ARTICLE

Centrifuge testing to simulate buried reinforced concrete pipejoints subjected to traffic loadingBoris Rakitin and Ming Xu

Abstract: Pipeline water leakage has become a serious problem in many countries. It has been widely noted that most of thedamage to the pipelines occurs in the joints where two pipes are connected to each other. This paper presents the results of ageotechnical centrifuge testing program in which the response of a 12 m long (in prototype scale) large-diameter reinforcedconcrete pipeline with gasketed bell-and-spigot joints subjected to three standard American Association of State Highway andTransportation Officials design load configurations has been investigated. The results show that most vertical pipe movementsoccurred during the first 10 cycles of traffic loading. Under design tandem loading, the pipe joint displacements were signifi-cantly higher than those under the other two traffic load configurations. An increase of soil cover depth resulted in a reducedinfluence of surface loading, the effect of which was the most significant for two single pairs of wheels of design trucks in passingmode. Furthermore, two pipes on the left side and two pipes on the right side from the tested joint were influenced significantlyby the surface loading, while the pipeline movements were not symmetrical. Although the joint directly under the loadexperienced the largest rotation, the possibility of leakage in the second joint in the spigot-to-bell direction was also high, dueto large differential displacement between the pipes.

Key words: reinforced concrete pipeline, joints, geotechnical centrifuge testing, cyclic loading, traffic load.

Résumé : Les fuites d’eau dans la canalisation sont devenues un sérieux problème dans un bon nombre de pays. On a noté que,généralement, la grande partie des dommages causés aux réseaux de canalisation surviennent dans les joints, a l’endroit où deuxtuyaux sont raccordés. Cette étude présente les résultats d’un programme d’essais a la centrifugeuse géotechnique dans le cadreduquel on a examiné la réponse d’une canalisation de 12 m de longueur (en échelle prototype) de grand diamètre en béton armémunie de joints étanches a emboîtement et bout uni qui a été soumise a trois configurations de charge de calcul selon les normesde l’American Association of State Highway and Transportation Officials. Les résultats montrent que la plupart des déplacementsverticaux des tuyaux se sont produits durant les 10 premiers cycles de chargement de trafic. Sous un chargement de calcul entandem, les déplacements des joints des tuyaux étaient considérablement plus élevés que ceux sous les deux autres configura-tions de charge de trafic. Une augmentation de l’épaisseur de la couverture de sol a atténué l’influence du chargement de surfaceet l’effet de l’augmentation fut le plus important sur les deux paires de roues simples des camions de conception en mode depassage. De plus, le chargement de surface a considérablement influencé les deux tuyaux du côté gauche et les deux tuyaux ducôté droit du joint en essai, alors que les déplacements de la canalisation n’étaient pas symétriques. Bien que le joint situédirectement sous la charge ait subi la plus importante rotation, la possibilité de fuite dans le deuxième joint dans la direction dubout uni orienté vers l’emboîtement était élevée aussi, a cause d’un important déplacement différentiel entre les tuyaux.[Traduit par la Rédaction]

Mots-clés : canalisation de béton armé, joints, essais a la centrifugeuse géotechnique, chargement cyclique, charge du traffic.

IntroductionReinforced concrete (RC) pipelines are widely used in under-

ground infrastructure construction. Joint performance plays animportant role in overall pipeline performance as joint failureleads to significant economic and water resource losses due towater infiltration or exfiltration. If the situation is critical, thewhole pipeline can reach its service limit, because the integrity ofpipe–soil system is affected (Elachachi et al. 2004; Vipulanandanand Liu 2005).

A significant amount of research has been carried out regardingpipe barrels (i.e., Lay and Brachman 2014; Rakitin and Xu 2014;Clayton et al. 2010), but little has been done in regards to pipejoints. However, there are some studies related to pipeline jointproblems and their importance. Based on a closed-circuit televi-sion (CCTV) inspection of an 1800 km long sewer network, Bucoet al. (2006) concluded that 26.7% of concrete pipe defects are

related to joint displacements or openings. Romer and Kienow(2004) reported that joints are a significant source of reportedconcrete pipe installation problems and pipe failures, becausemanufacturers of pipes are responsible for the design of gasketedconcrete pipe joints that need to be approved by the purchaser.However, there is little guidance available to the purchaser aboutthe procedure for joint design approval.

In addition, there are several American Society for Testing andMaterials (ASTM) standards related to joints and gaskets. ASTM(2014) C361 specifies basic manufacturing and performance re-quirements of joints and rubber gaskets. ASTM standards C443(ASTM 2005), C497 (ASTM 1998), and C1214 (ASTM 2002) describejoint leakage, shear, and integrity testing, respectively. Also ASTM(2006) C1628 covers rubber gasketed concrete pipe joints withmeasurable or defined infiltration or exfiltration. However, thesestandards mostly describe the performance requirements of pipe

Received 7 November 2014. Accepted 7 April 2015.

B. Rakitin and M. Xu. Department of Civil Engineering, Tsinghua University, Beijing 100084, China.Corresponding author: Ming Xu (e-mail: mingxu@mail.tsinghua.edu.cn).

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Can. Geotech. J. 52: 1762–1774 (2015) dx.doi.org/10.1139/cgj-2014-0483 Published at www.nrcresearchpress.com/cgj on 16 April 2015.

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joints and their testing, and there is uncertainty about the influ-ence of surface live loads on pipe joints.

