can arise from ground shaking,soil failures, surface fault ruptures, or tsunamis. However,only ground shaking and the potential for soil liquefactionare explicitly considered. Other hazards are addressedthrough planning and site selection. This may lead structuralengineers to assume that tsunamis are not critical and do notpose a significant loading event on structures. Meanwhile,more than 60% of major tsunamis recorded to date have or-iginated in the Pacific Ocean, highlighting the significantthreat to the coast of British Columbia. History has alsoshown that the eastern seaboard of Canada is also suscepti-ble to tsunamis, however, to a much lesser extent comparedto western Canada.

Received 24 September 2008. Revision accepted 24 June 2009.Published on the NRC Research Press Web site at cjce.nrc.ca on24 November 2009.

D. Palermo,1 I. Nistor, and Y. Nouri. Department of CivilEngineering, University of Ottawa, 161 Louis Pasteur, Ottawa,ON K1N 6N5, Canada.A. Cornett. Canadian Hydraulics Centre, National ResearchCouncil, 1200 Montreal Rd, M-32, Ottawa, ON K1A 0R6,Canada.

Written discussion of this article is welcomed and will bereceived by the Editor until 31 March 2010.

1Corresponding author (e-mail: [email protected]).

1804

Can. J. Civ. Eng. 36: 1804–1815 (2009) doi:10.1139/L09-104 Published by NRC Research Press

The objectives of this paper are to provide a review offorce components directly related to tsunami loading onnear-shoreline structures, present loading combinations forstatic analysis of structures, summarize results of an experi-mental program conducted by the authors, and discuss, usingthe experimental data, the appropriateness of the tsunami-in-duced force components. Yeh et al. (2005) developed aguideline for tsunami shelters, which also discusses tsunamiforce components. They concluded that hydrodynamic andimpact forces are the most probable forces for near-shorelinestructures, whereas this paper includes a surge force compo-nent in the proposed loading combinations. Other differen-ces and similarities are discussed herein.

Tsunami hazardGiven its geographical location, Canada, particularly the

west coast, is susceptible to tsunamis. Table 1 provides ahistorical list of tsunamis that have affected North America.

In addition to the 1700 Cascadia Earthquake and theGrand Banks Landslide, the tsunami of 28 March 1964 is ofsignificance to the western seaboard of Canada. The tsunamiwas triggered by a large earthquake that occurred in Alaska,and resulted in millions of dollars of damage in Port Al-berni, British Columbia.

The seismological features off the coast of British Colum-bia are similar to those in the Indian Ocean, which gener-ated the recent massive tsunami of December 2004. Fromnorthern Vancouver Island to northern California, the Casca-dia Subduction Zone delineates the boundary between thesmaller offshore Juan de Fuca Plate that is sliding under themuch larger North American Plate. Earthquakes of magni-tude 9.0 or greater are possible in this area. The 1700 Cas-cadia Earthquake of magnitude 9.0 was generated by a faultrupture over a length of approximately 1000 km. There arestriking similarities with the approximate 9.3 magnitudeearthquake, which occurred off the coast of Sumatra, Indo-nesia, in 2004. The earthquake was generated by a fault rup-ture of approximately 1300–1600 km along the subductionzone between the offshore Indian Plate, which is slidingunder the Burma Plate.

Popular belief suggests that major nearby cities on thewest coast, including Vancouver and Victoria, which are lo-cated on inland waterways rather than directly on the coast,would be sheltered from the full effect of a tsunami. This,however, has not been supported by past paleogeologicalevidence. The 2004 Indian Ocean Tsunami provides exten-sive evidence of the propagation of waves as they experi-ence diffraction and refraction phenomena when travelingaround islands or over underwater ridges. For example, thewest coast of Sri Lanka was severely affected by the tsu-nami, even though it was not directly in the path of the ini-tial tsunami waves.

For design and analysis purposes, the critical parametersare the expected coastal inundation (flood) levels and inlandvelocities of tsunami-induced hydraulic bores. This providesthe necessary information to estimate tsunami-induced loadson infrastructure located in tsunami-prone areas. For an en-gineer, the most effective tool would be a tsunami hazardmap through which tsunami inundation and velocities wouldbe readily available. However, such maps are currently not

available, and numerical modeling is required to estimate in-undation levels for a given earthquake. Xie et al. (2007)conducted numerical modeling of the Cascadia Fault to in-vestigate the tsunami risk for western Canada. A magnitude9.0 earthquake, similar to the 1700 Cascadia event, was con-sidered in the model. The analyses estimated maximumwave run-ups of up to 25 m along the western shore of Van-couver Island, with an estimated arrival time of 1 h and20 min.

