Failure mechanism of unsaturated landslide dam under seepage loading – Model tests and corresponding numerical simulations

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Failure mechanism of unsaturated landslide dam under seepage loading – Model tests and corresponding numerical simulations Xi Xiong a, Zhen Ming Shi b,c, Sheng Gong Guan b, Feng Zhang a,⇑ a

Department of Civil Engineering, Nagoya Institute of Technology, Nagoya, Japan b Department of Geotechnical Engineering, Tongji University, Shanghai, China c Key Laboratory of Geotechnical and Underground Engineering, Ministry of Education, Tongji University, Shanghai, China Received 7 November 2017; received in revised form 27 April 2018; accepted 30 May 2018 Available online 14 September 2018

Abstract Due to the rapid cracking, sliding and packing of its geomaterials, a landslide dam (LD) is usually weaker in structure than the undisturbed ground and is more vulnerable to seepage loading. In addition, since the naturally packed geomaterials of an LD are mostly in an unsaturated state, it is necessary to use a suitable constitutive model that can describe the mechanical behaviour of the soils under an unsaturated/saturated state in a unified way and whose material parameters can be determined in a rational way in order to accurately simulate the failure mechanism of the LD. In this paper, water retention tests and flume tests were conducted on model LDs prepared with three different ground materials. An unsaturated soil constitutive model was selected for the corresponding numerical simulations. Based on the results of water retention tests and triaxial tests, the parameters of the LD materials were properly determined. Using these parameters, soil-water-air coupling finite element analyses were conducted to simulate the flume tests on the model LDs. By comparing the calculated results with the test results, it was found that the numerical method used in this paper has satisfactory accuracy for describing the different failure mechanisms of the model LDs under seepage loading. The results indicate that the material properties of the LDs, especially the strength and the difference in void ratio between the unsaturated and the saturated states, play important roles. In addition, the influence of the rate of the rise in water head was also investigated by numerical tests. The purpose of this research is to provide a scientific basis for decision-making in the disaster mitigation process of LDs with a comprehensive method for understanding the failure mechanism of these LDs. Ó 2018 Production and hosting by Elsevier B.V. on behalf of The Japanese Geotechnical Society.

Keywords: Landslide dam; Failure mechanism; Unsaturated/saturated soil; Model flume test; Soil-water-air coupling FEM; Constitutive model

1. Introduction A landslide dam (LD) is a kind of natural dam that forms worldwide when the body of a landslide partially or completely blocks a river channel, mostly triggered by rainfall and/or an earthquake (Costa and Schuster, 1988; Ermini and Casagli, 2003). Once an LD has formed, water Peer review under responsibility of The Japanese Geotechnical Society. ⇑ Corresponding author. E-mail addresses: [email protected] (X. Xiong), [email protected] (Z.M. Shi), [email protected] (F. Zhang).

impoundment will result in landslide-dammed lakes (LD lakes). Due to the rapid packing of natural geomaterials, an LD is normally weak in structure, which can lead to its failure with the increase in the water level of the LD lake a short time after its formation. The flooding caused by an LD failure threatens the people and property in the downstream area. For example, an LD failure in the Yigong area of Tibet, China in 2000 caused about 100 people to lose their lives and 5000 people to lose their homes (Zhu et al., 2003). Although most LDs fail naturally or are removed artificially soon after their formation, some large LDs remain for a long period of time and are developed for

https://doi.org/10.1016/j.sandf.2018.05.012 0038-0806/Ó 2018 Production and hosting by Elsevier B.V. on behalf of The Japanese Geotechnical Society.

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hydroelectric generation or tourism (Zhao, 2008; Duman, 2009; Wang et al., 2013). Therefore, apart from disaster mitigation, and considering their potential usage, it is absolutely necessary to investigate the stability of LDs carefully. The failure modes of an LD generally include overtopping, piping and slope failure (Costa and Schuster, 1988). Since natural geomaterials are mostly in an unsaturated state, the increase in the water level of the LD lake leads to seepage loading and a change in the degree of saturation in the LD, causing piping or slope failure. If the inflow equals the outflow, the water level of the LD lake could basically be kept stable and no LD failure would occur for hundreds of years, e.g., Usoi LD (Schuster and Alford, 2004). However, once the water level of the LD lake reaches the dam crest, overtopping will be triggered and the entire breach process will continue and cannot be stopped (Yang et al., 2015). Therefore, it is more complex to estimate the stability of an LD under seepage loading. A few research works on the stability of LDs under seepage loading have focused on examining the influential factors. Case studies have shown that the following three factors are most influential to the longevity of an LD: (1) volume and rate of inflow to the impoundment, (2) size and shape of the dam and (3) geotechnical characteristics of the dam (Costa and Schuster, 1988). Both field and laboratory experiments (Awal et al., 2011; Chen et al., 2015; Gregoretti et al., 2010; Kazama et al., 2012; Okeke and Wang, 2016; Wang et al., 2013), as well as numerical simulations (Awal et al., 2009; Shi et al., 2015a; Sun et al., 2016), have been conducted to study how these factors work in the stability of LDs. With the water level rising in the LD lake, the increasing portion of the slope being filled with water contributes to the instability of the LD (Awal et al., 2009; Sun et al., 2016). For the size and shape of an LD, the failure modes of the LD could change from piping to slope failure with an increase in the downstream-face slope angle (Gregoretti et al., 2010; Shi et al., 2015a). LD materials usually have a wide grain size distribution, namely, from millimetres to meters (Ermini and Casagli, 2003). LDs composed of a mixture of materials are more likely to fail due to piping, and an increase in the fines content may raise the potential for piping. Dams composed of homogeneous materials, however, are more likely to fail due to downstream slope saturation (Kazama et al., 2012; Okeke and Wang, 2016). LDs consisting of high-density ground materials are very capable of retarding piping (Okeke and Wang, 2016; Wang et al., 2013). Moreover, according to large-scale outdoor experiments, dam-crest settlement, seepage-water turbidity and self-potential changes can be regarded as premonitory factors of LD failure under seepage loading (Wang et al., 2018). The above-mentioned studies, however, discussed little about the relationship between the influencing factors and the LD failure mechanism under seepage loading. Moreover, in the stability analyses of these studies, the material properties, such as the stress-strain relation of