To investigate the rotation and shear force demands of RC pipejoints under live loads, Becerril García and Moore (2015) per-formed laboratory tests on a pipeline assembled with two com-plete and two half-pipes cut from a third pipe, subjecting it tosurface live loads. They concluded that depending on the magni-tude and location of the surface live load relative to the joint, thejoint would exhibit rotation or shear displacement between thepipes being connected. They also mentioned that when rigid pipeswith moment–release joints (i.e., bell-and-spigot joints) are sub-jected to surface live loading, their response can be approximatedas rigid links (beams), due to their high flexural rigidity, undergo-ing rotation in the joints. In addition, Becerril García and Moore(2013) observed that the maximum rotation in the joint was ob-tained when the load was applied directly over the joint and themaximum shear (displacement between the bell and spigot) wasmeasured when the surface load was applied over the barrels.Wang and Moore (2013) developed design equations for estimat-ing shear force and rotation in bell-and-spigot joints of rigid pipes.The equations were closed form solutions based on modeling thepipe as an elastic beam resting on the Winkler soil model.

Buco et al. (2008) developed an experimental apparatus for full-scale testing of the joint between two concrete pipes under vari-ous loading conditions. During the tests, the pipes were placed ona loading frame and articulated supports. They were tested with-out the influence of the surrounding soil. Force-displacementsgraphs were obtained based on compression, shear, and bendingtests. These test results were used for developing and validating arheological bilinear model of the joint.

Despite the limited studies on reinforced concrete pipe joints,the pipelines in previous investigations mostly consisted of onlytwo complete pipes. Thus, the influence of adjacent pipes on thecentral tested pipe joint is uncertain. In many cases, surface liveloads were applied without taking into consideration the cyclicnature of traffic loading. Furthermore, in previous studies, pipejoints were mostly tested under a single pair of design truckwheels, but other live load configurations according to AASHTO(2010) and ACPA (2011) are possible as well.

The objective of the research presented in this paper is toreport the results of a geotechnical centrifuge testing program inwhich the behavior of a 12 m long reinforced concrete pipeline,consisting of four complete pipes and two half-segments, sub-jected to live loading, was studied. The pipes have an inner diam-eter of 1400 mm with gasketed bell-and-spigot joints. The liveloads were applied in accordance with AASHTO (2010). The pipe-line was tested under three possible live load configurations todetermine the most critical one. Since in reality pipelines workunder cyclical live loads, the RC pipe joints were tested undercyclic loading. Furthermore, the depth of soil cover was varied fordifferent pipeline installation conditions. Thus, the pipe jointswere tested under two common soil cover depths. Since a pipelineconsisting of pipes with bell-and-spigot joints is not symmetrical,the vertical pipe displacements on both the left and right sidesfrom the tested joint were measured.

Geotechnical centrifuge modeling

Test facilityThe 50 g-ton geotechnical beam centrifuge with a nominal ra-

dius of 2.0 m at Tsinghua University (Rakitin and Xu 2014) wasused for performing all tests presented in this paper (Fig. 1). Acentrifuge on-arm data acquisition system was controlled re-motely. Test data were monitored during the experiment from acentrifuge control room. An aluminum alloy strongbox with aninner space of 490 mm in length, 460 mm in width, and 520 mmin height was used for all centrifuge tests. Prototype stress condi-

tions were created by applying a 20g acceleration in this research.Some relevant centrifuge scaling laws are summarized in Table 1.

Pipeline modelThe tested pipeline consisted of four complete pipes and two

half-segments cut from a fifth complete pipe, as shown in Fig. 2.Five gasketed bell-and-spigot joints connected the six pipe seg-ments. The total length of the pipeline in prototype scale was12 m. Because the strongbox dimensions were limited, the pipe-line model was laid in the strongbox in a diagonal direction toincrease the total pipeline length. There was a distance of at least20 mm between the pipeline ends and the strongbox.

An aluminum model was used to simulate a typical reinforcedconcrete pipe: Sb-Gm 1400 × 2500, Type 1DNA EN 1916 according toBS EN 1916 (BSI 2002), with an inner diameter of 1400 mm, aworking length 2500 mm, and a wall thickness of 165 mm (Fig. 3a).

The gravitational acceleration, N, used in this study was 20g.Thus, all length dimensions and flexural stiffness per unit width(N·m2/m) between the model and the prototype were calculatedusing scaling laws, which are shown in Table 1.

The pipe model had an inner diameter of 70 mm. The model wasmade of aluminum alloy with a wall thickness of 6.2 mm (Fig. 3b).The thickness of the aluminum model wall was converted to thatof concrete with equivalent flexural stiffness (Rakitin and Xu2014). The elastic moduli for aluminum and concrete are 70 and30 GPa, respectively.

The prototype pipes and small-scale aluminum models havegasketed bell-and-spigot joints, as shown in Fig. 3c, which weresealed with a compressible rubber O-ring gasket for water tight-

Fig. 1. 50 g-ton geotechnical beam centrifuge at TsinghuaUniversity.

Table 1. Some relevant centrifuge scaling laws(Taylor 2005).

ParameterScaling law(model/prototype)

Gravity (m/s2) 1NLength (m) 1/NArea (m2) 1/N2

Volume (m3) 1/N3

Density (kg/m3) 1Unit weight (N/m3) 1NFlexural stiffness per unit

width (N·m2/m)1/N3

Flexural stiffness (N·m2) 1/N4

Stress (N/m2) 1Strain 1

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ness purposes. The gasket shape is believed to have no influenceon the mechanical behavior of the joint (Buco et al. 2008), and thegasket thickness was 1 mm on the model scale, which equals20 mm in the prototype scale. The geometrical parameters of thesmall-scale aluminum model joint were designed to be equal tothe pipe joint in a full scale, so as to simulate the actual pipe jointbehavior. This type of joint allows a restricted pipe rotation of 0°–5°(Rajani and Abdel-Akher 2013), and limited vertical (the gap betweenthe bell and spigot is 20 mm in the prototype scale) and horizontalrelative displacements of the two adjacent pipes.