Tsunami-induced force components

While the design of structures in flood-prone areas haspreviously been investigated and is well established, limiteddocumentation is available that specifically addresses the de-sign of near-shoreline structures built in tsunami-proneareas. The Federal Emergency Management Agency(FEMA) published the Coastal construction manual, knownas FEMA 55 (FEMA 2003), which recommends proceduresto calculate tsunami-induced flood and wave loads. The De-partment of Planning and Permitting of Honolulu, Hawaii,developed the City and County of Honolulu building code(CCH 2000). The code contains provisions that apply to dis-tricts located in coastal flood and tsunami-risk areas.

Three parameters are necessary to define the magnitudeand application of tsunami-induced forces: (i) the inundationdepth; (ii) the flow velocity; and (iii) the flow direction.These parameters mainly depend on the following: (a) tsu-nami wave height and wave period; (b) coast topography;and (c) roughness of the coastal inland. The extent of tsu-nami-induced coastal flooding, and therefore the inundationdepth at a specific location, can be estimated using varioustsunami scenarios involving varying magnitude and faultorientation, and modeling coastal inundation accordingly.However, while modeling the offshore tsunami propagationis advanced and reasonably accurate, the estimation of flowvelocity for the inland inundation is difficult. Inundationflow velocities can vary in magnitude from zero to signifi-cantly high values, while flow direction can also vary owingto onshore local topographic features, as well as soil coverand onshore roughness.

The force components explicitly considered by FEMA 55and CCH include hydrostatic forces, buoyant forces, hydro-dynamic forces, surge or wave-breaking forces, and debrisimpact forces. A brief description of these forces is furtherpresented.

Table 1. Historical tsunami events along Canada’s coastlines.

Date LocationMaximumrun-up (m)

4 Nov. 1994 Southern Alaska 7.64 Feb. 1965 Western Alaska 10.728 Mar. 1964 Gulf of Alaska 67.19 Mar. 1957 Central Alaska 22.823 Jun. 1946 British Columbia 30.010 Sept. 1899 Gulf of Alaska 60.018 Nov. 1929 Grand Banks, Newfound-

land26 Jan. 1700 Cascadia, British Columbia

Palermo et al. 1805

Published by NRC Research Press

Hydrostatic forceThe hydrostatic force is generated by still or slow-moving

water acting perpendicular onto planar surfaces. The hydro-static force per unit width, FHS, is calculated as

½1� FHS ¼1

2rg hþ

u2p

2g

!2

where r is the seawater density, g is the gravitational accel-eration, h is the water depth, and up is the normal compo-nent of flow velocity. Equation [1] is adopted by CCH andaccounts for the velocity head, whereas FEMA 55 assumesthe velocity head to be negligible and replaces h with ds,which is defined as the stillwater flood depth. For calcula-tion purposes, h and ds are equivalent. In the case of a bro-ken tsunami wave, which moves inland in the form of aturbulent hydraulic bore, the hydrostatic force is signifi-cantly smaller than the drag and surge forces. Conversely,the hydrostatic force becomes important when the tsunamiis similar to a rapidly rising tide.

Buoyant forceThe buoyant force is a vertical force acting through the

center of mass of a submerged body. Its magnitude is equalto the weight of the volume of water displaced by the sub-merged body, and is determined from eq. [2].

½2� Fb ¼ rgV

where Fb is the buoyant force, and V is the volume of waterdisplaced by the submerged structure. The buoyant force cangive rise to stability problems by reducing the resistance of astructure to sliding and overturning. The buoyancy effectgenerated by tsunami flooding was observed in field obser-vations (Ghobarah et al. 2006), where the rapidly risingwater level lifted and displaced flooring elements, houses,and large oil reservoirs over several hundreds of metres.

Hydrodynamic (drag) forceHydrodynamic (drag) forces arise as the tsunami bore

moves inland and travels around structures with moderate tohigh velocity. The general expression for drag is given as

½3� Fd ¼rCDAu2

2

where Fd is the drag force acting in the direction of flow, A

is the projected area of the body normal to the direction offlow, and u is the tsunami flow velocity. Values for the dragcoefficient (CD) vary between FEMA 55 and CCH. For ex-ample, values of 1.0 and 1.2 are recommended for circularpiles by CCH and FEMA 55, respectively. For the case ofrectangular piles, the drag coefficient recommended byFEMA 55 and CCH is 2.0. The flow is assumed to be uni-form, and therefore, the resultant force will act at the cen-troid of the projected area. The FEMA permits thehydrodynamic force to be converted to an equivalent hydro-static force for flood velocities not exceeding approximately3.05 m/s (10 ft./s).

