Importance of sidewall friction on manhole uplift during soil liquefaction

Soil Dynamics and Earthquake Engineering 119 (2019) 51–61

Contents lists available at ScienceDirect

Soil Dynamics and Earthquake Engineering journal homepage: www.elsevier.com/locate/soildyn

Importance of sidewall friction on manhole uplift during soil liquefaction ⁎

T

Zhiyong Zhang , Siau Chen Chian Department of Civil and Environmental Engineering, National University of Singapore, Singapore 117576, Singapore

A R T I C LE I N FO

A B S T R A C T

Keywords: Centrifuge modeling Earthquake Liquefaction Sidewall friction Uplift remediation

The friction at the interface of a manhole’s sidewall and soil is often ignored with the assumption of a fully liquefied state of granular soil. A detailed analysis of the component forces acting on a conventional manhole was carried out using centrifuge modeling. From the analysis, the sidewall friction at the interface of manhole sidewall and soil was determined and found to represent a substantial portion of residual resistance inhibiting manhole uplift despite its low magnitude relative to other component forces such as buoyancy, seepage force and self-weight of the manhole at static condition. In order to validate the importance of sidewall friction, further centrifuge tests with differing sidewall friction using sandpaper were conducted to demonstrate its contribution to the role of manhole uplift following soil liquefaction.

1. Introduction

fluid density and viscosity [12]. Based on this assumption, empirical Eq. [13] and safety factor [14] were proposed. However, there exists difficulties in determining the viscosity of liquefied soil which can vary between as low as 0.1 Pa-sec to 10,000 Pa-sec [15], thereby rendering the predicted uplift displacement to deviate significantly. Besides, the dynamic soil-structure interaction is also affected by factors such as the groundwater table and the self-weight of the structure, both of which are difficult to be considered in the viscous fluid analogy. Other researchers suggested considering the action on the underground structure invert in two parts: the hydrostatic buoyancy, which can be directly computed based on Archimedes Principle, and the dynamic uplift force due to the generation of the measurable excess pore pressure [11,16]. With this idea, another concept of safety factor was also proposed as shown in Eq. 1 [11]:

During a strong earthquake event, ground motions lead to a significant loss of saturated granular soil shear strength due to the decrease in effective stress, thus leading to soil liquefaction. Severe ground liquefaction was observed in recent major earthquakes, such as the 1999 Chi-Chi Earthquake [1,2], 2010–2011 Canterbury Earthquake Sequence [3] and 2011 Great East Japan Earthquake [4]. Due to lower unit weight as compared with the surrounding soil, buried structures such as manholes, underground tanks, pipelines and tunnels possess an inherent buoyancy which can lead to uplift when the soil liquefies. Damage to lifelines was observed to be the next most notable feature after landslides following the 2004 Niigata-Ken Chuetsu earthquake in Japan [5], and brought substantial financial losses and inconveniences to residents and businesses. In particular, manholes, having large crosssectional areas, are highly vulnerable to uplift. Such uplift of manholes has been observed as early as the 1964 Niigata Earthquake and 2004 Niigata-Ken Chuetsu Earthquake [6]. To date, the uplift failure of these lifelines continue to persist even in developed countries with robust seismic practices such as New Zealand and Japan as evident in recent events such as the 2010–2011 Christchurch Earthquakes [7] and 2011 Tohoku earthquake [8], indicating that this is still a pressing issue. Researchers in the past couple of decades had been focusing on the uplift failure of underground structures with significant efforts paid to shed light on the fundamental uplift mechanism and prediction of uplift magnitude [2,9–11]. Several methodologies were put forward. One idea was to assume the liquefied soil as a viscous liquid and the uplift mechanism alike to the classic drag equation in fluid dynamics considering

⁎

Fs =

FM + FSH FEPP + FB

(1)

where FM is the manhole’s self-weight; FSH is the shear resistance on the sidewall; FEPP is the seepage force acting on the manhole invert due to the generated excess pore pressure; FB is the buoyant force due to hydrostatic water pressure. This concept of safety factor was adopted by several researchers [2,9,17,18]. When computing the safety factor in their studies, the wall frictional resistance in the liquefied layer was often ignored, and the safety factors were reported below 1, particularly between 0.7 and 0.9, at the onset of uplift of their underground structures. Safety factor of 0.7 and 0.9 implies that the sum of the resistance forces is 30% and 10% lower than the sum of uplift forces respectively. This indicates that the

Corresponding author. E-mail addresses: [email protected] (Z. Zhang), [email protected] (S.C. Chian).

https://doi.org/10.1016/j.soildyn.2018.12.028 Received 21 August 2018; Received in revised form 28 December 2018; Accepted 28 December 2018 0267-7261/ © 2018 Elsevier Ltd. All rights reserved.