The short half-segments at the end of the pipeline were used fora representation of a remaining part of a much longer pipeline.Both ends of the pipeline were sealed by round plastic plates toprevent the soil from moving into the pipeline.

Loading system and configurationThe loading and measuring systems with instrumentations are

shown in Fig. 4. The loading system with a pneumatic pressurecylinder was installed on the top U-section steel beam (beamlength 900 mm and width 100 mm), while the measuring system

with linear variable differential transformers (LVDTs) was in-stalled on another U-section steel beam (beam length 900 mm andwidth 160 mm). Both beams were fixed directly to the strongbox,and there was no contact between these two separate beams, thusthe loading of the model and the deformations of the loadingbeam did not influence the measuring beam with instrumentations.

During the centrifuge testing, the pipeline was subjected tomaximum service loads in accordance with AASHTO (2010). Threelive load configurations were considered: (i) design truck singlepair of wheels at the end of an axle, (ii) two single pairs of wheelsof design trucks in passing mode, and (iii) design tandem, asshown in Figs. 5a, 5b, and 5c, respectively.

The load configuration of a single pair of wheels is representedin Fig. 5a. When the vehicle is moving in a given direction, whichis directly above the pipe but not perpendicular to the pipelineaxis, then the second pair of wheels at another end of the axlewould not be over the pipeline axis and located at some distancefrom it. Thus, the second pair of wheels would have a limitedinfluence on the pipe joint. In this situation, only a single pair of

Fig. 2. Schematic views of centrifuge strongbox: (a) diagonal view with one loading pad; (b) diagonal view with two or four loading pads;(c) plane view. Soil cover depth H = 35 or 70 mm. (All dimensions in millimetres.)

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wheels over the joint should be used as a design case, becausethis configuration induces the largest demands across the joint(Becerril García and Moore 2014).

The second configuration (ACPA 2011) was two single pairs ofwheels of design trucks in passing mode (Fig. 5b). In this configu-ration, the distance between the two truck wheels was 1219 mm(4 ft.), which is closer than the distance between two wheels ofsingle truck (1829 mm or 6 ft.). Because of the closer distancebetween the wheels, the induced demand across the tested joint ishigher and, as a result, the influence of the two single pairs ofwheels of design trucks in passing mode is expected to be moresignificant. Furthermore, if the vehicles move in a direction per-pendicular to the pipeline axis, the presence of four pairs ofwheels above the pipeline at the same time (Fig. 5d) would reducethe live load influence on the tested joint. Therefore, the config-uration of two loading pads used in tests 3 and 4 (Fig. 5b) was more

critical for the tested pipe joint, and the test results are on themore conservative (safer) side.

Design tandem load configuration is shown in Fig. 5c. It consistsof a pair of axles spaced 1219 mm (4 ft.) apart (AASHTO 2010).During the tests, the largest load was applied using this configu-ration.

For a single pair of wheels, the dimensions of the tire contactarea were width, B, of 254 mm (10 in.) and length, L, of 508 mm(20 in.) in prototype scale, which gives 12.7 mm × 25.4 mm in themodel scale (see Fig. 6). According to the scaling law for stress inTable 1, stresses beneath the loading pads in model scale wereequal to those in prototype scale. For further information on cal-culating surface loads during the centrifuge tests, refer to Rakitinand Xu (2014).

For design truck single pair of wheels, the unfactored designload was 71 kN, which was then multiplied by a multiple presence

Fig. 3. (a) Prototype of reinforced concrete pipe and (b) its aluminum centrifuge model (c) Gasketed bell and spigot joint at model scale. (d)Photographic view of aluminum pipe model with a gasket. (All dimensions in millimetres.)

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factor and a dynamic load allowance to obtain the maximumservice load (Lmax). The multiple presence factor is 1.2. The dy-namic load allowance, which accounted for the truck load beingnonstatic, is 1.24 and 1.14 for the 700 and 1400 mm soil coverdepth, respectively (AASHTO 2010). For unification, a commonload of 100 kN (in prototype scale) was applied to represent themaximum service load (Lmax) per single pair of wheels in testswith one loading pad for both soil cover depths (tests A, 1, and 2).

For tests 3 and 4 two loading pads were used to represent twosingle pairs of wheels (see Fig. 5b). The maximum service load persingle pair of wheels was also 100 kN, thus the maximum serviceload per two single pairs of wheels (Lmax) applied during the testswas 200 kN (in prototype scale).

In tests 5 and 6 four loading pads were used to represent adesign tandem load (see Fig. 5c). According to AASHTO (2010), theunfactored load per single pair of wheels was 55.6 kN. The unfac-tored load was then multiplied by a multiple presence factor anda dynamic load allowance, as described in the previous paragraph.Therefore, the maximum service load per single pair of wheels ofa design tandem load was chosen as 75 kN (in prototype scale).Because the design tandem load consisted of four pairs of wheels,the maximum service load per four single pairs of wheels (Lmax)was 300 kN (see Table 2). The surface loads applied during the testsare summarized in Table 2.