Surge forceThe surge force is generated by the first impact of the ad-

vancing water front of a tsunami bore on a structure. On ac-count of a lack of detailed experiments specificallyapplicable to tsunami bores running up the shoreline, thecalculation of the surge force exerted on a structure is sub-ject to substantial uncertainty. Accurate estimation of theimpact force in the laboratory is a challenging and difficulttask. The CCH recommends the following formula, adoptedfrom Dames and Moore (1980), which is applicable to verti-cal walls subjected to a hydraulic bore-like wave:

½4� FS ¼ 4:5rgh2

where FS is the surge force per unit width of wall, and h isthe surge height, usually assumed equal to the inundationdepth or flood level.

Wave-breaking forcesClassic wave breaking, as assumed in FEMA 55, considers

the force generated by wave breaking at the shoreline. Anexample is the 1946 Aleutian Tsunami, where the destruc-tion of Scotch Cap Lighthouse, Unimak Island, was appa-rently the result of a wave breaking directly on the structure(Yeh et al. 2005). Tsunami waves, however, tend to breakoffshore and approach the shoreline in the form of a rapidlyadvancing hydraulic bore. Therefore, wave-breaking forcesare not directly applicable to the case of tsunami bores,and are not considered as a force component in this paper.Furthermore, the focus herein is on inland structureslocated close to the shoreline, which would not be affectedby the actual surf zone wave breaking.

Fig. 1. Tsunami bore velocity versus inundation depth.

1806 Can. J. Civ. Eng. Vol. 36, 2009


Debris impact forceA high-speed tsunami bore traveling inland carries debris

such as floating automobiles, floating pieces of buildings,drift wood, as well as boats and ships. The impact of float-ing debris can induce significant forces on a building, lead-ing to structural damage or collapse (Saatcioglu et al. 2006).Both FEMA 55 and CCH account consistently for debris im-pact forces, using the following expression:

½5� Fi ¼ mu

Dt

where Fi is the impact force, m is the mass of the body im-pacting the structure, u is the approach velocity of the float-ing body (assumed equal to the flow velocity), and ~t isthe impact duration taken equal to the time between the in-itial contact of the floating body with the structure and theinstant of maximum impact force. The recommendationsfor impact duration vary between CCH and FEMA 55. Forexample, CCH recommends impact duration of 0.1 s forconcrete structures, while FEMA 55 provides different va-lues for walls and piles for various construction types. Forconcrete construction, wall elements are assigned an impactduration between 0.2 and 0.4 s, and piles, between 0.3 and0.6 s.

According to FEMA 55, the impact force (a single con-centrated load) acts horizontally at or near the water surface.Its magnitude is equal to the force generated by 455 kg(1000 lb) of debris carried by the hydraulic bore. The impactforce is to be applied to the structural element at its mostcritical location, as determined by the structural designer.The debris impact load using the mass suggested by FEMAtypically does not represent a significant component to thetotal lateral load imposed on a building relative to the otherforce components. However, the debris impact force is sig-nificant in the design of the structural component that is as-sumed to be subject to the impact.

VelocityDifferences in estimating hydrodynamic forces exerted on

structures by tsunami-induced bores, as well as impact ofdebris, are due to differences in estimating flow velocity.The hydrodynamic force is proportional to the square of theflow velocity. Thus, uncertainties in estimating velocities in-duce large differences in the magnitude of the resulting hy-drodynamic force. Tsunami inundation velocity magnitudeand direction can vary significantly during a major tsunamiinundation. Current estimates of the velocity are crude; aconservatively high flow velocity impacting the structure ata normal angle is usually assumed. Also, the effects of run-up, backwash, and direction of velocity are not addressed incurrent design codes.

The general form of the tsunami-induced flow velocity isshown below:

½6� u ¼ Cffiffiffiffiffiffiffigds

pwhere u is the flow velocity, ds is the inundation depth, andC is a constant coefficient. A number of codes and research-ers have proposed estimates of velocity for given tsunamiinundation levels, such as FEMA 55 (Dames and Moore1980), Iizuka and Matsutomi (2000), CCH (2000), Kirkoz

(1983), Murty (1977), Bryant (2001), and Camfield (1980).Figure 1 illustrates the wide scatter in calculated velocities.The FEMA assumes a constant coefficient of 2, whereas theCCH considers the velocity to be equal in magnitude to theinundation level. Note, Camfield (1980) suggested a similarvelocity to FEMA 55.