Soil Dynamics and Earthquake Engineering 119 (2019) 51–61

Z. Zhang, S.C. Chian

computation of resistance and uplift forces are incomplete according to law of equilibrium. Another tool to study the uplift problems is numerical simulation. However, due to the deficiency of current numerical platforms and constitutive models, simulation of liquefaction induced deformation still remains a challenge [19]. Some researchers have attempted to study the underground structure uplift failure numerically [20,21], but extensive calibration of parameters based on experiments are essential prior to conducting the simulation and only qualitive conclusions were yielded. Although many researchers working on this topic strive to propose possible remediation methods against uplift failure and predict the uplift displacement, little work has been carried out to investigate the contributions of the component forces acting on the manhole during earthquake induced soil liquefaction and verify their relationships with the manhole’s onset and cessation of uplift, both of which are the crux of the uplift mechanism. This absence of detailed understanding of uplift mechanism has therefore hampered the effectiveness of remediation methods proposed which yielded varying degree of successes in many occasions. In this paper, following a conventional manhole experiment, results of another three centrifuge experiments with different manhole wall roughness are presented to ascertain the contribution of wall frictional resistance. The altering of the manhole sidewall roughness is achieved with sandpapers of different grit, which their frictional angles with the soil are measured with a modified direct shear test setup. A comprehensive force equilibrium analysis is thereafter carried out and validated with the observed uplift response of the manhole. After that, comparison between the centrifuge tests with different manhole wall roughness would be conducted and findings are substantiated with the theoretical force analysis to determine the influence of sidewall roughness on manhole uplift failures in liquefied soil. 2. Centrifuge modeling

Fig. 1. Prototype manhole and model manhole (mm).

The geotechnical beam centrifuge at the National University of Singapore (NUS) Centrifuge Laboratory has an in-flight platform radius of approximately 2 m and a capacity of 40 g-tons. Geotechnical centrifuges can realistically recreate the response of liquefiable soil, associated with the identical stress and strain states in the prototype according to the scaling law [22]. With miniaturized sensors, it is feasible to monitor the variations of the model deformation, pore water pressure, and acceleration during earthquake shaking. The centrifuge glevel experienced by the soil model was 25 g. A simple sinusoidal shaking with amplitude of 0.2 g, frequency of 1 Hz for a duration of 25 s in prototype scale was applied to simulate the earthquake shaking. In order to eliminate the boundary effect and offer a semi-infinite extent alike to field condition, the test model was prepared in a laminar box, which is made up of rectangular rings supported by ball bearings that allows the box to deform corresponding to the natural response of the soil.

Table 1 Comparison of model and prototype parameters. Parameter

Prototype scale

Required model scale

Fabricated model manhole

Length (cm) Diameter (cm) Wall thickness (cm) Mass (g) Volume (cm3) Density (g/cm3)

300 110 7.5 2.29 × 106 2.43 × 106 0.94

12 4.4 0.3 146 156 0.94

12 4.4 0.3 176 182 0.96

2.1. Centrifuge setup Fig. 1 shows a prototype manhole typically used in Japan and simplified manhole model. The model manhole used in the centrifuge test was made of aluminium alloy, and assembled with a cap, cylindrical wall and a base. The dimensions of the model manhole were scaled 1/25 times of the prototype manhole. Table 1 shows the detailed comparison of specifications between the model and prototype manhole. The sand (W9 sand) used in the tests was silica in nature and derives from the Logan River in Australia. It is a uniformly graded sand with properties ϕ =34° , emax= 0.86, emin= 0.49, Gs= 2.65. The grain size distribution is shown in Fig. 2. When preparing the sand model, the uniformity and reproducibility was ensured by means of air pluviation using a semi-automatic sand pourer. By controlling the sand pourer’s

Fig. 2. Grain size distribution of W9 sand and ranges of liquefiable soil [23].

52

Soil Dynamics and Earthquake Engineering 119 (2019) 51–61

Z. Zhang, S.C. Chian

Fig. 3. A dry pluviated sand model with embedded manholes and instrumentation.

Fig. 4. Centrifuge test model layout of MH-Phi-20-0(unit: mm).

2.3. Experimental results and analysis

drop height, travel speed and flow rate, a consistent sand relative density of 48% was achieved. A vacuum cleaner was used to remove excess sand and produce a level soil surface for the convenience of placing instrumentations and the instrumented manhole in the soil model. Fig. 3 shows a picture of a prepared sand model with two buried manholes. One conflict in dynamic centrifuge is the time scale between seepage and dynamic. To address this conflict, the pluviated sand model is saturated with a viscous solution, prepared with Kelco-Crete at a viscosity of 22.3 cP, 25 times the viscosity of water at 25 degrees Celsius [24]. The viscous fluid is introduced into the dry soil model from the bottom of the box with a designed saturation system, which can provide an 85% vacuum environment and thus ensure a high saturation degree in the sand sample. During the saturation process, care must be taken to ensure that the hydraulic gradient is not excessively high to cause disturbance to the model. The maximum flow rate during saturation can be determined by Darcy’s law [25].