unsaturated soil, were not described by a rational constitutive model. Therefore, a rational relationship between the material properties and the LD failure mechanism has not yet been established. As a result, current methods for analysing the stability of LDs and failure modes (Takahashi and Kuang, 1988; Meyer et al., 1994.; Ermini and Casagli, 2003; Awal et al., 2008; Dong et al., 2009; Dong et al., 2011; Weidinger, 2011) have given little consideration to the element behaviour of the LD materials. In this paper, water retention tests and flume tests were conducted on model LDs with three typical LD materials prepared by mixing silica sand with different particle sizes. Based on the results of the water retention tests conducted in this paper and the permeameter tests conducted by Shi et al. (2015b) on these LD materials, the parameters of the LD materials for the water retention curves (WRCs) were determined. Meanwhile, large-size triaxial compression tests under a saturated condition, available in literature, were used to determine the material parameters of the LD materials. Using these parameters, numerical analyses were conducted based on a rational constitutive model for a saturated/unsaturated model (Zhang and Ikariya, 2011) to simulate the flume tests on the LDs as a boundary value problem. By comparing the results of the model tests and the calculations, it was verified whether or not the numerical method used in this paper could describe the LD failure mechanism. Moreover, the influence of the rate of the rise in water head was investigated through numerical tests. 2. Water retention tests 2.1. Selection of model LD materials Since LD materials have a wide range of grain size distributions, mentioned in the works by Casagli et al. (2003), Ermini and Casagli (2003) and Braun et al. (2017), the proper selection of the LD materials for the flume tests and numerical analyses is a very important issue. In the work by Casagli et al. (2003), a detailed discussion of the grain size distribution of soils from many landslide dams in the Northern Apennines was conducted. Three typical grading curves were selected from them for this research, as shown in Fig. 1. In order to assure the operability and repeatability of the model tests and their corresponding numerical calculations, three different manmade LD materials, prepared by mixing silica sand with different particle sizes and the same grading curves as those shown in Fig. 1, were utilized as the LD materials in the model tests and the numerical calculations to investigate the failure mechanism of LDs under seepage loading. These three grading curves represent (1) sand, (2) a well-graded mixture and (3) a gapgraded mixture. Hereafter, for simplicity, the manmade LD materials are just called the LD materials. Permeameter tests on these LD materials under a saturated condition were conducted by Shi et al. (2015b), in which it was found that the permeability of the sand, the well-graded mixture

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the discharge of the samples were measured to calculate the water content under different levels of controlled suction. A large-scale unsaturated permeameter, whose sample size is / = 16 cm in diameter and h = 8 cm in height, as shown in Fig. 3, was designed and manufactured at Tongji University based on the concept proposed by Uno et al. (1990), in which the suction is controlled by the axistranslation method. Two ceramic disks, installed at upper and lower loading pedestals with an air entrance value (AEV) of 100 kPa, were used. They were saturated with vacuumed water and desiccated for one week before the tests. The LD materials were prepared by mixing silica sand with different particle sizes according to the grading curves shown in Fig. 1, and then different samples were prepared by the static compaction method shown in Fig. 4. Since the void ratio is an important state variable in the constitutive model for unsaturated soils (Zhang and Ikariya, 2011), the initial void ratio of the samples was arranged to be the same for all the samples, as e0 = 0.48, in order to eliminate its influence on the WRCs. The other physical properties are listed in Table 1. Six tests were conducted under the same vertical load of rv ¼ 9:8 kPa applied by a counterweight. The WRCs of these LD materials were measured in Tests 1–1, 2–1 and 3–1, respectively. After determining the ranges in suction for the main curves, scanning curves in these ranges were also measured in Tests 1–2, 2–2 and 3–2. In the tests, the samples were firstly compacted and immersed in water under a zero-suction condition (s0 = 0 kPa) to reach an initial degree of saturation. Then, the samples were loaded or unloaded with suction, step by step, and the quantity of water absorption or discharge from the samples was recorded after it reached an approximately constant condition. During the tests, the air pressure was always controlled to be less than the AEV of the ceramic disk. The vertical displacement of the samples was measured, by which the degree of saturation can be calculated accurately. Fig. 5 shows the suction-void ratio relations of the water retention tests. At the initial state of s0 = 0 kPa, volumetric

Sand Well-graded Gap-graded

60

40

20

0 100

10

1

0.1

0.01

0.001

Grain [mm]

Fig. 1. Grain size distribution curves of landslide dam materials.

and the gap-graded mixture at a saturated state were kw = 8.72
104, 3.33
104 and 1.29
104 m/s, respectively, with an average void ratio of e0 = 0.48. Meanwhile, the results of numerous large-scale triaxial compression tests on dam materials are available in literature. Three of those works, conducted by Wei et al. (2005), Araei et al. (2010) and Zhang et al. (2009), were selected for discussion. The reason why these three works were chosen is quite simply because the grain size distribution curves of the soils considered in their works are very similar to those of the three LD materials in this study, as shown in Fig. 2. Utilizing this advantage, it is possible to determine the material parameters of these LD materials on the assumption that non-cohesive soils with the same grain size distribution curve and composed of similar particles (e.g., silica sand) have the same mechanical behaviour. 2.2. Water retention tests Water retentivity is one of the most important properties of unsaturated soils; it is generally expressed by the so-called water retentivity curve (WRC). In order to investigate the mechanical properties of unsaturated LD materials and to rationally explain the failure mechanism of LDs from these properties, water retention tests were conducted in this paper. In the water retention tests, the volume and 120 Sand LD material Material in Wei et al. (2005)

100

Percentage finer by weight [%]

Percentage finer by weight [%]

120

80 60 40 20 0

120 Well-graded LD material Material in Araei st al. (2010)

100 80 60 40 20

Percentage finer by weight [%]

Percentage finer by weight [%]

100

80

0 10

1

0.1

Grain [mm]

0.01

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Gap-graded LD material Material in Zhang et al. (2009)

100 80 60 40 20 0

100

10

1

0.1

Grain [mm]

0.01

0.001

100

10

1

0.1

Grain [mm]

Fig. 2. Similarity of grading curves between manmade LD materials and dam materials.

0.01

0.001

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Fig. 3. Schematic diagram of large-scale unsaturated permeameter for landslide dam materials.

(a) Sannd

(b) Well-graded mixture

(c) Gap-graded mixture

Fig. 4. Samples used in water retention tests.

Table 1 Physical properties of LD materials in tests. LD materials

Specific gravity

Initial void ratio e0

Maximum void ratio emax

Minimum void ratio emin

Relative density Dr (%)

Sand Well-graded Gap-graded

2.65 2.65 2.65

0.480 0.480 0.480

0.903 0.862 1.015

0.474 0.469 0.441

98.6 97.2 93.2

shrinkage occurred in all the samples along with a rise in the degree of saturation, which was most significant in the gap-graded LD material. During the drying process, along with the rise in suction, the effective stress of the samples increased and consequently led to a decrease in void ratio. Fig. 6 presents the tested WRCs obtained from the above-mentioned water retention tests. By comparing these three LD materials, it is known that the degree of saturation of the sand LD material is the lowest both at the residual state and the zero-suction state, implying that the water retentivity is poor, while the water retentivity of the gapgraded LD material is the best. It should be pointed out

here that the initial degree of saturation of the sand, 67%, is really lower than that given in most previously published results. In the test, each equilibration time during the first water immersion in the WRC tests was set to be at least 12 h, judging by the condition that the variation in water content stopped. The test was repeated two times and almost the same results were obtained. The only reason for this abnormality is that the sand used in this research is not thought to be pure clean sand and it was compacted to a very highly dense state (e = 0.48, relative density Dr (%) = 98.6). According to the grading curves given in Fig. 1, it is seen that within these LD materials, the sand has the lowest fines

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Fig. 5. Suction-void ratio relationship in water retention tests.