Discussion of tsunami-induced forcecomponents

There has been significant debate concerning tsunami-in-duced force components, particularly the surge force. Fur-thermore, the tsunami inundation flow velocity is ofquestion, which will significantly affect the drag and debrisimpact forces. The surge force given in eq. [4] results fromthe summation of hydrostatic and hydrodynamic force com-ponents at the instant the tsunami bore impacts a structure asgiven by the following expression:

½7� Fs ¼1

2rgh2bþ 1

2Cdru2hb

For infinitely long walls, Cd is taken as 2, and the velocityu is assumed equal to 2

ffiffiffiffiffigh

pat the leading edge of the tsu-

nami. The water depth h is assumed equal to the surgeheight and typically is taken as equal to the inundation level,and b is the width of the structure. Substitution into eq. [7]results in the following:

½8� Fs ¼1

2rgh2bþ 4rgh2b ¼ 4:5rgh2

The hydrostatic and hydrodynamic force terms in eq. [8]are added, resulting in a surge force with the same magni-tude as assumed by CCH (eq. [4]). Therefore, CCH implic-itly assumes a larger velocity for the calculation of the surgeforce than for the drag force as given in eq. [3]. Thus, theinitial velocity of the leading edge of the tsunami-inducedbore is assumed greater than that of the semi-steady flowaround the structure. In Japan, Okada et al. (2005) devel-oped the Structural Design Method of Buildings for Tsu-nami Resistance (SMBTR), which also assumes a surgeforce as given by eq. [4]. However, the SMBTR documentconsiders the surge force only in calculating the total tsu-nami load. Both CCH and SMBTR assume the surge forcecan be represented by a hydrostatic pressure distribution ex-tending 3h from the base of the wall, with the point of appli-cation of the resultant force located at a distance h above thebase. The pressure at the base of the wall is directly propor-tional to the height over which the pressure is assumed toact and is equal to 3rgh. This methodology is applicable forwalls with heights equal to, or greater than, 3h. Walls lessthan 3h require surge forces to be calculated using a combi-nation of hydrostatic and drag force components as given ineq. [7]. The magnitude of the surge force as calculated byeq. [8] is nine times the hydrostatic force for the same surgedepth. Yeh et al. (2005) have suggested that the surge asgiven in CCH may result in excessively overestimatedforces. The pressure distribution gives the impression thatthe surge height is three times the inundation design level.Furthermore, it assumes that the point of application of theresultant force is at the same height as the inundation level.

Palermo et al. 1807


The authors also argue that such an approach does not ap-pear to have a correct physical interpretation. The magnitudeof the surge height is significantly lower, as documented byvarious post-tsunami surveys.

Laboratory testing has shown that surge forces can ariseunder certain laboratory conditions. Arnason (2005) ob-served the initial impact (surge force) overshoot the dragforce owing to the passing bore in the case of a square col-umn for small bore heights. The maximum surge force wasequivalent to 1.5 times the subsequent hydrodynamic force.On the other hand, this overshoot was not observed forlarger bore heights. Also, no overshoot was recorded for thecase of circular or rhomboidal columns. Ramsden (1993) in-vestigated the surging forces on a vertical wall and noted anovershooting force of approximately 1.5 times the hydrody-namic force under conditions of a turbulent hydraulic bore.However, this overshoot did not occur at the initial impactof the bore on the structure, but occurred shortly after.Ramsden (1993) also observed a lack of overshoot for con-ditions of dry flume beds. However, a tsunami typicallybreaks offshore and travels inland as a broken hydraulicbore, which is not well simulated by dry bed conditions.

The validity of eq. [4] for calculating the tsunami loadingwas investigated by Nakano and Paku (2005). The investiga-tion was based on surveyed structures in Sri Lanka andThailand that were damaged during the 2004 Indian OceanTsunami. Data on surge levels was recorded on walls andcolumns. The pressure coefficient 3, as assumed by CCHand SMBTR for the surge force, was assigned as a parame-ter a, which represented the dividing line between damageand no damage conditions. The study indicated that a = 3.0is appropriate for walls and a = 2.0 for columns. Hence, thisstudy suggests that the coefficient for walls is in agreementwith the design surge forces prescribed by CCH andSMBTR.

Tsunami-induced loading combinations

Force components that are directly associated with tsu-nami-induced turbulent hydraulic bores consist of (1) hydro-static force, (2) hydrodynamic (drag) force, (3) buoyantforce, (4) surge force, and (5) debris impact. The designdocuments previously discussed, specifically CCH andFEMA 55, do not specifically provide loading combinationsto estimate the maximum tsunami load for design. In thecase of SMBTR, the tsunami load is determined only fromthe surge force component. The FEMA 55 provides loadcombinations for flood loads, which includes wave breaking.However, modifications are necessary to derive loadingcombinations that are directly applicable to tsunamis. Yehet al. (2005) recommended that tsunami shelters located inthe inundation zone, but inshore, be designed for hydrody-namic (drag) and debris impact. The surge force that is gen-erated owing to the formation of a turbulent bore isneglected, since Yeh et al. (2005) have based their recom-mendation considering the effects of dry-bed surges, wherethe initial impacting force (surge force) does not overshootthe drag force. Dias et al. (2005) proposed two loading com-binations: point of impact and post-submergence. The pointof impact considers the initial impact of the tsunami waveand is estimated as the sum of drag and hydrostatic force