2.3.1. Test results Experiments related to the conventional underground structures have been carried out by some researchers [11,26,27]. As observed in the field, the most typical failure of a conventional manhole due to earthquake loading is uplift. From Fig. 5, it can be observed that, as soon as the shaking started, the conventional manhole displaced upward steeply for the first 5 s shaking before transiting to a lower uplift rate with time. After the shaking ceased, excess pore pressure dissipated, and the manhole settled with the ground. These observations are similar to those reported in past experiments with similar underground structures [11,27]. The maximum excess pore pressure measured beneath the manhole (P2) was lower than that measured in far-field at the same soil depth (P3). This is attributed to the lower overall unit weight of the manhole (0.96 g/cm3) as compared to the surrounding saturated sand (2.01 g/ cm3), hence leading to a lower overburden stress at P2.

2.2. Experimental program

2.3.2. Component forces calculation Fig. 6 shows the component forces acting on a manhole after the soil has liquefied. They are the self-weight (FM ), the buoyancy (FB ), the sidewall frictional resistance (FSH ) and the seepage force (FEPP ) at the manhole structure invert. During the earthquake shaking, the four component forces has a relationship as shown in Eq. 1. If the combined downward forces (FM + FSH ) are insufficient to resist the sum of upward forces (FB + FEPP ), the net force (Fnet ) obtained from Eq. 2 would be greater than zero (i.e. net upward force), which causes the uplift of manhole.

In this study, four cases of manhole tested in the centrifuge were introduced below. The first case, labelled as MH-Phi-20-0, involves a conventional manhole with water table at the ground surface. The second manhole case, MH-Phi-20-1, was similar to the first case, except that the water table was located at 1 m below the ground surface. The remaining manhole cases (MH-Phi-30-1 and MH-Phi-32-1) were identical with MH-Phi-20-1 apart from their roughened sidewalls (the soil -sidewall friction angles were 30° and 32° respectively). All the cases in this study are tabulated in Table 2. The labelling of test indicates the friction angle and depth of water table level, i.e., MH-Phi-20-0 refers to a manhole with friction angle of 20° and water table at ground. Fig. 4 shows the arrangement of the manholes and instrumentation. Laser displacement transducers were hung over the manholes and the ground surface to obtain their vertical displacement time history during shaking. Pore pressure transducers and accelerometers were placed inside and outside the manhole at different depths in the soil model.

Soil-sidewall interface friction angle (°)

Apparent density (g/cm3)

Groundwater level (m)

MH-Phi-20-0 MH-Phi-20-1 MH-Phi-30-1 MH-Phi-32-1

20 20 30 32

0.96 0.96 0.96 0.96

0 −1 −1 −1

(2)

Fnet = FB + FEPP − FSH − FM

(3)

In this section, the computation of component forces acting on the manhole along with the shaking are presented. A typical manhole with water Table 1m below the ground surface is consider here. Fig. 7 shows the positions of the manhole and nearby instrumentation before and after earthquake shaking according to observations during experiments. In order to confirm the final location of the pore pressure transducers, the soil model was excavated carefully to expose their locations after the centrifuge test was completed. In Fig. 7, s , w and s1 are used to depict the changes in ground settlement, depth of water table, and the location of the pore pressure transducers accordingly. Fig. 8 shows the time history of the normalized ground settlement, which is computed as the instantaneous ground settlement divided by the final ground surface settlement after earthquake. The time history of this normalized ground settlement during test is represented

Table 2 Centrifuge test program. Case

FSH + FM ≤ FB + FEPP

53

Soil Dynamics and Earthquake Engineering 119 (2019) 51–61

Z. Zhang, S.C. Chian

Fig. 5. Layout and results of MH-Phi-20-0.

w (t ) = w∙f (t )

as f (t ) for the convenience of acquiring the variation in groundwater table (w (t )) and location of pore pressure transducer (s1 (t )) during the earthquake shaking. Water in the voids are expelled as the ground densifies. Since the model was fully saturated below the groundwater table, it can be assumed that the volume of expulsed pore water was proportional to the volume of ground settled. Thus, the variation in the groundwater table with time (t) can be expressed as:

(4)

Besides, to avoid any possible disturbance, the pore pressure transducer at the invert of the manhole was not attached to the manhole in the centrifuge tests and it was observed to separate from the manhole when the latter uplifted. The transducer also settles with the soil over time following the onset of the earthquake, as shown in Eq. 5:

s1 (t ) = 54

H−h ∙s∙f (t ) H

(5)