Fig. 6. Tested water retention curves.

content (particle size under 0.075 mm). Therefore, the water retentivity of the LD material is mainly determined by the fines content, and an increase in the fines content will lead to an increase in water retentivity. In recent research (Lu et al., 2014), it was pointed out that positive pore water pressure will also have an influence on the WRCs and the soaked degrees of saturation. In the present case, however, no positive pore water pressure was applied to the samples. Therefore, the influence cannot be confirmed in the present case, but it is worthy of investigation as a future task. The drying curve of the well-graded LD material is close to the wetting curve, meaning that the moisture hysteresis is also poor, while the differences in the drying and wetting curves are larger in the gap-graded LD material with significant moisture hysteresis. It is natural, therefore, to suggest

that the reason for this difference in moisture hysteresis is due to the uniformities of the LD materials. The voids in the soil consist of a relatively wide range of radii bounded by narrow channels. The drying process depends mainly on the narrow channels connecting the small-radius pores, while the wetting process depends more significantly on the large pores with maximum radii; this is called the ink-bottle effect (Maqsoud et al., 2012). Under the same void ratio, the well-graded LD material has the most uniform pore sizes, which results in the small difference between the narrow radius and the maximum radius. For this material, the drying process and the wetting process are similar and the moisture hysteresis is not obvious. Therefore, the moisture hysteresis of the LD materials is related to the distribution of the void radii.

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3. Flume tests on model landslide dams As most LDs form in mountainous areas after rainfall or earthquakes, it is difficult to conduct field tests on them. Model tests, however, are controllable, easy to observe and monitor, and have fine operability and repeatability. Therefore, it is useful to study the failure mechanism of LDs through model tests. In this paper, flume tests on model LDs using the above-mentioned three LD materials were conducted to study the failure mechanism of unsaturated LDs under seepage loading. 3.1. Experiment setup The rectangular flume utilized in this paper is shown in Fig. 7; it is 50 m in length, 0.8 m in width and 1.25 m in height, and had a slope angle of 0°. The walls of the flume

were made of transparent glass that allowed easy visibility of the test phenomenon. The bottom of the flume had an impermeable concrete surface. As shown in Fig. 8, each model LD was made of a homogeneous ground; it was 390 cm in bottom width, 50 cm in crest width and 90 cm in height, and had a slope angle of 30° in both upstream and downstream directions. All the LD materials were prepared using silica sand and had a specific gravity of 2.65. When creating the model LDs, the layered compaction method was used to control the void ratio to be e0 = 0.48, exactly the same as that in the permeameter tests discussed in the previous section. The values of the physical properties are listed in Table 1. The boundary conditions of the LDs are given as: (a) the bottom and the side surfaces are in contact with the impermeable boundaries of the flume, and the other surfaces are free and permeable; (b) the initial water level on the left

Fig. 7. Flume used in model tests of landslide dam.

Camera 3 Camera Side view Upstream

Downstream 90 cm

50 cm

30°

30°

Front view

390 cm

Camera 2 80 cm

Top view

Camera 1

Fig. 8. Layout of landslide dam model and setup of cameras.

X. Xiong et al. / Soils and Foundations 58 (2018) 1133–1152

Fig. 9. Changes in water level of different model landslide dams in flume tests.

surface of the LD lake is 0 cm; and (c) the water level increases under a constant inflow of q = 1130 ml/sec. The measurement system consists of three cameras, as shown in Fig. 8. The development of the seepage line and the incremental rate of the water level of the LD lake can be recorded by Camera 1, and the failure process of the model LDs can be recorded by all the cameras. 3.2. Test methodology

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and 10. With the highest permeability coefficient, the incremental rate of the water level of the sand LD was lower than that of the others. The reason is that there was an obvious outflow from the downstream surface of the sand LD after the seepage line reached the downstream surface. The sand LD, well-graded LD and gap-graded LD failed at water levels of 83, 90 and 57 cm, respectively. The failure modes occurring in the LDs with the different materials were also different. For the sand LD, the crest width was shortened by the sliding of the downstream slope, which led to the dam failure, as shown in Fig. 11. The seepage line of the sand LD reached the downstream slope quickly, and then it increased, which resulted in an increase in the hydraulic gradient. Under seepage loading, the toe of the slope on the upstream side failed first, due to the increase in pore water pressure and the decrease in effective stress. Then, the toe of the slope on the downstream side slid. Finally, the sliding area on the downstream side gradually expanded to the entire downstream slope, a typical sliding failure. For the well-graded LD, it was stable under seepage loading, but finally failed due to overtopping, as shown in Fig. 12. Compared with the sand LD, the seepage line development of the well-graded LD was slower. Even though the hydraulic gradient increased significantly, the

Three flume tests on the model LDs were conducted with sand (Case1), a well-graded mixture (Case2) and a gapgraded mixture (Case3) as the LD materials. Firstly, the LD materials were prepared by mixing dry silica sand with different grain sizes according to Fig. 1. Then, the model LDs were prepared by compacting every layer to have a thickness of 15 cm. After that, the valve of the water pipe was opened and the upstream water level began to rise with the prescribed inflow. At the same time, the cameras started to take pictures at regular intervals. The water level at the upstream increased continually until the LD failed. Three failure modes were considered for the LDs in the tests: (1) the sliding failure of the downstream slope, (2) overtopping failure and (3) piping failure in which the upstream water level stopped increasing for the enlarged piping channel. As to the judgement of the piping failure, a more detailed explanation should be given here. A constant inflow is maintained in the flume tests. According to the permeability of the LD materials at the saturated state (Shi et al., 2015b) and the model test observation, the seepage through the downstream surface of the LD models is much smaller than the inflow. However, the outflow will increase significantly because of the enlargement of the piping channel. Therefore, the state at which the upstream water level stops increasing because of the enlargement of the piping channel is judged as piping failure. 3.3. Test results Changes in the water level and the seepage line were seen from the beginning of the tests to the LD failure in Figs. 9

Fig. 10. Changes in seepage lines of different model landslide dams in flume tests.

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000 min

030 min

060 min

090 min

120 min

150 min

180 min

210 min

240 min

270 min

300 min

330 min

360 min

390 min

420 min

Fig. 11. Failure process and seepage line of model landslide dam in flume test (Case1, sand).

downstream slope of the well-graded LD was generally stable. After the water level reached the crest, the flow eroded the crest and the downstream surface, leading to the failure of the whole LD, a typical overtopping failure. For the gap-graded LD, the development of the seepage line was the slowest among the three LDs, as shown in Fig. 13. With the increase in the water level of the LD lake, many horizontal cracks occurred under the seepage line of the LD, as shown in Fig. 14(a). When water flowed through these cracks, the loss in water head and energy was reduced, which means it took more fine grains away. Therefore, these cracks were more easily broadened by the flow and connected with each other, which eventually led to the formation of the piping channel shown in Fig. 14(b). Finally, the water level stopped rising, meaning that the outflow of the LD flowing through the piping channel was equal to the inflow, which means that the gap-graded LD failed in piping, a typical piping failure. It can be found from the above-mentioned test results that under constant-inflow seepage loading, the failure modes of the LDs differed from those of the ground materials.

4. Numerical simulation of flume tests 4.1. Constitutive model In order to properly describe the mechanical behavior of the LD materials under an unsaturated/saturated condition due to the rising of the water level in the LD lake, a constitutive model (Zhang and Ikariya, 2011) for unsaturated soil was adopted in the numerical calculation. After Alonso et al. (1990) proposed the Barcelona Basic Model (BBM), many constitutive models for unsaturated soils were developed. The main difference in these models is the choice of independent state variables, which can be divided into three types: (1) net stress and suction, (2) Bishop-type skeleton stress and suction and (3) Bishoptype skeleton stress and degree of saturation (Cui and Delage, 1996; Chiu and Ng, 2003; Kohgo et al., 1993; Loret and Khalili, 2002; Sun et al., 2007; Zhang and Ikariya, 2011; Ohno et al., 2013). In the first two types of constitutive models, although the stress-suction-strain relationship of unsaturated soils is well described, the saturation is not directly considered. However, even under the

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000 min

030 min

060 min

090 min

120 min

150 min

180 min

240 min

270 min

300 min

330 min

360 min

Fig. 12. Failure process and seepage line of model landslide dam in flume test (Case2, well-graded).