components on the upstream face of the structure. The post-submergence includes drag on the upstream face, hydrostaticforces on the upstream and downstream faces, and buoy-ancy. The impact of debris is not explicitly included in ei-ther of the load combinations. Pacheco and Robertson(2005) analyzed structures to various inundation levels. Toestimate the tsunami load, loading combinations wereadopted from FEMA 55. In the estimation of the tsunamiload, wave-breaking forces were omitted. For columns di-rectly exposed to the tsunami wave, the load was estimatedas a combination of hydrodynamic and debris impact. Thetsunami load for structural walls placed parallel to the shore-line, and perpendicular to the flow of the tsunami wave, wasconsidered the maximum of two combinations. The firstcombination included the combined effect of hydrodynamicand debris impact forces, whereas the second consideredsurge and debris impact forces. Nouri et al. (2007) proposedloading combinations specifically for turbulent bores gener-ated by tsunamis, as shown in Fig. 2. Two combinationswere developed, which were based on modifications of thoserecommended by Dias et al. (2005). The first combination(Initial Impact) considers the first arrival of the tsunamibore on a structure, and includes the combined effect ofsurge and debris impact forces. These force components canbe evaluated using eqs. [4] and [5], respectively. The secondcombination (Post Impact) considers the flow of the tsunamibore around the structure. Hydrodynamic (drag), debris im-pact, and hydrostatic forces are combined to determine thelateral loading, and can be calculated from eqs. [3], [5], and[1], respectively. The net hydrostatic force is assumed to bezero. Consideration is also given to the buoyancy force,which can cause sliding and overturning instability.Although the surge force has been included, further investi-gation is required to provide a more accurate estimate of thisforce component for structural walls and columns. Analysesof structures, considering the loading combinations of Nouriet al. (2007), have been performed to estimate the magni-tude of tsunami loads on structures relative to seismic loads(Palermo et al. 2007; Nouri et al. 2007). The work of Pa-lermo et al. (2007) illustrated that tsunami-induced baseshears for an inundation level of 3 m could exceed seismicbase shears for structures up to 20 storeys in height in the

Fig. 2. Proposed tsunami loading combinations: (a) initial impact;(b) post impact.

1808 Can. J. Civ. Eng. Vol. 36, 2009


Vancouver area. The resulting tsunami loads were dependanton the portion of exterior nonstructural elements that wereassumed to remain intact during the loading.

The tsunami load can readily be incorporated in buildingcodes and combined with other loads. Given that a tsunamiis an extreme event, load cases adopting the philosophy ofseismic loading are suggested as a preliminary framework.Following the format of the 2005 NBCC (NRCC 2005),three load cases are considered. The first load case considersTsunami (T) and Dead (D) loads only. The second load caseincludes companion loads, specifically Live (L) and Snow(S) loads. The third case should only be considered if earlywarning systems provide sufficient warning to allow occu-pants to exit buildings safely. However, this load caseshould not be applied to tsunami shelters.

½9�1:0T þ 1:0D

1:0T þ 1:0Dþ 0:5Lþ 0:25S

1:0T þ 1:0Dþ 0:25S

Note that in the case of the Cascadia Subduction Zonealong the western coastline of British Columbia, damage tostructures may initially occur owing to the triggering earth-quake before the tsunami loading arrives. In such cases, en-gineers should consider the effects of the tsunami load onsoftened or damaged structures.

Experimental programA comprehensive and exhaustive experimental program

was conducted at the Canadian Hydraulics Centre in Ottawa,Canada. The main objective of this study was to betterunderstand the force components associated with turbulentbores, which are representative of tsunami waves travelinginland. The testing was carried out in a high-discharge flumemeasuring 10 m long, 2.7 m wide, and 1.4 m deep and isserviced by pumps with a variable discharge flow up to1.7 m3/s. The flume was partitioned to 1.3 m in width and7.3 m in length to create a narrower channel for testing. Togenerate a turbulent bore, a hinged gate was designed andfastened near the upstream end of the channel. In the closedposition, the gate supported (impounded) a specified waterlevel. The hinging mechanism of the gate allowed a rapidopening, which resulted in the formation of a turbulent hy-draulic bore that moved downstream from the gate. Thismechanism is typical of a dam break phenomenon, and hasbeen shown to realistically simulate turbulent bores gener-ated during tsunamis (Nouri et al. 2008). Figure 3 illustratesthe features of the flume used during the experimental pro-gram. Additional details of the experimental program areavailable elsewhere (Nouri et al. 2008; Nouri 2008).