Soil Dynamics and Earthquake Engineering 119 (2019) 51–61

Z. Zhang, S.C. Chian

considered as fully liquefied, and thus the deductions should include two parts: 1) the pressure due to the water depth change of P2 (the second item in Eqs. 6) and 2) the effective weight of the suspended sand filling in to the gap (the third item in Eq. 6). Therefore, the excess pore pressure acting on the manhole base can be obtained as shown in Eq. 6:

pexc = pm − ρW ∙g∙ [s1 (t ) + w (t )] − (ρsat − ρw ) g (U (t ) + s1 (t ))

(6)

where pm is the excess pore pressure measured by P2, pexc is the excess pore pressure acting on the manhole invert, U (t ) is the time history of the manhole uplift during test, ρW is the density of water, and g is the gravitational acceleration. The buoyancy (FB ) is calculated based on Archimedes Principles, where the uplift force equals the submerged volume of the manhole multiplied by the unit weight of water. Due to the earthquake loading, the soil undergoes settlement and the groundwater level is also raised. Considering the variation of the groundwater table and manhole uplift, the instantaneous buoyancy (FB ) can therefore be computed based on the formula:

Fig. 6. Force components acting on a manhole during soil liquefaction.

FB (t ) = ρw g [L − U (t ) − (hw − w (t ))] πr 2

(7)

where ρw is the density of water, h w is the initial groundwater depth, r is the radius of the manhole base, and L is the manhole length. The calculation of seepage force (FEPP ) acting on the manhole invert can be carried out through Eq. 8, in which the excess pore pressure is assumed to be uniformly distributed across the manhole invert.

FEPP = πr 2∙pexc

(8)

For a case where the groundwater table is below ground surface, the frictional resistance along the manhole sidewall (FSH ) is obtained as the sum of wall-soil frictional resistance in the dry and saturated soil regions. The calculation is based on horizontal earth pressure at rest, which utilizes the modified K0 factor proposed by Jaky [28]. To simplify the calculation, the contact property of the soil-sidewall is assumed as unchanged during liquefaction. This is aligned with the existing component forces computation method in literatures [29–32]. Besides, the vertical length of the sand-structure contact decreases, as shown in Fig. 7, due to 1) the manhole uplift and 2) the ground surface settlement during the shaking. They should be considered during the calculation of the sidewall friction above and below groundwater table accordingly. Therefore, the frictional resistance of the manhole sidewall at the dry soil layer (FSH − dry ) is calculated as shown in Eq. 9:

Fig. 7. Elevation of manhole and ground before and after shaking.

FSH − dry (t ) =

∫0

[hw − w (t ) −s(t)]

ρd gzdz∙2πr∙K 0tanδ

(9)

where ρd is the density of the dry sand, s (t ) is the time history of the ground surface settlement, z is the integrated depth, K 0 is the coefficient of lateral earth pressure, and δ is the friction angle of interface between soil and the manhole sidewall. In the case of the frictional resistance at the saturated soil layer (FSH − wet ) , past studies have largely ignored the contribution as explained earlier. However, in this present study, the friction on the sidewall is computed based on the measured excess pore pressure ratio of the surrounding soil of the manhole. The manhole sidewall friction below the water table can be obtained by Eqs. 10 and 11.

Fig. 8. Time history of normalized ground settlement.

where h is the depth of the transducers embedded, H is the model height in prototype scale, and s is the final settlement of ground surface. As shown in Fig. 7, due to the uplift of the manhole, a gap would be developed between the pore pressure transducer (P2) and the manhole base along with the shaking. The pressure due to the elevation difference caused by this gap should hence be deduced from the measured pore pressure reading ( pm ) to reflect the excess pore water pressure ( pexc ) acting on the manhole base. In light of the evident movements of the manhole and sand around the manhole base, this region is

FSH − wet (t ) =

[L − U (t ) − s (t )]

∫[h

w − w (t ) −s(t)]

σ ′dz∙2πr∙K 0 ∙tanδ (1−ru (t ) ⎤ ⎥ ⎦

σ ′ = ρd g (h w − w (t )) + (ρsat − ρw ) g [z − (h w − w (t ))] where ru (t ) is the excess pore water pressure ratio, and ρsat is the density of the saturated soil. 55

(10) (11)