The surface live loads on underground infrastructure are con-stantly increasing because of the development of the automotiveindustry. As a result, the U.S. Federal Highway Administration(U.S. Federal Highway Administration 2000) is planning to in-crease truck weight limits from the current 36 to 44 short tons(1 short ton = 0.907 Mg).

To take into consideration the possibility of higher live loads,after the first 10 cycles of loading, the maximum service live load(Lmax) was increased by 50% and applied for 10 more cycles ofloading–unloading in tests 1–6, as shown in Figs. 7b, 7c, and 7d.

During all centrifuge tests, the surface live loads were appliedusing a pneumatic pressure cylinder installed on the loadingbeam, as shown in Fig. 4. The air pressure inside the cylindercould be changed during the centrifuge test to simulate differentmagnitudes of surface live loads. A pneumatic pressure cylinderwith a 40 mm inner diameter was used for applying a verticalforce on one and two loading pads. Because this cylinder could notapply the amount of load necessary for four loading pads, a larger50 mm inner diameter cylinder was used during tests 5 and 6.

Measuring systemFigure 2 illustrates the locations of the LVDTs installed on the

centrifuge strongbox to investigate responses of the pipes in thelongitudinal directions. The vertical displacements of the pipeswere measured using LVDTs with rigidly connected extensionrods (see Fig. 8a) because the length of the LVDTs was not enoughto reach the top of the pipes. The connections between extensionrods and the top of the pipes were flexible, to allow free piperotation (see Fig. 8b). The extension rods were encased in hollowaluminum tubes to avoid contact with the surrounding soil dur-ing the centrifuge tests. In addition, the extension rods were lu-bricated with silicon grease to minimize friction with the hollowaluminum tubes. A special recess was made on the top of the pipesto prevent any slipping of the LVDT extension rod. The hollowaluminum tubes were sealed at the bottom with soft paste toprevent disturbance from the sand.

A preliminary numerical analysis was performed to estimatethe influence of the loading pads on the LVDT extension rods. Thedistance was chosen to be sufficiently large (see Fig. 4), while thevertically installed aluminum tubes had a small diameter of only6 mm and very smooth surfaces. Therefore, the influence of theLVDT extension rods on soil movements under the loading padscould be neglected.

As the bell-and-spigot joint is not symmetrical relative to thevertical axes passing through its center, the vertical pipeline dis-placements on both left and right sides of the tested joint weremeasured. In total, 10 LVDTs installed on the measuring beamwere used during the tests (see Figs. 2a and 2b). Because the lon-gitudinal vertical displacements along the pipeline were not uni-form, LVDTs with different measurement ranges were used.

Because the largest vertical movements were expected near thetested joint, LVDTs with a measurement range of 0–20 mm and aresolution of 0.0005 mm (LVDT 1, 1=, 2, and 2=) were installed here.As the distance from the tested joint increased, the vertical move-ments of the pipeline decreased, thus, LVDTs with a smaller mea-surement range and a higher resolution were installed in thisregion. For pipes P2 and P2=, LVDTs 3, 3=, 4, and 4= (with a range0�10 mm and a resolution of 0.000 25 mm) were used. For pipes 3and 3=, which were located at the largest distance from the testedjoint and had the smallest vertical movements, LVDTs 5 and 5=(with a range of 0�5 mm and a resolution of 0.000 125 mm) wereused. The measurement accuracy of all the LVDTs used in thecentrifuge tests was 0.05%, and they were connected to the 16-bit32-channel on-arm data acquisition system, which took readingsonce every second during the tests.

The data from periods (about 2 min) during the tests beforeloading, and during the calibration when there was no change inthe applied displacement, was employed to evaluate the precisionof the measurements, which reflects the repeatability of the mea-surements. It was found that the system gave a stable reading withonly slight variation. The standard deviation was calculated foreach LVDT and was found to be less than 0.0003 mm, indicating agood precision.

Model preparationFigure 2 shows the cross section of the strong box with geomet-

rical parameters. The pipeline was tested under two different soilcover depths: 700 mm and 1400 mm (or 35 mm and 70 mm inmodel scale), which is larger than the minimum soil cover depthof 600 mm required by AASHTO (2010). Thus, the ratios of soilcover depth to pipeline inner diameter were 0.5 and 1, respec-tively. The soil depth below the pipeline was 4100 mm (or 205 mmin model scale), the ratio of which to the pipeline inner diameterwas 2.9.

The pipeline was laid on compacted materials (90% of standardProctor compaction). As the centrifuge acceleration was 20g,the soil used in the centrifuge tests corresponds to gravelly sand

Fig. 4. Strong box installed on centrifuge with loading andmeasuring systems fixed on separate and independent U-sectionsteel beams.

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(category 1) in prototype scale (ACPA 2011). Thus, AASHTO (2010)type 2 installation was employed in all centrifuge tests.

For each centrifuge test, the strongbox was filled with sandusing a dry pluviation technique to ensure a uniform density.During the model preparation, the density of the soil sample wascontrolled by the sand dropping height, which was chosen as800 mm. Sand dry unit weight was 15.7 kN/m3 and average particlesize was D50 = 0.2 mm. The maximum and minimum void ratioswere found to be emax = 0.982 and emin = 0.598, respectively. Thespecific gravity of the sand solids was Gs = 2.65. A relative density,

Dr, of 0.85 was achieved during the tests. The particle size distri-bution of the sand is shown in Fig. 9.

Before each centrifuge test, the model preparation procedure was

1. The pipeline model was assembled using four complete alu-minum model pipes and two half-segments. The pipes’ bell-and-spigot joints were sealed with compressible O-ring gaskets.To ensure test quality and repeatability, new gaskets wereused for each centrifuge test.