Forces were measured on two structural components:square or diamond, and circular sections (Fig. 4). The circu-lar section was made of PVC pipe and measured 0.32 m indiameter, whereas the square or diamond section was as-sembled from acrylic Plexiglas and measured 0.2 m �0.2 m. The circular section was mounted to a six-axis dyna-mometer, allowing base shears and moments to be recordeddirectly. In addition, 10 pressure transducers were placedvertically and flush with the column surface on the circularsection. This was used to establish the time–history of thepressure profiles. The square or diamond section, on the

other hand, was instrumented with five pressure plates,which recorded local forces. The flume was equipped withacoustic doppler velocimeter (ADV) sensors and wavegauges to record flow velocities and depths, respectively.The testing program consisted of 11 test series, and includedvarying upstream impoundment depths, debris weights, andconstrictions. For brevity, a sample of the experimental re-sults for the circular section will be discussed herein, specif-ically to provide a comparison with tsunami-induced forcecomponents commonly used and previously discussed. Addi-tional test results are available elsewhere (Nouri et al. 2008;Nouri 2008).

Global force–time histories of the base shear in the direc-tion of the hydraulic bore were recorded for the circular sec-tion, for upstream water levels (impoundment depths) of0.50, 0.75, 0.85, and 1.0 m, as shown in Fig. 5. Note thatthe impoundment depth is denoted by h0 in the figure.

For each impoundment depth, an abrupt rise in force(surge force) occurs at the instant the hydraulic bore impactsthe circular column section. Approximate surge forces of 56,128, 180, and 295 N were recorded for the 0.5, 0.75, 0.85,and 1.0 m impoundment depths, respectively. This impulsiveforce was followed by a decrease in the base shear. For the0.75, 0.85, and 1.0 m impoundment depths, the reductionwas approximately 55%–60% of the initial impacting force.For the 0.50 m impoundment, the drop was approximately30%. Figure 6 shows the drop in force for the 1.0 m im-poundment depth. This drop was followed by a gradual in-crease in force and was caused by run-up of the hydraulicbore on the column, which is more apparent with the0.85 and 1.0 m impoundment depths as shown in Fig. 5.Run-up forces of 80, 157, 242, and 305 N were recorded,respectively, for impoundments of 0.5, 0.75, 0.85, and1.0 m. The run-up force was greater than the initial impact.The run-up was followed by an approximately constantforce (drag force) caused by the flow of the bore around thestructure. The measured drag forces were approximately115, 210, 230, and 240 N, respectively, for the 0.5, 0.75,0.85, and 1.0 m impoundments.

The drag force was the largest force component in thehistory of the recorded base shears for the 0.5 and 0.75 mimpoundments, whereas the run-up force was critical for the0.85 and 1.0 m impoundments. This indicates that for a cir-cular column section, the drag force is the most criticalforce component for small bore depths; however, the surgeand subsequent run-up force become increasingly critical asthe bore depth increases. Note that this is based on im-poundment depths not exceeding 1.0 m. Additional testingincluding larger bore depths and varying aspect ratio of thestructural testing components is required to better under-stand the individual force components.

Pressure–time histories were recorded for the circular sec-tion subjected to impoundment depths of 0.5, 0.75, 0.85, and1.0 m. Figure 7 provides the recorded pressures for the1.0 m impoundment depth along the height of the test speci-men. Similar responses were recorded for all other impound-ments depths.

Four pressure readings are shown in Fig. 7, starting at50 mm up to 250 mm from the base of the test structure.This was the range over which pressures were recorded atthe instant the turbulent bore impacted the section. The re-

Palermo et al. 1809


sponse demonstrates an initial spike in pressures, which cor-responds to the initial impact or the so-called ‘‘surge force.’’Immediately following is a drop in pressures, quickly fol-lowed by a gradual increase, corresponding to the run-upforce. Later in the response, the pressures converge, whichis associated with the drag force. The pressure–time historiesare in agreement with the forces recorded at the base of thestructure as shown in Fig. 5.

Loading on the circular section due to impact of floatingdebris conveyed by the generated hydraulic bores was alsoinvestigated. The Indian Ocean Tsunami of 2004 highlightedthe destructive effects as a result of debris transported by the

tsunami flow. In this study, two wooden logs (large andsmall) were used to simulate debris impact. The largewooden log was 445 mm long, with a 90 mm � 90 mmcross section. The small log was 435 mm long, with a rec-tangular cross section of 60 mm � 45 mm. The masses were1.479 and 0.474 kg for the large and small logs, respec-tively. The small debris was conveyed by 0.75 m impound-ment depths, and the large debris with 0.75 and 1.0 mimpoundments. The small wooden log induced debris impactforces ranging from 158 to 196 N, with an average force of177 N and an average rise time of 0.0078 s. The largewooden log generated debris impact forces between 365–

Fig. 3. Wave flume: (a) elevation view; (b) plan view (after Nouri et al. 2008).

Fig. 4. Structural components: (a) circular; (b) square or diamond.