Soil Dynamics and Earthquake Engineering 119 (2019) 51–61

Z. Zhang, S.C. Chian

the earthquake. The seepage force (FEPP ), as one of the uplift-promoting forces, was generated as a result of the build-up of excess pore pressure at the manhole invert. Besides, the buoyancy force acting on the manhole (FB ) shrank slightly as the manhole displaced upward as the submerged volume of manhole below the water table was reduced. For the case of frictional resistance along the manhole sidewall beneath the water table (FSH − wet ) , it was significantly reduced since a liquefied soil has minimal shear strength to provide the resistance. However, following the aforementioned calculation procedure, a small amount of frictional resistance still existed when the soil was liquefied. By summing up these component forces, the resultant force acting on the manhole is shown in Fig. 9. It can be observed that, when the resultant force turned positive (i.e., upward), the manhole commenced its rapid uplift. As soon as the shaking ended, the resultant force (i.e., net upwards force) fell below 0 (i.e., downward), and the manhole ceased uplift despite full liquefaction still retained in the soil. This is similar to the observation of tunnel uplift by Chian, Tokimatsu and Madabhushi [27]. In addition, as aforementioned in the introduction, the factor of safety calculated in the past literatures were all below 1 (i.e., 0.7–0.9) with the precondition that they ignored sidewall friction below groundwater Table [2,9,17,18]. In order to demonstrate the importance of considering sidewall friction, this study presents both ‘factor of safety’ curves with and without considering the sidewall friction in Fig. 9. It can be seen that the factor of safety at the onset of manhole uplifting in this study was also below 1 (i.e., 0.8) when the frictional resistance on the sidewall is neglected. However, the factor of safety hovers around 1 in most of the shaking course when the sidewall friction is considered. Therefore, it can be speculated that ignoring the sidewall friction in these literatures was one of the possible reasons for failing to attain a higher factor of safety nearer to the value of unity at 1. This claim would be verified in the next section. 3. Impact of resultant force on manhole uplift response Accurate prediction of uplift displacement of manholes in the field is still extremely challenging despite recent advancements in experimental and numerical techniques. This section seeks to confirm the theoretical factor of safety at equilibrium to be 1 during the earthquake and the difficulty of using the safety factor (or the resultant force) to predict the extent of uplift displacement. Fig. 10 demonstrates the uplift time history of MH-Phi-20-0. A fitting curve is applied, and the matching formula is expressed in Eq. 14. This equation composes of two parts: 1) the natural logarithm function (s1) to describe the backbone curve in Eqs. 15, and 2) the cosine (s2) for the fluctuation in Eq. 16. Fig. 9. Forces acting on the conventional manhole (MH-Phi-20-0).

Lastly, the self-weight(FM ) is the weight of the manhole which is assumed to be unchanged during shaking. It is calculated based on the formula:

FM = ρm ∙g∙πr 2L

(12)

where ρm is the apparent density of the manhole. With the above-introduced component forces, the net resultant force experienced by the manhole in vertical direction can be expressed as

Fnet = FB + FEPP − FM − FSH − dry − FSH − wet

(13)

Following the above analysis, time histories of the five component forces (the seepage force (FEPP ), the buoyancy (FB ), manhole’s selfweight (FM ), Sidewall friction above groundwater table (FSH − dry ), sidewall friction below groundwater table (FSH − wet )) acting on the conventional manhole can be obtained accordingly, as shown in Fig. 9. The self-weight of the manhole (FM ) was taken to be constant throughout

Fig. 10. Typical uplift time history curve during earthquakes (shaking started at t = 5 s). 56

Soil Dynamics and Earthquake Engineering 119 (2019) 51–61

Z. Zhang, S.C. Chian

Fig. 12. Manholes fitted with sandpaper.

which cannot be measured directly since no standard instrumentation can offer this capability. Hence, pore pressure transducer, P1, located at the side of the manhole exterior wall 1 m below the groundwater table, as shown in Fig. 5, is adopted for the determination of assumed average degree of liquefaction across the depth of the manhole. Based on the matching resultant force from the experiment and the above theory, the computation of the sidewall frictional resistance within the saturated layer in the experiment is assured.

Fig. 11. Manhole uplift curve with varying α values (shaking started at t = 5 s).

s = s1 + s2

(14)

s1 = 0.32ln(t)−0.56

(15)

s2 =

7 cos(2π(t − 0.6)) 1000

(16)

By considering the uplifting manhole as a free body in motion, the acceleration of the manhole can be acquired from double differentiation of the uplift displacement time history. Thereafter, the resultant force can be acquired by Eq. 17 according to Newton’s law.

4. Experimental validation of sidewall friction In order to validate the above computation of component forces and the influence of the sidewall friction, a series of centrifuge tests with roughened-sidewall manholes was carried out. These manholes were wrapped with sandpapers of different coarseness around the sidewall as shown in Fig. 12. P220 grit sandpaper has an average particle diameter of 68 µm, while P60 grit has a much coarser average particle diameter of 269 µm. P220 grit and P60 grit sandpaper were applied to MH-Phi30-1 and MH-Phi-32-1 respectively. The friction angle of the soilsandpaper interface was checked using a modified direct shear test. In this test setup, the lower ring was substituted with an aluminium plate overlaid with sandpaper. In this way, the friction angle of the interface of the soil and sandpaper can be measured following the procedure of a conventional direct shear test. Table 2 summarises the properties of the manholes in these test cases, and these parameters are presented in Fig. 13 for a clearer comparison.