2. Using a dry pluviation technique, the strongbox was filledwith sand until the level of the pipeline bottom installation.

Fig. 5. Live load configurations (on prototype scale): (a) design truck single pair of wheels (one loading pad), (b) two single pairs of wheels ofdesign trucks in passing mode (two loading pads), (c) design tandem (four loading pads), and (d) two design trucks in passing modeconfiguration above pipeline. Live loads are in accordance with AASHTO (2010) and ACPA (2011).

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The voids for placing the protruding pipe bells were pre-excavated.

3. The pipeline model was laid in the strongbox. The measuringbeam with LVDTs was installed on the top of the strongbox.The LVDT extension rods encased in hollow aluminum tubeswere connected to the top of the pipes (Fig. 8a).

4. The LVDTs and their extension rods were temporarily re-moved and hollow aluminum tubes were supported by a struc-tural frame during sand pluviation. In addition, the tube endswere sealed to prevent sand particles from going inside. Plu-viation of sand continued until the sand sample reached thenecessary height.

5. The hollow aluminum tubes were unsealed. The measuringbeam with LVDTs and their extension rods was fixed on thestrongbox. The temporary structural frame for the aluminumtubes support was removed. After that, the loading beam witha pneumatic pressure cylinder and loading pads was installed.

Test program and procedureIn total, seven geotechnical centrifuge tests were carried out

(see Table 2). As shown in Figs. 2 and 5, three live load configura-tions for two typical soil cover depths of 700 and 1400 mm werestudied during the tests.

Each type of live load was applied cyclically. For the first cycle,precautions were taken and the maximum service live load (Lmax)was applied gradually in four steps (see Fig. 7), which correspondto 25%, 50%, 75%, and 100% of the maximum service load, and thenthe load was decreased with the same pattern. From cycles 2 to 10,the maximum service load was applied with the same symmetricloading–unloading pattern for each cycle. After the first 10 cycles,the maximum service live load was increased by 50% (1.5Lmax) andapplied for 10 more cycles of loading–unloading. In total, during

each centrifuge test (excluding test A), the pipeline model wastested for 20 cycles of loading–unloading: 10 cycles with maxi-mum service load (Lmax) and 10 cycles with increased maximumservice load (1.5Lmax). During test A, maximum service load (Lmax)was applied for 20 cycles without increase, as shown in Fig. 7a.

The following procedure was performed for each centrifugetest:

1. The centrifuge model was prepared in a strongbox at 1g. Afterthat, the strongbox was weighed and installed on the centri-fuge platform.

2. The centrifuge was gradually accelerated to the test accelera-tion of 20g. Before the surface live load was applied, sufficienttime was allowed to ensure that no further soil settlementoccurred. Then the data from all LVDTs measured at 20g accel-eration were taken as the initial readings.

3. The cyclical surface live load was applied in flight using thepneumatic pressure cylinder. After each loading step, suffi-cient time was provided to allow all LVDTs readings to stabi-lize before continuing to the next step of loading.

4. After all loading cycles were completed, the centrifuge wasstopped.

Results and discussion

Influence of the number of cyclesIn this research the pipe joints were tested under cyclic loading.

The first centrifuge test (test A) was performed to identify theminimum number of loading cycles required for the investiga-tion.

In test A the pipeline joints were tested under 700 mm soil coverdepth while the maximum service cyclic load (Lmax) of 100 kN (inprototype scale) was applied using a single loading pad (Fig. 2a).During the centrifuge test, the vertical pipe displacements weremonitored. The test continued for 20 cycles. Unfortunately, LVDT 3,which was located at a distance of 3 m (in prototype scale) fromthe load application point, was not functioning properly duringthe test. Thus, no useful data was obtained from this LVDT.

In total, the live load was applied for 20 cycles. The results ofthis centrifuge test are shown in Fig. 10. As can be seen from thisgraph, as the pipeline was loaded for more cycles, the increase invertical movements per cycle decreased.

Fig. 6. (a) Track tire contact areas on model scale and (b) loading pads photographic views. (All dimensions in millimetres.)

Table 2. Centrifuge tests details.

TestNo.

Number ofloading pads

Soil coverdepth, H (mm)

Total surfaceload Lmax (kN)

Total surfaceload 1.5Lmax (kN)

A 1 700 100 —1 1 700 100 1502 1 1400 100 1503 2 700 200 3004 2 1400 200 3005 4 700 300 4506 4 1400 300 450

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After Fig. 10 in all following graphs, all dimensions are given inprototype scale, unless mentioned otherwise.

The graphs in Fig. 11 display the data from LVDT 1 with maxi-mum vertical displacement, which is located 1 m right (in proto-type scale) from the center of the load application point. After thefirst cycle of loading, the vertical displacement per one cycle was

the largest, equal to 0.52 mm (see Fig. 11a). During the 10th cycleof loading, the accumulated displacement was 1.23 mm. Whenthe surface load was applied for 10 more cycles (from 11th to20th cycles), the vertical displacement increased from 1.23 to 1.5 mm.

Figure 11b shows the relationship between vertical displace-ment and the number of surface load applications. As can be seen,the average vertical displacement curve became much flatter afterthe first 10 cycles of loading. Therefore, 10 cycles of surface loadapplication are deemed to be sufficient for investigation of thecyclic behavior of pipe joints.

Fig. 7. Applied live load versus number of cycles for (a) one loadingpad during test A, (b) one loading pad during tests 1 and 2, (c) twoloading pads during tests 3 and 4, and (d) four loading pads duringtests 5 and 6.