000 min

030 min

060 min

090 min

120 min

150 min

173 min

180 min

193 min

Fig. 13. Failure process and seepage line of model landslide dam in flume test (Case3, gap-graded).

same net stress and suction path, the stress-strain relationship and strength of unsaturated soil with different degrees of saturation differ from each other. Therefore, the third type of constitutive model (Zhang and Ikariya, 2011) was used in the calculation. Based on the experimental results, Zhang and Ikariya (2011) proposed an unsaturated soil constitutive model,

using the skeleton stress and the degree of saturation as the independent variables. In the model, it is assumed that the normally consolidated line in the unsaturated state (N. C.L.S.) is parallel to the normally consolidated line in the saturated state (N.C.L.), but in a higher position than the N.C.L. (see Fig. A1). The skeleton stress is a kind of Bishop effective stress, defined as follows:

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(a) Cracks

(b) Piping channel from front view

Fig. 14. Close-up view of failure pattern in model landslide dam after flume test (Case3, gap-graded).

Table 2 Parameters of WRC for LD materials. LD materials

Sand

Well-graded

Gap-graded

Degree of saturation at which suction is zero S sr Residual degree of saturation S rr Parameter corresponding to drying AEV (kPa) Sd Parameter corresponding to wetting AEV (kPa) Sw Initial stiffness of scanning curve (kPa) k esp

0.68 0.21 0.90 0.20 4000

0.92 0.61 1.00 0.10 3000

0.88 0.46 5.00 1.00 3000

Parameter of shape function c1 Parameter of shape function c2 Parameter of shape function c3

0.65 1.0 50.

0.45 1.2 10.

0.40 0.40 9.0

r00ij ¼ rtij  ua dij þ S r ðua  uw Þdij ¼ rnij þ S r sdij

ð1Þ

where r00ij is the skeleton stress tensor, rtij is the total stress tensor, rnij is the net stress tensor, S r is the degree of saturation, ua is the air pressure, uw is the water pressure and s is the suction. The constitutive model includes nine material parameters, of which five parameters, M, N, k, j and m, have the same meaning as those of the Cam-Clay model. Parameters a, b and b control the degree of over consolidation and can be determined by triaxial tests. Nr is the void ratio at the N.C.L.S. under reference mean skeleton stress pr when the degree of saturation is in the residual state. In the unsaturated soil constitutive model, a WRC is also proposed considering the moisture hysteresis of the unsaturated soil. Depending on the state of moisture, the moisture characteristics are given in three different tangential and arc-tangential functions as the primary drying curve, the secondary drying curve and the wetting curve (see Fig. A2), which include eight parameters. Three of the parameters, c1, c2 and c3, are determined by the curve fitting method, while the other five parameters, k s0 ; S sr ; S rr ; Sd and Sw, have a definite physical meaning and can be determined by water retention tests. The assumption adopted in the model has been verified by an experimental method (Kurimoto et al., 2017). And there have been some studies on the application of the proposed constitutive model (Xiong et al., 2014; Zhang and Maeda, 2015; Zhang et al., 2016). Based on the constitutive model, Xiong et al. (2014) developed a soil-water-air coupling FEM program, called SOFT, and simulated model

tests on the slope failure of an unsaturated Shirasu ground. The calculated results indicated that the slope failure of the unsaturated Shirasu ground can be properly described, on the whole, with satisfactory accuracy. Thus, it is reasonable to select the constitutive model in the analysis of the unsaturated LD failure mechanism under seepage loading. 4.2. Determination of parameters In this paper, to assure the accuracy of the calculation in boundary value problems, all the material parameters involved in the constitutive model are determined based on the element tests. The parameters involved in the WRCs for the LD materials were determined from the results of water retention tests; their values are listed in Table 2. Since large-scale unsaturated triaxial compression tests would take too long to perform and would not be feasible at the present time, unsaturated triaxial tests on the LD materials were not conducted in this paper. The material parameters were basically determined by the results of saturated triaxial tests (Wei et al., 2005; Araei et al., 2010; Zhang et al., 2009), considering the available physical quantities obtained from the water retention tests, such as parameters b and Nr, which can be determined from the water retention tests by comparing the relation between the incremental suction and the change in void ratio. It should be emphasized here that the model can describe both saturated and unsaturated soils with a unified set of parameters. Compared with the saturated state, the unsaturated state requires two additional parameters, that is, b

X. Xiong et al. / Soils and Foundations 58 (2018) 1133–1152 Table 3 Material parameters of LD involved in constitutive model. LD materials

Sand

Well-graded Gap-graded

Compression index k 0.105 Swelling index j 0.007 1.50 Critical state parameter Rcs Void ratio N (p0 = 15 kPa on N.C.L.) 0.46 Poisson’s ratio m 0.30

0.095 0.0050 4.29 0.44 0.30

0.140 0.0050 3.45 0.38 0.30

Parameter of overconsolidation a Parameter of suction b Parameter of overconsolidation b Void ratio Nr (p0 = 15 kPa on N.C.L.S.)

80. 0.50 1.0 0.48

1.0 0.50 3.0 0.48

5.0 0.50 3.0 0.48

and Nr. Parameter Nr can be determined uniquely by water retention tests. While an accurate value for b should be verified by unsaturated triaxial tests, water retention tests can also yield proximate estimations of the value. The calibrated material parameters of the LD materials are listed in Table 3. Fig. 15 shows the element simulation of the triaxial compression tests for the saturated soils with the parameter values listed in Tables 2 and 3, from which it is known that the magnitude and the stress-strain relations are basically the

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same. Therefore, the present constitutive model for unsaturated/saturated soils can accurately describe the mechanical properties of the LD materials under a saturated condition. To compare the mechanical properties of the unsaturated and the saturated LD materials, consolidateddrained triaxial tests are calculated with the same parameters as those listed in Tables 2 and 3. The confining pressure was kept constant at 9.8 kPa during the tests, and the initial degrees of saturation and suction are listed in Fig. 16. It is found from this figure that the strength of the unsaturated LD materials is larger than that of the saturated LD materials. Moreover, the strength of sand is much lower than that of the other materials, and the strength of the well-graded LD material is the highest among the three materials. In the calculation, the value of the saturated permeability of the LD soils is the same as that determined in the works by Shi et al. (2015b). It should be pointed out that the LD soils used in this research are exactly the same as those used in the works by Shi et al. (2015b). The unsaturated permeability coefficients of the sand, the well-graded, and the gap-graded LD materials were set by the formula of Van Genuchten (1980).

Fig. 15. Element simulation of triaxial compression tests for saturated soils (Test data from Wei et al., 2005; Araei et al., 2010; Zhang et al., 2009).

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X. Xiong et al. / Soils and Foundations 58 (2018) 1133–1152

Fig. 16. Stress-strain relationship of three LD materials in unsaturated/saturated condition (r3 = 9.8 KPa).

Fig. 17. Finite element mesh used in numerical calculation with 2D FEM.