(17)

F = ma

in which, m and a are the manhole’s mass and acceleration respectively. To investigate the difference in the resultant forces acting on a manhole with different uplift magnitude, α is introduced into Eq. 14, yielding Eq. 18. Values of α are assigned to be 0.2 and 2 to obtain two different uplift curves (the maximum displacement is 0.1 m and 1.15 m when α = 0.2andα = 2 respectively), as shown in Fig. 11.

s′ = α(0.32ln(t) −0.56) +

7 cos(2π(t − 0.6)); 1000

s1′ = α(0.32ln(t) −0.56);

t≥5

t≥5

(18) (19)

Since the manhole’s maximum uplift displacement is controlled by the backbone curve s1′, focus is placed on this parameter. By double differentiating Eq. 19, the manhole’s acceleration is expressed as Eq. 20. The ratio of the resultant force acting on the manhole and its selfweight (Eq. 21) is calculated to be merely 0.13% when t = 5 s (α = 1). Even at much larger maximum uplift displacement of 1.1 m (α = 2 ), the ratio is still ignorable (0.27%). This implies that the resultant force acting on the manhole during uplift is rather low relative to the manhole self-weight. When the soil liquefies, the resultant force hovers about 0 (i.e., the factor of safety is close to 1) and the vibration term(s2) breaches positive resultant force values momentarily. This finding is consistent with the net force obtained from these experiments conducted as shown in Fig. 9.

a=

d2s1′ 0.32α =− 2 ; dt 2 t

Fnet ma 0.32α = =− 2 ; FM mg t ∙g

t ≥5

t ≥5

(20)

(21)

Looking back at the component forces, the manhole’s self-weight is measured directly and thus reliable. The buoyancy force is also accurate since it is calculated directly using the classical Archimedes’ principle. As for the computation of seepage force, it is computed using two measurements, i.e., the excess pore pressure at the manhole invert and its base area. The remaining component force is the sidewall friction

Fig. 13. Interface friction angles and groundwater tables in tests. 57

Soil Dynamics and Earthquake Engineering 119 (2019) 51–61

Z. Zhang, S.C. Chian

Fig. 14. Centrifuge test results of conventional and treated cases.

4.1. Tests Results of Model MH-Phi-20-1, MH-Phi-30-1, MH-Phi-32-1

conventional manhole (MH-Phi-20-1) suffered an uplift of 0.52 m. However, the sidewall-roughened manholes (MH-Phi-30-1 and MH-Phi32-1) suffered a lower uplift displacement, thereby indicating the effectiveness of roughened manhole wall against uplift.

Fig. 14 shows the model instrumentation layout and test results of model MH-Phi-20-1, MH-Phi-30-1 and MH-Phi-32-1. In this figure, the 58

Soil Dynamics and Earthquake Engineering 119 (2019) 51–61

Z. Zhang, S.C. Chian

Fig. 15. Comparison of component forces.

residual frictional resistance below the groundwater table, both of which contributed to the manhole stability. Due to the reduced uplift in the case of the roughened-sidewall manholes, greater seepage force on the manhole invert was produced due to accumulation of pore water pressure under a stationary manhole. This seepage force is one of the key factors leading to the manhole’s uplift [6,11,33]. However, with the aid of the component force computation, it is observed that this increase in seepage force is countered by a proportional increase in sidewall friction. This is confirmed by the attainment of a near zero resultant force for all three manhole cases, regardless of their differences in soil-sidewall interface frictional angle. This presents an interesting finding that the equilibrium of component forces and uplift of the manhole are co-dependent on each other and should not be analysed separately on their own. On the other hand, this study presents another intriguing finding that despite the obvious reduction in uplift displacement due to the adoption of remediation technique with roughened sidewall, the near zero resultant force do not provide easy

The manhole with coarser sidewall (MH-Phi-30-1 and MH-Phi-32-1) possesses a higher soil-sidewall friction angle and therefore is expected to offer a higher frictional resistance along the manhole sidewall. This discouraged uplift of manhole which in turn inhibit relieve of the excess pore pressure. Therefore, excess pore pressure P2 located at the manhole invert was kept higher as compared to the manhole without sandpaper (MH-Phi-20-1). 4.2. Component forces Similar to the earlier centrifuge tests, component forces are computed for the three manhole cases (MH-Phi-20-1, MH-Phi-30-1 and MHPhi-32-1). Fig. 15 shows the comparison of the seepage force (FEPP ), buoyancy (FB ), self-weight (FM ), side frictions (FSH − dry and FSH − wet ), the net resultant force (Fnet ), and the uplift displacement. From Fig. 15, it can be found that the roughened sidewall possesses a larger frictional resistance in the dry sand region. It also improved the 59