Fig. 8. (a) Photographic view of centrifuge strongbox with pipelinemodel during preparation. (b) Flexible LVDTs extension rodconnection with the top of the pipes.

Fig. 9. Sand particle-size distribution.

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Influence of the load configurationIn tests 1, 3, and 5, the soil cover depths were all the same (i.e.,

700 mm). However, the live load configurations were differentand they were applied in accordance with Fig. 5 and graphs inFig. 7. Thus, the influence of different kinds of traffic loads on pipe

joint behavior can be studied. These tests results are presented inFigs. 12, 13, 14, and Table 3.

In test 1 (Fig. 12), the live load was applied using one loading pad(i.e., design truck single pair of wheels). After the first cycle of100 kN loading (Lmax), the maximum vertical displacement of thepipe joint was 0.75 mm. After 10 cycles of loading, the verticaldisplacement increased by nearly two times to 1.35 mm. When themaximum service live load was increased by 50% (1.5Lmax), thevertical displacement further increased to 1.5 mm. After 10 cyclesof loading with the increased load, the final vertical displacementwas 2.05 mm.

For the first 10 cycles of loading, both tests A and 1 were per-formed under the same soil cover depth and the same surfaceloading conditions. Thus, the data from these two tests can becompared directly to evaluate the repeatability of the test data.After the first cycle of loading, the maximum vertical displace-ment of the pipe joints was 0.73 mm for test A and 0.75 mm fortest 1 (the difference between the data of the two tests is 3%). After10 cycles of loading, the maximum vertical displacement became1.44 and 1.35 mm, respectively (difference of about 6%). Verticaldisplacements measured by other LVDTs during tests A and 1 werealso similar. Therefore, the repeatability of the test data are en-sured.

Furthermore, the results of test 1 were compared with thosepublished by Becerril García and Moore (2015), who carried outfull-scale laboratory tests on a 1200 mm inner diameter RC pipe-line with a slightly shallower soil cover depth (600 mm) and abedding depth of 1.17 m, subjected to a maximum surface load of100 kN. The 6.6 m long pipeline consisted of two complete pipesand two segments cut from a third complete pipe, with flexibleretaining wall as the boundary. A maximum vertical joint dis-placement of 0.21 mm was observed in their test. The maximumdisplacement in test 1 during the first cycle was larger (i.e.,0.75 mm), which is probably due to the larger bedding thickness(4100 mm in prototype scale), as larger soil compression wouldhappen under the same surface load. A longer pipe length (12 m in

Fig. 10. Vertical displacements of pipeline under maximum servicecyclic load of 100 kN applied for up to 20 cycles (test A). (Alldimensions in millimetres and in prototype scale.)

Fig. 11. Vertical displacements at LVDT1 of pipeline under 20 cyclesof loading (test A): (a) vertical displacement over the surface load,and (b) number of cycles over the vertical displacement. (Alldimensions in millimetres and in prototype scale.)

Fig. 12. Vertical displacements of pipeline with 700 mm soil coverdepth loaded with one loading pad (test 1). (All dimensions inmillimetres and in prototype scale.)

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prototype scale) in the centrifuge test might also contribute to alarger maximum displacement. However, the general pattern ofthe tested pipe joint movements (the joint between pipes P1 andP1=) is similar.

In test 3 (Fig. 13) the live load was applied using two loading pads(design trucks two single pairs of wheels in passing mode). During

this test, the maximum vertical displacement of the joint was onaverage 2.6 times higher than in the previous test with one pad(see Table 3). In test 5 (Fig. 14) the live load was applied using fourloading pads (design tandem). On average, the maximum verticaldisplacement of the joint was 3.85 times higher than in the testwith one pad.

During centrifuge tests 1, 3, and 5, the pipe joint behavior wasstudied under the same soil cover depth but with different liveload configurations. In Table 3, the results from these tests aresummarized and presented for four different cycles of loading. Inaddition, the ratios of the maximum vertical displacements be-tween different tests are also given in Table 3. For example, thevertical displacement ratio of test 3 / test 1 represents the relation-ship between the maximum vertical displacement in tests 3 and 1.Furthermore, the average vertical displacement ratio of test 3 /test 1, shown at the bottom of the table, represents the arithmeticmean value of the four vertical displacement ratios. It can be seenthat, when the data from these three tests were compared, designtandem (four loading pads) was found to be the least favorablecondition due to the fact that the maximum vertical displace-ments of the pipe joint were on average 3.85 times larger than thatof the design truck single pair of wheels (one loading pad).

Influence of soil cover depthTo study the influence of the soil cover depth on pipe joint

behavior, tests 2, 4, and 6 were carried out with a soil cover depthof 1400 mm. The results are compared with those of tests 1, 3, and5, in which the soil cover depth was 700 mm (see Table 1).

Figure 15 presents the vertical displacements of the pipelineduring different stages of loading in test 2. Tests 1 and 2 wereperformed under similar surface loading conditions, but differentsoil cover depths (700 and 1400 mm, respectively). Comparing theresults in tests 1 and 2, the average vertical displacement ratio oftest 2 / test 1 was 0.72 (see Table 3). Thus, significant increase in thesoil cover depth resulted in a relatively small decrease of thetested joint’s maximum vertical displacement.

In test 4 (see Fig. 16) two loading pads were used and the pipe-line was tested under 1400 mm soil cover depth. The results arecompared with those from test 3 with different soil cover depths.It is noticed that in these two tests the increase in the soil coverdepth resulted in a significant decrease in the tested joint’s max-imum vertical displacement: the average vertical displacementratio of test 4 / test 3 was 0.35 (see Table 3).