4.3. Numerical model and initial conditions The flume tests on the model LDs were simulated with a soil-water-air coupling finite element method (FEM) and the program called SOFT (Xiong et al., 2014), using a finite element–finite difference scheme (FE–FD) for soil–water– air three-phase coupling problems. The values of the soil parameters are the same as those listed in Tables 2 and 3. Fig. 17 shows the finite element mesh used in the simulation. The size of the FEM mesh, composed of 1581 nodes and 1500 4-node isoparametric elements, is the same as that of the model tests under plane strain conditions. The drainage and vented boundary conditions, given in Fig. 18(a), are exactly the same as those in the model tests. For the displacement boundary condition, given Fig. 18(b), the bottom surface is fixed in the x- and y-directions and the other surfaces are free in both directions.

In the numerical simulation, the settings of the initial conditions are very important and greatly affect the accuracy of the calculation. Based on the WRCs, the initial degree of saturation, the pore pressure and the suction of the LDs were set to have the values as those listed in Table 4. Since the model LDs were prepared by the layered compaction method, the materials were heavily compacted at the initial condition. An extra mean effective stress of 12 kPa was added to the whole area to consider the compaction effect. Fig. 19 shows the initial vertical effective stress field in the calculations. In the numerical simulations, a prescribed increment in water head was applied to simulate the rise of the water level shown in Fig. 8. With the same calculation steps, various time intervals were adopted in the different cases to investigate the influence of the rate of the rise in water head on the failure of the LDs, as shown in Table 5. In Case1,

X. Xiong et al. / Soils and Foundations 58 (2018) 1133–1152

(a) Displacement boundary

Drained and unvented condition Drained and vented condition Undrained and unvented conditon

(b) Hydraulic boundary

Fig. 18. Boundary conditions of numerical model.

Table 4 Initial conditions of numerical calculations. LD materials

Degree of saturation Sr0

Pore pressure u0 (kPa)

Suction s0 (kPa)

Sand Well-graded Gap-graded

0.21 0.62 0.49

9.0 9.0 9.0

9.0 9.0 9.0

12

14

16

18

20

22

24

Fig. 19. Initial mean effective stress field of landslide dam model (unit: kPa).

Case2 and Case3, the time interval is 1 sec/step, and their rates of the rise in water head are the same as those of the flume tests. 4.4. Simulations of flume tests In order to verify whether or not the numerical method proposed in this paper can properly describe the failure

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mechanism of LDs under seepage loading, the calculation results for Case1, Case2 and Case3 are compared with the experimental results. Fig. 20 shows the distribution of the degrees of saturation at different times, which coincides quite well with the test results shown in Fig. 9. With the increase in water head, the degrees of saturation increased from the toe of the upstream slope, and seepage lines developed in upstream and downstream directions. Therefore, it is proved that the WRCs evaluated from the water retention tests and the estimated unsaturated permeability coefficients are correct. This assures the accuracy of the calculations in boundary value problems. Fig. 21 shows a comparison of the calculated displacement vectors in different test cases for the three different LDs. At the end of the calculation, sliding was seen to have only occurred in Case1 (the sand LD), while the displacement of the other LDs consists of mainly downward settlement at the rest, which is totally consistent with the phenomenon observed in the flume tests. Although the displacement values are different, slope failure is not triggered directly by the settlement under seepage loading. This is because, at this stage, the largest settlement of the LD happened in Case3 (the gap-graded LD), but it did not fail due to sliding at all. According to Fig. 16, because the strength of the sand is much lower than the other materials, the increase in excessive pore water pressure and the decrease in effective stress might cause the slope failure to occur more easily in Case1 than in the other cases. Therefore, displacement vectors can be utilized to identify the slide failure of LDs under seepage loading. Fig. 22 shows a comparison of the large volume strain areas in the three LDs at the end of the calculation, at which time the volume strain (shrinkage) of the gapgraded LD is the largest among the three LDs. According to Table 3, as the difference in pore ratio between N and Nr of the gap-graded LD is the largest among the three cases, an increase in the degree of saturation might cause the significant shrinkage of the soil. Compared with the test results for the gap-graded LD, shown in Fig. 13, in which volumetric shrinkage caused the formation of cracks and finally led to piping failure, the calculation also shows the same phenomenon, implying that piping failure also

Table 5 Calculating cases considering influence of rate of rise in water head. Materials of LD

Total water head h (cm)

Time interval (sec/step)

Relative total water head rising rate (v/v0)*

Case1 Case1-1 Case1-2

Sand

83

1.0 0.50 4.0

1.0 2.0 0.25

Case2 Case2-1 Case2-2

Well-graded

90

1.0 0.50 4.0

1.0 2.0 0.25

Case3 Case3-1 Case3-2

Gap-graded

57

1.0 0.50 4.0

1.0 2.0 0.25

*

v0: real rate of rise in water head in flume tests; v: rate of rise in water head used in numerical simulation.

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X. Xiong et al. / Soils and Foundations 58 (2018) 1133–1152 Case1 h=20 cm

h=57 cm

0.21 Case2 h=30 cm

0.30

h=83 cm

0.40

0.49

0.59

h=57 cm

0.61 Case3 h=20 cm

0.68

h=88 cm

0.74

0.80

0.86

0.57

0.92

h=57 cm

h=40 cm

0.49

0.68

0.65

0.72

0.80

0.88

Fig. 20. Changes in distribution of degrees of saturation in different test cases.

Case1 h=20 cm

h=57 cm

0.00 Case2 h=30 cm

0.25

h=83 cm

0.50

0.75

1.00

h=57 cm

0.00 Case3 h=20 cm

h=88 cm

0.02

0.04

0.06

0.60

0.08 h=57 cm

h=40 cm

0.00

1.25

1.20

1.80

2.40

3.00

Fig. 21. Changes in displacement vectors in different test cases (unit: cm).

Case1 h=83 cm

Case3

Case2 h=88 cm

5.0

5.2

h=57 cm

5.4

5.6

5.8

6.0

Fig. 22. Comparison of volume strain at failure stage in different test cases (unit: %).

occurred in the simulation. Moreover, as shown in Fig. 23, the areas of large volume strain progressed gradually: (1) h = 20 cm, a large volume strain area did not form, (2) h = 40 cm, a horizontal large volume strain area formed under the seepage line and (3) h = 57 cm, a large volume strain area extended from the upstream slope to the downstream slope surface, and the largest volume strain occurred at the toe of the slope to the downstream side,

Case3 h=20 cm

which can be regarded as the developing process of the piping channel. Therefore, under seepage loading, the piping failure of LDs can easily be identified by the volume strain. pffiffiffiffiffiffiffi Fig. 24 shows the changes in shear strain 2I 2 in the different test cases, where I 2 is the second invariant of the deviatoric strain tensor. As the water head rose from the toe of the upstream slope, it was found that shear strain firstly occurred at the toe of the upstream slope in all cases.

h=57 cm

h=40 cm

5.0

5.2

5.4

5.6

5.8

6.0

Fig. 23. Evolution of volume strain in gap-graded landslide dam in Case3 (unit: %).

X. Xiong et al. / Soils and Foundations 58 (2018) 1133–1152 Case1 h=20 cm

h=57 cm

0

4

Case2 h=30 cm

h=83 cm

8

12

h=57 cm

0

Case3 h=20 cm

16 h=88 cm

3

6

9

12 h=57 cm

h=40 cm

0

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3

6

9

pffiffiffiffiffiffiffi Fig. 24. Development of shear strain 2I 2 in different test cases (unit: %).