Soil Dynamics and Earthquake Engineering 119 (2019) 51–61

Z. Zhang, S.C. Chian

manhole uplift failure. It was also concluded that the onset and cessation of the manhole uplift can be determined effectively using the resultant of the component forces, where the uplift commences as soon as the resultant force turned positive (i.e., upward) and stops when the resultant force returned to negative soon after the termination of shaking. By considering the sidewall friction in the resultant force computation, it was observed that the forces equilibrate at zero during earthquake loading, which confirms the robustness of the computation methodology of the component forces as well as the need to incorporate sidewall friction for more accurate estimate of the time of uplift. By carrying out a theoretical evaluation of the resultant force based on the uplift time history, the resultant force acting on the underground structure during earthquakes made up only about 0.27% of the manhole’s self-weight. This is in agreement with observations in the centrifuge experimental results. This finding highlights the extreme challenges in predicting the liquefaction induced underground uplift from the perspective of forces analysis. In order to further justify the importance of the frictional resistance between the soil and sidewall of the manhole, a novel remediation measure was proposed by roughening the manhole’s sidewall. Results showed that the uplift displacement was reduced significantly with the aid of the roughened sidewall of the manhole. Acknowledgements The authors would like to thank the financial support for the first author's Ph.D. programme and the Tier 1 Grant titled Novel Remediation against Damage to Lifelines in Liquefied Soil by the Ministry of Education in Singapore. References [1] Wang C-Y, Dreger DS, Wang C-H, Mayeri D, Berryman JG. Field relations among coseismic ground motion, water level change and liquefaction for the 1999 Chi-Chi (Mw= 7.5) earthquake, Taiwan. Geophys Res Lett 2003;30. [HLS1. 1-HLS1. 4]. [2] Tobita T, Kang G-C, Iai S. Centrifuge modeling on manhole uplift in a liquefied trench. Soils Found 2011;51:1091–102. [3] Carter L, Green R, Bradley B, Cubrinovski M. The influence of near-fault motions on liquefaction triggering during the Canterbury earthquake sequence. Soil Liq Recent Large-Scale Earthq 2014;57. [4] Hyodo M, Orense R, Noda S. Slope failures in residential land on volcanic fills during the 2011 Great East Japan earthquake. Soil Liq Recent Large-Scale Earthq 2014:95. [5] Scawthorn C, Miyajima M, Ono Y, Kiyono J, Hamada M. Lifeline aspects of the 2004 Niigata ken Chuetsu, Japan, earthquake. Earthq Spectra 2006;22:89–110. [6] Yasuda S, Kiku H. Uplift of sewage manholes and pipes during the 2004 NiigatakenChuetsu earthquake. Soils Found 2006;46:885–94. [7] Aydan Ö, Ulusay R, Hamada M, Beetham D. Geotechnical aspects of the 2010 Darfield and 2011 Christchurch earthquakes, New Zealand, and geotechnical damage to structures and lifelines. Bull Eng Geol Environ 2012;71:637–62. [8] Bhattacharya S, Hyodo M, Goda K, Tazoh T, Taylor C. Liquefaction of soil in the Tokyo Bay area from the 2011 Tohoku (Japan) earthquake. Soil Dyn Earthq Eng 2011;31:1618–28. [9] Wei Y-C, Chuang W-Y, Hung W-Y, Chen H-T, Lee C-J. Centrifuge Modeling on Uplift Behavior Of Immersed Tunnel Embedded In Liquefied Soils, In: Proceedings of the 8th International Conference on Urban Earthquake Engineering 6; 2011. [10] Chian SC, Madabhushi SPG. Excess pore pressures around underground structures following earthquake induced liquefaction. Int J Geotech Earthq Eng (IJGEE) 2012;3:25–41. [11] Koseki J, Matsuo O, Koga Y. Uplift behavior of underground structures caused by liquefaction of surrounding soil during earthquake. Soils Found 1997;37:97–108. [12] Nishio N. Mechanism of projection of sewerage manholes above ground due to soil liquefaction. Doboku Gakkai Ronbunshu 1994;1994:51–4. [13] Sasaki T, Tamura K. Prediction of liquefaction-induced uplift displacement of underground structures, In: Proceedings of the 36th Joint Meeting US-Japan Panel on Wind and Seismic Effects, 2004, p. 191–98. [14] Ling HI, Mohri Y, Kawabata T, Liu H, Burke C, Sun L. Centrifugal modeling of seismic behavior of large-diameter pipe in liquefiable soil. J Geotech Geoenviron Eng 2003;129:1092–101. [15] Hwang J-I, Kim C-Y, Chung C-K, Kim M-M. Viscous fluid characteristics of liquefied soils and behavior of piles subjected to flow of liquefied soils. Soil Dyn Earthq Eng 2006;26:313–23. [16] Chian SC, Madabhushi SPG. Effect of soil conditions on uplift of underground structures in liquefied soil. J Earthq Tsunami 2012;6:1250020. [17] Otsubo M, Towhata I, Hayashida T, Shimura M, Uchimura T, Liu B, Taeseri D, Cauvin B, Rattez H. Shaking table tests on mitigation of liquefaction vulnerability