1810 Can. J. Civ. Eng. Vol. 36, 2009


591 and 558–692 N for the 0.75 and 1.0 m impoundmentdepths, respectively. The corresponding average forces andaverage rise times were 487 N and 0.0073 s, and 621 N and0.0075 s, respectively, for the 0.75 and 1.0 m impound-ments. The scatter in recorded debris impact forces was aresult of the differences in angle of impact between themasses and structure, which was confirmed by videos takenduring testing. Figure 8 illustrates the results of typical re-sponses for 1.0 and 0.75 m impoundment depths subjectedto the larger wooden log.

Similarities in recorded force–time histories of the baseshear are evident. The hydraulic bore causes an initial im-pact (surge force) followed by a large spike in force a shorttime after, which is caused by the impact of the debris. De-

bris impact forces of 692 N with a rise time of 0.0075 s and591 N corresponding to a rise time of 0.008 s were recordedfor the 1.0 and 0.75 m impoundment depths, respectively.Bore front velocities of 4.6 and 3.3 m/s were estimated forthe 1.0 and 0.75 m impoundments, respectively, which ex-plains the differences in debris impact forces. The velocitieswere determined using wave gauges, and corresponded tothe leading edge of the hydraulic bore and the time of im-pact of the debris.

A ‘‘bounce back’’ phenomenon was observed for a num-ber of debris impact tests. The bounce back effect was theresult of a second debris impact immediately following thefirst impact. In all cases, the magnitude of the second impactforce was less than the first; however, similar rise timeswere recorded.

Fig. 5. Force–time histories for circular section.

Fig. 6. Force–time history for 1.0 impoundment depth.

Palermo et al. 1811


Fig. 7. Pressure–time histories for circular section.

Fig. 8. Debris impact testing.

Fig. 9. Individual force components on circular section.

1812 Can. J. Civ. Eng. Vol. 36, 2009


Discussion of resultsThe force–time history for the circular test specimen sub-

jected to a 1.0 m impoundment depth with superimposedbore depth, which was measured in front of the circular sec-tion, is shown in Fig. 9. The response is used to evaluate thetsunami-induced force components previously discussed.Similar responses were recorded for impoundment depths of0.5, 0.75, and 0.85 m.

The surge, run-up, and drag force components are high-lighted. Typically, the force components are calculated froma given inundation level, usually corresponding to the flowdepth associated with the hydrodynamic (drag) flow condi-tions. Based on the response in Fig. 9, the inundation levelcorresponding to the drag flow is approximately 0.7 m,based on the approximately constant bore height from 14 to18 s. Using eq. [4], a surge force of 6920 N is calculated,which excessively overestimates the recorded surge force of295 N. A closer examination of the response indicates thatthe bore depth was approximately 245 mm at the instant ofthe initial impact. For a bore depth of 245 mm, a surge forceof approximately 850 N is estimated. Further improvementsare obtained if the surge force is computed as a combinationof hydrostatic and hydrodynamic force components as dis-cussed for eq. [7], assuming a bore front velocity given byeq. [6] with a coefficient of 2, and a drag coefficient of 1.2prescribed by FEMA 55. For impoundment depths of 0.5,0.75, 0.85, and 1.0 m, the corresponding bore depths were125, 140, 230, and 245 mm, respectively. The calculatedsurge forces according to eq. [7] are 142, 178, 482, and 546N, corresponding to ratios of calculated to measured surgeforces of 2.54, 1.39, 2.67, and 1.87, respectively, for the0.5, 0.75, 0.85, and 1.0 m impoundments. While this ap-proach provides improved predictions, the calculated surgeforce is still overestimated, with significant scatter in the re-sults. The drag force component is significantly larger thanthe hydrostatic force component, thus the discrepancies area result of the high velocity of the leading edge of the hy-draulic bore, also indicated by other researchers (Yeh et al.2005).

An evaluation of the drag forces using eq. [3] also high-lights the importance of more accurate estimates of the ve-locity of turbulent bores around structural elements. Twovelocity estimates are investigated: FEMA 55 using eq. [6]with a coefficient of 2; and CCH, which assumes the veloc-ity is equal to the depth of water at the structure. The meas-

ured bore depths corresponding to the drag forces wereapproximately 370, 550, 600, and 700 mm for the 0.5, 0.75,0.85, and 1.0 m (Fig. 9) impoundment depths, respectively.The FEMA 55 estimates drag forces of 1030, 2280, 2710,and 3690 N, whereas the CCH predicts drag forces of 8, 27,35, and 55 N, respectively, for the given impoundments. Themeasured drag forces were approximately 115, 210, 230,and 240 N, respectively. Therefore, FEMA 55 grossly over-estimated the drag forces, while CCH significantly underes-timated the drag. Velocities for the 0.5, 0.75, 0.85, and1.0 m impoundment depths, of 1.27, 1.41, 1.41, and 1.34 m/s,respectively, are approximated from eq. [3] using the re-corded drag forces as input, along with a drag coefficientof 1.2 as prescribed by FEMA 55. The approximated ve-locities are smaller than FEMA 55, but larger than CCH.The average velocity is 1.35 m/s, and the average velocityconstant coefficient, C, determined from eq. [6] is 0.59. Itis evident that improved estimates of the flow velocity arenecessary to properly evaluate drag forces. The coefficientof 2 used in FEMA 55 appears to be more appropriate forthe leading edge of the turbulent bore, and does not appearto be appropriate for the hydrodynamic flow condition.