Case1-1

Case1

Case1-2

Max䠖1.29

Max䠖1.27 0.00

0.25

0.50

0.75

Max䠖1.30 1.00

1.25

Fig. 25. Displacement vectors of sand landslide dam at same total water head (h = 83 cm) in different calculated cases (unit: %, Case1).

However, the development of shear strain was different in the different LDs. For the sand LD (Case1), shear bands formed on both sides of the slopes at the end of the calculation, but only the shear band of the downstream slope reached the crest, implying that the sand LD failed due to the sliding of the downstream slope the end of calculation. For the well-graded LD (Case2), even with a water head of 88 cm, the shear strain was very small, except for in the vicinity of the upstream slope, which did not influence the entire stability of the well-graded LD. For the gap-graded LD (Case3), the shear strain propagated significantly with the development of the seepage line, and the maximum shear strain was located at the toe of the slope on the downstream side, indicating that the gap-graded LD failure was caused by the change in the degree of saturation in the simulation. Therefore, based on the calculated shear strain, displacement vectors and volume strain, the three failure modes of the LDs can be clearly identified. By comparing the results of the flume tests with the corresponding numerical calculations, it can be concluded that, based on a rational constitutive model and the proper determination of the material parameters for the ground materials, the numerical analysis can clearly describe the different failure mechanisms of the LDs observed in the model flume tests with satisfactory accuracy. Based on the numerical calculations, a method for judging the LD failure modes can be proposed as follows: (1) Select a rational constitutive model; (2) Determine the material parameters of the soils from elementary tests; (3) Simulate the boundary value problem of the LDs with proper initial and boundary conditions; (4) Judge the failure mode by the

calculated results of the displacement vectors, volume strain and shear strain. 4.5. Influence of rate of rise in water head After the formation of a landslide dam, the rate of increase in the water level of the LD lake, which is mainly controlled by its catchment area and the local weather, is a very important decision-making factor in risk assessment and disaster mitigation. Therefore, it is absolutely needed in order to analyse the influence of the rate of the rise in water head on the failure mechanism of the LD. The accurate numerical method proposed in this paper is utilized to analyse the failure mechanism of the LD under different water head rising rates, which can be called a kind of numerical test. Three different water head rising rates are considered in the numerical tests; all cases are listed in Table 5. Case1-1, Case2-1 and Case3-1, with a time internal of 0.5 sec/step, represent extreme weather conditions, such as a typhoon or strong rainfall, while Case1-2, Case2-2 and Case3-2, with a time internal 4.0 sec/step, are used to analyse the seepage stability of LDs with long lifetimes. In order to quantitatively analysis the stability of LDs, some elements were selected for which the calculated results have been plotted. For the sand LD, no matter what kind of water head rising rate there might be, downstream slope failure occurred. Figs. 25 and 26 show that the slope of the sand LD slid downstream significantly and a shear band occurred under all the water head rising rates. Moreover, as shown in

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X. Xiong et al. / Soils and Foundations 58 (2018) 1133–1152

Case1-1

Case1

Case1-2

Max䠖15.8

Max䠖17.6 0

Fig. 26. Shear strain

4

Max䠖16.4

8

12

16

pffiffiffiffiffiffiffi 2I 2 in sand landslide dam at same total water head (h = 83 cm) in different calculated cases (unit: %, Case1).

8

Deviator stress q [kPa]

7

B

A

6

A B C.S.L.

Case 1-1

5 4 3 2 1 0 0.0

5.0

10.0

15.0

20.0

15.0

20.0

Mean stress p" [kPa] 8

6

7

Case 1

Deviator stress q [kPa]

Deviator stress q [kPa]

7

8 A B C.S.L.

5 4 3 2 1 0 0.0

6

A B C.S.L.

Case 1-2

5 4 3 2 1

5.0

10.0

15.0

20.0

Mean stress p" [kPa]

0 0.0

5.0

10.0

Mean stress p" [kPa]

Fig. 27. Skeleton stress path at selected elements of sand landslide dam in different calculated cases (Case1).

Fig. 27, the skeleton stress path of point A has reached the C.S.L. and that of point B has also reached an unstable state in all cases. This means that the rate of the rise in water head has little influence on the failure mechanism of the sand LD under seepage loading. It is the low strength of the material that leads to a short lifetime of the sand LD. However, an interesting phenomenon can be seen in Fig. 26 whereby along with the decrease in the rate of the rise in water head, the shear strain at the toe of the downstream slope increases. In addition, it can be found from Fig. 27 that the deviator stress at point B in Case1-2 is larger than that in the other cases. Therefore, a decrease in the water head rising rate may reduce the stability of the sand LD. For the well-graded LD, it was stable under seepage loading in all cases. Fig. 28 shows that, although the shear

strain at the toe of the upstream slope rises with the increase in the water head rising rate, it is small in other parts of the well-graded LD. Moreover, as shown in Fig. 29, compared with the skeleton stress paths of point C in the upstream slope toe, the paths of point D at the toe of the downstream slope have not reached an unstable state for any of the water head rising rates. Therefore, a rise in the water head can cause the settlement of the wellgraded LD and hasten the overtopping, but will not trigger a failure because of the high strength of the well-graded material. Overtopping is the main reason for the erosion of the surface of the downstream side and finally the cause of the failure of the slope. For the gap-graded LD, its failure mechanism was influenced by the water head rising rate. As a result of low permeability, the quick rising of the water head limited the

X. Xiong et al. / Soils and Foundations 58 (2018) 1133–1152

Fig. 28. Shear strain

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pffiffiffiffiffiffiffi 2I 2 in well-graded landslide dam at same total water head (h = 88 cm) in different calculated cases (unit: %, Case2).

25 C D C.S.L.

C

Deviator stress q [kPa]

20

D

Case 2-1 15

10

5

0 0.0

5.0

10.0

15.0

20.0

15.0

20.0

Mean stress p" [kPa] 25

25

C D C.S.L.

C D C.S.L.

20

Case 2

Deviator stress q [kPa]

Deviator stress q [kPa]

20

15

10

Case 2-2 15

10

5

5

0 0.0

5.0

10.0

15.0

0 0.0

20.0

5.0

10.0

Mean stress p" [kPa]

Mean stress p" [kPa]

Fig. 29. Skeleton stress path at selected positions in well-graded landslide dam in different calculated cases (Case2).

Case3-1

Case3

Case3-2

Max䠖7.0

Max䠖6.4 5.0

5.2

5.4

5.6

Max䠖5.7 5.8

6.0

Fig. 30. Volume strain of gap-graded landslide dam at same total water head (h = 57 cm) in different test cases (unit: %, Case3).

development of the seepage line, and it had not reached the downstream slope surface by the end of the calculation in Case3-1. Hence, as shown in Fig. 31, the shear strain of the downstream slope surface in Case3-1 is much smaller than that in the other cases. This means that a piping channel did not form and piping failure did not occur within the gap-graded LD at the water head of h = 57 cm. The same conclusion can be drawn from Fig. 32, in which the skele-

ton stress paths of points F and E have not reached an unstable state in Case3-1. However, the volume strain is the largest in Case3-1, as shown in Fig. 30. It is reasonable, therefore, to say that the LD in Case3-1 will fail due to the piping under seepage loading if the water head continues to rise. In conclusion, both the soil properties and the water head rising rate play important roles in the failure mechanism of the gap-graded LD.