Fig. 15. (continued)

estimate of the uplift displacement. This supports the earlier theoretical discussion that the resultant force is very sensitive and a minute change in resultant force would lead to a significant change in uplift displacement. 5. Conclusion For underground structures in liquefied soil, the frictional resistance on the soil-structure interface in past studies was often ignored due to its low magnitude after the excess pore pressure was generated. In this paper, the effect of the frictional resistance on the uplift of manholes was investigated through a series of dynamic centrifuge tests. Developing from the computation of the component forces acting on the manhole during earthquake loading, the drastic increase in seepage force and significant drop in shear resistance at the soil-manhole interface were observed to be the two main reasons leading to the 60

Soil Dynamics and Earthquake Engineering 119 (2019) 51–61

Z. Zhang, S.C. Chian

[26] Kang G-C, Tobita T, Iai S, Ge L. Centrifuge modeling and mitigation of manhole uplift due to liquefaction. J Geotech Geoenviron Eng 2012;139:458–69. [27] Chian SC, Tokimatsu K, Madabhushi SPG. Soil liquefaction–induced uplift of underground structures: physical and numerical modeling. J Geotech Geoenviron Eng 2014;140:04014057. [28] Jaky J. The coefficient of earth pressure at rest. J Soc Hung Archit Eng 1944;78:355–8. [29] Tobita T, Kang G-C, Iai S. Estimation of liquefaction-induced manhole uplift displacements and trench-backfill settlements. J Geotech Geoenviron Eng 2012;138:491–9. [30] Tobita T, Iai S, Kang GC, Hazazono S, Konishi Y. Estimation of the uplift displacement of a sewage manhole in liquefied ground. 2008 SEISMIC ENGINEERING CONFERENCE: Commemorating the 1908 Messina and Reggio Calabria Earthquake. AIP Publishing; 2008. p. 456–63. [31] Nishio N. Mechanism of projection of sewerage manholes above ground due to soil liquefaction, In: Proceedings of the Japan Society of Civil Engineers, Dotoku Gakkai, 1994, pp. 51–51. [32] Koseki J, Matsuo O, Koga Y. Uplift behavior of underground structures caused by liquefaction of surrounding soil during earthquake, In: Proceedings of the Soil and Foundation of Japanese Geotechnical Society, 37, p. 97-108; 1997. [33] Koseki J, Matsuo O, Ninomiya Y, Yoshida T. Uplift of sewer manholes during the 1993 Kushiro-Oki earthquake. Soils Found 1997;37:109–21.

for existing embedded lifelines. Soils Found 2016;56:348–64. [18] Watanabe K, Sawada R, Koseki J. Uplift mechanism of open-cut tunnel in liquefied ground and simplified method to evaluate the stability against uplifting. Soils Found 2016;56:412–26. [19] Ramirez J, Barrero AR, Chen L, Dashti S, Ghofrani A, Taiebat M, Arduino P. Site response in a layered liquefiable deposit: evaluation of different numerical tools and methodologies with centrifuge experimental results. J Geotech Geoenviron Eng 2018;144:04018073. [20] Kang G-C, Tobita T, Iai S. Seismic simulation of liquefaction-induced uplift behavior of a hollow cylinder structure buried in shallow ground. Soil Dyn Earthq Eng 2014;64:85–94. [21] Kawabata T, Uchida K, Ooishi J, Nakase H. Uplift Mechanism of Underground Structures in Dry Sand Subjected to Cyclic Simple Shear by DEM, In: Proceedings of the Fifteenth International Offshore and Polar Engineering Conference, International Society of Offshore and Polar Engineers; 2005. [22] Schofield AN. Dynamic and earthquake geotechnical centrifuge modelling; 1981. [23] Tsuchida H. Prediction and countermeasure against the liquefaction in sand deposits. Abstract of the Seminar in the Port and Harbor Research Institute. 1970. p. 31–333. [24] Kestin J, Sokolov M, Wakeham WA. Viscosity of liquid water in the range− 8 °C to 150 °C. J Phys Chem Ref Data 1978;7:941–8. [25] Stringer M, Madabhushi S. Novel computer-controlled saturation of dynamic centrifuge models using high viscosity fluids. Geotech Test J 2009;32:559–64.

61