In test 6, four loading pads were used and the soil cover depthwas 1400 mm (for the test data, see Fig. 17). When the data fromthis test were compared with the data for test 5 (four pads loading,but 700 mm soil cover depth), the average vertical displacementsratio test 6 / test 5 was 0.53 (Table 3).

In summary, the influence of soil cover depth on pipe jointbehavior varied under different kinds of live load. For one loadingpad, when a soil cover depth was increased from 700 mm to1400 mm, the average vertical displacement ratio of test 2 / test 1was 0.72. For two and four loading pads, the average vertical dis-placement ratios of test 4 / test 3 and test 6 / test 5 were 0.35 and0.53, respectively.

Thus, the decrease in pipe joint movement associated with in-creased soil cover depth was most significant for the case of twopads loading (i.e., two single pairs of wheels of design trucks inpassing mode), and least significant for the case of one pad load-ing (i.e., design truck single pair of wheels).

Influence of surface load on other pipes of the systemOne of the main questions this research set out to address is

how far along the pipeline the vertical movements caused by thesurface live loading could be transferred in a pipeline with gas-keted bell-and-spigot joints. As shown in Fig. 2, the pipeline de-scribed in this paper consists of four complete pipes and twohalf-segments. The pipeline total length was 12 m in prototype

Fig. 13. Vertical displacements of pipeline with 700 mm soil coverdepth loaded with two loading pads (test 3). (All dimensions inmillimetres and in prototype scale.)

Fig. 14. Vertical displacements of pipeline with 700 mm soil coverdepth loaded with four loading pads (test 5). (All dimensions inmillimetres and in prototype scale.)

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scale. Data from tests 1–6 during the 10th cycle of the maximumservice live loading (Lmax) were used to study the transfer of verti-cal movements in a pipeline (see Fig. 18).

As can be seen from Fig. 18 and from data in Table 4, the verticaldisplacements of the two half-segments (P3 and P3=) were notinfluenced significantly by the amount or type of live loading, asthe measurements made by LVDTs 5 and 5= are small and constantduring all the tests. Thus, in Fig. 18, the vertical displacementpoints for LVDTs 5 and 5= are located close to each other.

Table 3. Maximum vertical displacement during tests.

Maximum vertical displacement ratio

Maximum vertical displacement (mm) Loads influence Soil cover influence

Cycle of loading Test 1 Test 2 Test 3 Test 4 Test 5 Test 6Test 3Test 1

Test 5Test 1

Test 2Test 1

Test 4Test 3

Test 6Test 5

Lmax–1st 0.75 0.55 2 0.68 3.5 1.8 2.67 4.67 0.73 0.34 0.51Lmax–10th 1.35 0.92 3.1 1.15 4.8 2.5 2.3 3.56 0.68 0.37 0.521.5Lmax–1st 1.5 1.18 4.3 1.45 5.7 3.2 2.87 3.8 0.78 0.34 0.561.5Lmax–10th 2.05 1.40 5.2 1.72 6.9 3.62 2.54 3.37 0.68 0.33 0.52

Average* 2.6 3.85 0.72 0.35 0.53

Note: Lmax–1st, first cycle of maximum service live load application (Lmax); Lmax–10th, 10th cycle of maximum service live load application (Lmax); 1.5Lmax–1st, firstcycle of increased maximum service live load application (1.5Lmax); 1.5Lmax–10th, 10th cycle of increased maximum service live load application (1.5Lmax).

*Average vertical displacement ratio, which represents an arithmetic mean value of the four vertical displacement ratios of the different cycles of loading.

Fig. 15. Vertical displacements of pipeline with 1400 mm soil coverdepth loaded with one loading pad (test 2). (All dimensions inmillimetres and in prototype scale.)

Fig. 16. Vertical displacements of pipeline with 1400 mm soil coverdepth loaded with two loading pads (test 4). (All dimensions inmillimetres and in prototype scale.)

Fig. 17. Vertical displacements of pipeline with 1400 mm soil coverdepth loaded with four loading pads (test 6). (All dimensions inmillimetres and in prototype scale.)

Fig. 18. Vertical displacements of pipeline during the 10th cycle ofthe maximum service live load (Lmax) application in tests 1–6. (Alldimensions in millimetres and in prototype scale.)

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In contrast with P3 and P3=, the data from LVDTs 3, 3=, 4, and 4=showed that the vertical displacements of P2 and P2= were influ-enced significantly by the type and amount of surface live loading.It can be also seen in Fig. 18 that the vertical displacements of P2and P2= were different during tests 1–6.

During tests 1–6, the vertical displacements of the P1 and P1=pipes, whose joint was tested, were influenced greatly by the typeand amount of surface live loading. Since shear displacement isevident in the joints between P1 and P2, and between P1= and P2=,this suggests that P2 and P2= provide significant support to P1 andP1= through shear force transfer in the joints between them.

In conclusion, two pipes on the right side (P1 and P2) and twopipes on the left side (P1= and P2=) of the tested joint are influencedsignificantly by the surface live loading. The influence of the sur-face live loads on P3 and P3= is relatively small, which indicatesthat P3 and P3= provide much less support to P2 and P2=.