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X. Xiong et al. / Soils and Foundations 58 (2018) 1133–1152

Fig. 31. Shear strain

pffiffiffiffiffiffiffi 2I 2 of gap-graded landslide dam at same total water head (h = 57 cm) in different test cases (unit: %, Case3).

25 E F G C.S.L.

G F E

Deviator stress q [kPa]

20

Case 3-1 15

10

5

0 0.0

5.0

10.0

15.0

20.0

25.0

20.0

25.0

Mean stress p" [kPa]

25

25 E F G C.S.L.

20

Deviator stress q [kPa]

Deviator stress q [kPa]

20

Case 3 15

10

Case 3-2 15

10

5

5

0 0.0

E F G C.S.L.

5.0

10.0

15.0

20.0

25.0

Mean stress p" [kPa]

0 0.0

5.0

10.0

15.0

Mean stress p" [kPa]

Fig. 32. Skeleton stress path at selected positions in gap-graded landslide dam in different calculated cases (Case3).

5. Conclusions In this paper, a series of water retention tests and flume tests for model LDs with three different ground materials were systematically conducted. Meanwhile, based on an unsaturated soil constitutive model (Zhang and Ikariya, 2011), in which the material parameters of the model LDs were carefully determined by water retention tests and triaxial tests, soil-water-air coupling finite element analyses were conducted to simulate the flume tests as boundary value problems. Furthermore, after confirming the validity of the proposed numerical method, numerical tests were also conducted to investigate the influence of the water head rising rate on the failure mechanism of the LDs. The following conclusions can be made:

1. The water retentivity of the LD materials is mainly determined by the fines content, and an increase in the fines content will lead to an increase in the water retentivity. Due to the ink-bottle effect, the moisture hysteresis of the LD materials is related to the distribution of void radii. Under the same void ratio, because the well-graded LD material has relatively uniformed pore sizes, its moisture hysteresis is not obvious compared with the gap-graded LD material. 2. It is known from the flume tests on the model LDs that the soil properties have a great influence on the failure mode of the model LDs. Under the present test condition, that is, under a constant-inflow seepage loading condition, the failure mode of the sand LD is a typical sliding failure, the failure mode of the

X. Xiong et al. / Soils and Foundations 58 (2018) 1133–1152

3.

4.

5.

6.

7.

well-graded LD is a typical overtopping failure and the failure mode of the gap-graded LD is a typical piping failure. By comparing the calculated results with the test results, it was found that based on a rational constitutive model and the proper determination of the material parameters with sufficient element tests, the proposed numerical method can describe the different failure modes of the LDs with satisfactory accuracy. The calculated failure modes of the LDs were exactly the same as those observed in the flume tests for the different ground materials. Nevertheless, the comparison between the tests and the calculations was only at a qualitative level. Further research should be done to validate the applicability at a quantitative level. Based on the numerical calculations, a method of judging the failure mode of LDs was proposed, which can provide some applicable information for evaluating the stability of LDs. As the sand material had low strength, it is the decrease in the average effective stress due to the increase in excessive pore water pressure that caused the slope failure of the sand LD. Although the rate of the rise in water head had little influence on the failure mechanism of the sand LD under seepage loading, its decrease may have reduced the stability of the sand LD. The rise in water head caused the settlement of the wellgraded LD and hastened the overtopping, but did not trigger a failure, because of the relatively high strength of the well-graded material. Overtopping was the main reason for the erosion of the slope surface on the downstream side and was finally the cause of the failure of the slope. The development of a seepage line in the gap-graded LD was the slowest among the three LDs. As the difference in pore ratios between the N and Nr of the gap-graded LD was the largest, the rise in the water head firstly generated many horizontal cracks within the LD due to shrinking, caused by the increase in the degree of saturation, and then these cracks were broadened and connected with each other, which eventually led to the formation of a piping channel. Finally, when the outflow of the LD flowing through the piping channel was equal to the inflow, the gap-graded LD failed with the piping mode. Since the water head rising rate influenced the development of the seepage line, both the material property of the soil and the water head rising rate played important roles in the failure mechanism of the gap-graded LD. Due to limited data obtained from the model flume tests and the lack of verification of the material parameters in the element tests under an unsaturated condition, the numerical analyses were only able to provide a qualitative description of the failure mechanism of LDs subjected to seepage loading. Further research should be done in the future, using both model tests and element tests, to improve the accuracy of the arguments in this paper.

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Acknowledgements This research was substantially supported by the Natural Science Foundation of China (Nos. 41372272 and 41402257), the Shanghai Sailing Program (14YF1403800), and the China Scholarship Council (201706260252). Financial support from Grant-in-Aid Scientific Research (B), No. 17H03304, JSPS, is also greatly appreciated. Appendix A. See Figs. A1 and A2.

Fig. A1. e–ln p relation considering moving up of N.C.L. and C. S.L. due to instauration (Zhang and Ikariya, 2011).

Fig. A2. Image of WRC of unsaturated soil (Zhang and Ikariya, 2011).

References Alonso, E.E., Gens, A., Josa, A., 1990. A constitutive model for partially saturated soils. Ge´otechnique 40 (3), 405–430. Araei, A.A., Soroush, A., Rayhani, M., 2010. Large-scale triaxial testing and numerical modeling of rounded and angular rockfill materials. Scientia Iranica. Trans. A Civ. Eng. 17 (3), 169. Awal, R., Nakagawa, H., Kawaike, K., Baba, Y., Zhang, H., 2008. Experimental study on prediction of failure mode of landslide dams.

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X. Xiong et al. / Soils and Foundations 58 (2018) 1133–1152