The pressure–time histories illustrated in Fig. 7 providepressure profiles corresponding to the individual force com-ponents shown in Fig. 9 by plotting the pressures along theheight of the structure at the instant the respective forcecomponents arise. Figure 10 provides the pressure distribu-tion for the 1.0 m impoundment depth, which can be usedfor static and dynamic analysis; the latter would also requireduration of the loading.

The initial impact (surge force) resulted in a highly trian-gular pressure distribution, with the maximum pressure re-corded at 100 mm from the base of the structure. Thepressure at 250 mm was negligible, indicating that the boreheight at initial impact was approximately 250 mm. Thus,for the 1.0 m impoundment depth on the circular section,the maximum pressure occurred at 40% of the height of thebore from the base of the structure. The run-up and dragforces induced pressures up to 550 mm. These pressures areslightly more constant over the height relative to the surgeforce. Note that variations in the velocity of the flow, aswell as a significant level of air entrainment throughout thedepth of the bore, resulted in deviations from a constant-pressure distribution. The pressure profile for the drag forcecomponent is consistent with the recommendations of

Fig. 10. Vertical pressure distributions corresponding to force components.

Palermo et al. 1813


FEMA 55, which suggests that the drag force can be con-verted to an equivalent hydrostatic force for flow velocitiesnot exceeding 3.05 m/s. The drag force, as indicated inFig. 9, is approximately 240 N. Using eq. [3], with a boredepth of 0.7 m and a drag coefficient of 1.2 for round pilesas prescribed in FEMA 55, an average velocity of 1.34 m/sis calculated, which is less than the limiting value for a hy-drostatic distribution, thus confirming the recommendationsof FEMA 55.

Debris impact forces are calculated by FEMA 55 and CCHusing eq. [5], which is based on the impulse–momentumapproach. The estimated velocities were 3.3 and 4.6 m/s forthe 0.75 and 1.0 m impoundment depths. The averagerecorded rise times were 0.0078, 0.0073, and 0.0075 s forthe small log, large log with 0.75 m impoundment, and largelog with 1.0 m impoundment, respectively. Based on eq. [5],debris impact forces of 200, 669, and 907 N are predicted,respectively. The average measured forces were 177, 487,and 621 N, respectively, indicating that the impulse–momen-tum approach provides a reasonable estimate of debris load-ing for this test program. Note that the discrepancy betweencalculated and recorded forces is, for the most part, a resultof the variable angle of impact between the masses and struc-ture, and the approximate estimate of flow velocity. Furtherwork is required to investigate other debris impact loadingformulations available in the literature.

ConclusionsA detailed discussion of the force components expected

during a tsunami-induced inundation for structures locatedinland was presented. In addition, sample test results from acomprehensive experimental program conducted to investi-gate the interaction between turbulent hydraulic bores andsimple structures, with the objective of quantifying the forcecomponents associated with tsunami-induced loading, werepresented. Some preliminary findings are provided:

1. Currently, building codes are either silent or provideconflicting guidance for estimating tsunami loads for in-land structures located in tsunami-prone regions.

2. The Cascadia Subduction Zone in Canada, which has si-milar seismological features to the subduction zone offthe coast of Indonesia, where the 24 December IndianOcean Tsunami surfaced, has great potential to generatesignificant tsunamis if megathrust earthquakes arise.

3. It was observed that surge, run-up, and drag forces aregenerated when turbulent bores impact the structuralshapes tested.

4. The drag force was the largest force component for thesmaller impoundment depths (shallower bores), and thesurge and run-up were more critical for the larger im-poundment depths (deeper bores).

5. The run-up force was greater than the surge force for allbore depths.

6. The surge force was best estimated using a combinationof hydrostatic and drag force components.

7. The force–time and bore height – time histories indicatethat different bore heights are associated with the surgeand drag force components, and therefore, using a singlebore height in the estimation of the force componentsmay not be suitable.

8. The drag force was not well predicted using currentlyavailable design formulations. This was a direct result ofthe inaccurate flow velocity assumed by design codes.

9. The debris impact forces could be reasonably estimatedusing the impulse–momentum approach.

10. Time histories of the tsunami flow velocity would pro-vide improved estimates of the force components.

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