Pipeline movements under surface loadsIn this research, the vertical displacements of four complete

pipes and two half-segments were measured. Based on the exper-imental data, it was noticed that the influence of load on thepipeline was not symmetrical relative to the vertical axis comingthrough the center of the tested pipe joint. As shown in Fig. 19, theinfluence of live loads can extend in two directions: the bell-to-spigot direction (i.e., to the right from the tested joint) and thespigot-to-bell direction (i.e., to the left).

Based on the data in Fig. 18 and Table 4, it can be seen that thevertical displacements of points measured by LVDTs 1, 2, 3, 4, and5 in the bell-to-spigot direction were generally larger than verticaldisplacements of points measured by LVDTs 1=, 2=, 3=, 4=, and 5= inthe spigot-to-bell direction. Clearly, the vertical movements in

opposite directions from the tested joint were different and notsymmetrical.

In Fig. 20, typical results of vertical displacements of the pipe-line during the 10th cycle of the maximum service live load (Lmax)application in tests 5 are shown. Larger differential displacementwas observed between pipes P1= and P2= in the spigot-to-bell direc-tion (0.8 mm), in comparison with that between P1 and P2 in thebell-to-spigot direction (0.3 mm). As higher differential displace-ment results in higher shear force demand on the joint, the pos-sibility of the leakage in the joint between P1= and P2= is thereforehigher.

Rotation at the joint is calculated as the change in angle be-tween the central axes of the two adjacent pipes. The largest piperotation was generated in the pipe joint between P1 and P1= duringthe live loading (see Fig. 20). Higher leakage resistance is thereforerequired for the central joint.

In summary, the vertical movements of large-diameter rein-forced concrete pipelines with gasketed bell-and-spigot jointsunder surface loading are nonsymmetrical and discontinuous,involving significant concentration of shear and rotation at thejoints.

Summary and conclusionsThis paper presents the results of a geotechnical centrifuge

testing program in which the response of a 12 m long (in prototype

Table 4. LVDTs data for 10th cycle of maximum service live loading.

TestNo. LVDT 5= LVDT 4= LVDT 3= LVDT 2= LVDT 1= LVDT 1 LVDT 2 LVDT 3 LVDT 4 LVDT 5

A −0.3175 −0.4200 −0.615 −0.890 −0.960 −1.220 −0.960 −– −0.690 −0.61001 −0.2450 −0.2850 −0.420 −0.730 −0.820 −1.070 −0.790 −0.495 −0.455 −0.42502 −0.2025 −0.2900 −0.440 −0.500 −0.570 −0.800 −0.690 −0.555 −0.505 −0.45253 −0.2625 −0.4007 −0.6531 −1.820 −2.050 −2.260 −1.990 −0.935 −0.605 −0.43754 −0.2725 −0.3500 −0.4550 −0.760 −0.830 −0.980 −0.940 −0.685 −0.605 −0.54505 −0.2375 −0.4300 −0.7500 −2.050 −2.560 −2.830 −2.170 −0.995 −0.635 −0.41006 −0.2675 −0.4550 −0.8100 −1.430 −1.570 −1.750 −1.470 −0.850 −0.795 −0.5600

Fig. 19. Directions of live loads influence. Fig. 20. Vertical displacements of pipeline during the 10th cycle ofthe maximum service live load (Lmax) application in test 5. (Alldimensions in millimetres and in prototype scale.)

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scale) large-diameter reinforced concrete pipeline with gasketedbell-and-spigot joints subjected to three standard AASHTO designload configurations has been investigated.

The influence of the number of load cycles on pipe joint behav-ior was analyzed. The results show that most vertical pipe dis-placement occurred during the first 10 cycles of traffic loading,while further increase in the number of load cycles did not lead toa significant increase in pipe joint movements.

The influence of three load configurations has been investi-gated. Under design tandem loading (four loading pads), the pipejoint movements were significantly higher than those under theother two traffic load configurations (e.g., 3.85 times higher thana single pair of wheels).

An increase of soil cover depth from 700 to 1400 mm resulted inreduced pipe joint movements under surface loading and an over-all improvement in pipe joint service conditions. For one padloading, the average vertical displacement ratio of test 2 / test 1was 0.72. For two and four pads loading, the average vertical dis-placement ratios of test 4 / test 3 and test 6 / test 5 were 0.35 and0.53, respectively. Therefore, the effect of increasing soil coverdepth was most significant for two pads loading (i.e., two singlepairs of wheels of design trucks in passing mode).

The influence of surface load can be transferred to other pipesin the system through shear forces in the joints. Two pipes on theleft side and two pipes on the right of the tested joint are influ-enced significantly by the surface load. The influence of surfaceload on the third pipes was relatively small compared with thepipes closer to the tested joint.

The pipeline movements were not symmetrical. Larger differ-ential displacement of the pipes in the spigot-to-bell direction wasobserved when compared with those in the bell-to-spigot direc-tion. Therefore, the shear force demand and overall possibility ofleakage in the joints are different. Although the joint directlyunder the load experienced the largest rotation, the possibility ofleakage in the second joint in the spigot-to-bell direction was alsohigh, due to large differential displacement between the pipes.

During the geotechnical centrifuge tests presented in this pa-per, the surface load was applied directly above the pipe joint,which experienced the maximum rotation. Further researchwould be necessary to examine the response of the pipeline sub-jected to surface load at other locations, which might cause max-imum shear demand to the joint.

AcknowledgementsThe authors are grateful for the research support received from

the National Natural Science Foundation of China (51450110097,41272280, and 51350110231). The assistance of Dehai Jin and DaweiShen with conducting the centrifuge testing, and the manufactur-ing of the pipeline model and loading pads by Hongxing Zhu, isalso gratefully acknowledged.

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