In: Proceedings of Fourth International Conference on Scour and Erosion (ICSE-4), pp. 655–660. Awal, R., Nakagawa, H., Kawaike, K., Baba, Y., Zhang, H., 2009. Three dimensional transient seepage and slope stability analysis of landslide dam. Disaster Prevent. Res. Inst. Annu. B 52, 689–696. Awal, R., Nakagawa, H., Kawaike, K., Baba, Y., Zhang, H., 2011. Experimental Study on Piping Failure of Natural DAM. J. Jpn. Soc. Civil Eng. Ser. B1 (Hydraul. Eng.) 67 (4), I_157–I_162. Braun, A., Cuomo, S., Petrosino, S., Wang, X., Zhang, L., 2017. Numerical SPH analysis of debris flow run-out and related river damming scenarios for a local case study in SW China. Landslides, 1–16. Casagli, N., Ermini, L., Rosati, G., 2003. Determining grain size distribution of the material composing landslide dams in the Northern Apennines: sampling and processing methods. Eng. Geol. 69 (1–2), 83–97. Chen, S., Lin, T., Chen, C., 2015. Modeling of natural dam failure modes and downstream riverbed morphological changes with different dam materials in a flume test. Eng. Geol. 188, 148–158. Chiu, C.F., Ng, C.W., 2003. A state-dependent elasto-plastic model for saturated and unsaturated soils. Ge´otechnique 53 (9), 809–830. Costa, J.E., Schuster, R.L., 1988. The formation and failure of natural dams. Geol. Soc. Am. Bull. 100 (7), 1054–1068. Cui, Y.J., Delage, P., 1996. Yielding and plastic behaviour of an unsaturated compacted silt. Ge´otechnique 46 (2), 291–311. Dong, J.J., Tung, Y.H., Chen, C.C., Liao, J.J., Pan, Y.W., 2009. Discriminant analysis of the geomorphic characteristics and stability of landslide dams. Geomorphology 110 (3), 162–171. Dong, J.J., Tung, Y.H., Chen, C.C., Liao, J.J., Pan, Y.W., 2011. Logistic regression model for predicting the failure probability of a landslide dam. Eng. Geol. 117 (1), 52–61. Duman, T.Y., 2009. The largest landslide dam in Turkey: Tortum landslide. Eng. Geol. 104 (1), 66–79. Ermini, L., Casagli, N., 2003. Prediction of the behaviour of landslide dams using a geomorphological dimensionless index. Earth Surf. Proc. Land. 28 (1), 31–47. Gregoretti, C., Maltauro, A., Lanzoni, S., 2010. Laboratory experiments on the failure of coarse homogeneous sediment natural dams on a sloping bed. J. Hydraul. Eng. 136 (11), 868–879. Kazama, M., Kataoka, S., Uzuoka, R., 2012. Volcanic mountain area disaster caused by the Iwate-Miyagi Nairiku Earthquake of 2008, Japan. Soils Found. 52 (1), 168–184. Kohgo, Y., Nakano, M., Miyazaki, T., 1993. Theoretical aspects of constitutive model for unsaturated soils. Soils Found. 33 (4), 49–63. Kurimoto, Y., Kobayashi, S., Tsunemoto, T., Zhang, F., 2017. Uniaxial compression test of unsaturated Masado under constant degree of saturation condition and its modeling. Int. J. GEOMATE 12 (31), 83– 87. Lu, N., Alsherif, N., Wayllace, A., Godt, J., 2014. Closing the loop of the soil water retention curve. J. Geotech. Geoenviron. Eng. 141 (1), 1. Loret, B., Khalili, N., 2002. An effective stress elastic–plastic model for unsaturated porous media. Mech. Mater. 34 (2), 97–116. Meyer, W., Schuster, R.L., Sabol, M.A., 1994. Potential for seepage erosion of landslide dam. J. Geotech. Eng. 120 (7), 1211–1229. Maqsoud, A., Bussie`re, B., Aubertin, M., Mbonimpa, M., 2012. Predicting hysteresis of the water retention curve from basic properties of granular soils. Geotech. Geol. Eng. 30 (5), 1147–1159. Ohno, S., Kawai, K., Iizuka, A., Tachibana, S., Kanazawa, S.I., Ohta, H., 2013. Elasto-plastic constitutive model for unsaturated soils with subloading surface concept. Geotech. Pred. Practice Deal. Geohazards, 215–228. Okeke, A.C.U., Wang, F., 2016. Hydromechanical constraints on piping failure of landslide dams: an experimental investigation. Geoenviron. Disasters 3 (1), 4.

Schuster, R.L., Alford, D., 2004. Usoi landslide dam and lake sarez, Pamir mountains, Tajikistan. Environ. Eng. Geosci. 10 (2), 151–168. Shi, Z.M., Xiong, X., Peng, M., Lin, M.L., 2015a. Stability analysis of landslide dam with high permeability region: a case study of Hongshihe landslide dam. J. Hydraul. Eng. 46 (10), 1162–1171 (In Chinese). Shi, Z.M., Guan, S.G., Peng, M., Xiong, X., 2015b. Research on the influence of material permeability to landslide dam seepage stability. 10th Asian Regional Conference of IAGE, pp. 1–7. Sun, D.A., Cui, H.B., Matsuoka, H., Sheng, D.C., 2007. A threedimensional elastoplastic model for unsaturated compacted soils with hydraulic hysteresis. Soils Found. 47 (2), 253–264. Sun, G., Huang, Y., Li, C., Zheng, H., 2016. Formation mechanism, deformation characteristics and stability analysis of Wujiang landslide near Centianhe reservoir dam. Eng. Geol. 211, 27–38. Takahashi, T., Kuang, S.F., 1988. Hydrograph prediction of debris flow due to failure of landslide dam. Annuals of Disaster. Prevent. Res. Inst. 31 (B-2), 601–615. Uno, T., Sato, T., Sugii, T., Tsuge, H., 1990. Method of test for permeability of unsaturated sandy soil with controlled air pressure. J. Geotech. Eng. JSCE 418, 115–124 (in Japanese). Van Genuchten, M.T., 1980. A closed-form equation for predicting the hydraulic conductivity of unsaturated soils. Soil Sci. Soc. Am. J. 44 (5), 892–898. Wang, F., Dai, Z., Okeke, C.A.U., Mitani, Y., Yang, H., 2018. Experimental study to identify premonitory factors of landslide dam failures. Eng. Geol. 232, 123–134. Wang, G., Huang, R., Kamai, T., Zhang, F., 2013. The internal structure of a rockslide dam induced by the 2008 Wenchuan (M w 7.9) earthquake, China. Eng. Geol. 156, 28–36. Wei, S., Zhu, J.G., Wang, J.J., Yu, X.J., 2005. Consolidated-undrained triaxial test study on steady state strength of sand. Chin. J. Rock Mech. Eng. 24 (22), 4151–4157 (in Chinese). Weidinger, J.T., 2011. Stability and life span of landslide dams in the Himalayas (India, Nepal) and the Qin Ling Mountains (China). Natl. Artif. Rockslide Dams, 243–277. Xiong, Y., Bao, X., Ye, B., Zhang, F., 2014. Soil–water–air fully coupling finite element analysis of slope failure in unsaturated ground. Soils Found. 54 (3), 377–395. Yang, Y., Cao, S.Y., Yang, K.J., Li, W.P., 2015. Experimental study of breach process of landslide dams by overtopping and its initiation mechanisms. J. Hydrodyn. Ser. B 27 (6), 872–883. Zhao, Y.H., 2008. Seepage control of the earthquake-induced barrier bar of Xiaonanhai Reservoir. Adv. Sci. Technol. Water Resour. 28 (5), 39– 44 (In Chinese). Zhang, F., Ikariya, T., 2011. A new model for unsaturated soil using skeleton and degree of saturation as state variables. Soil Found 51 (1), 67–81. Zhang, W.J., Maeda, K., 2015. Numerical simulations of slope and levee failure under heavy rainfall using the three-phase SPH model. J. Jpn. Soc. Civ. Eng. Ser. A2 (Appl. Mech. [AM]) 70 (2), 483–494. Zhang, W.J., Maeda, K., Saito, H., Li, Z., Huang, Y., 2016. Numerical analysis on seepage failures of dike due to water level-up and rainfall using a water–soil-coupled smoothed particle hydrodynamics model. Acta Geotech. 11 (6), 1401–1418. Zhang, X.D., He, C.R., Wang, C., Xiao, J., Chen, Y.J., Zhao, H.F., 2009. Study on stress – strain characteristics of gravel - modified gravel soil in core walls. Sichuan Archit. 29 (1), 69–72 (in Chinese). Zhu, P.Y., Wang, C.H., Wang, Y.C., 2003. Large-scale landslide-debris avalanche in Tibet, China, (2) formation of an exceptionally serious outburst flood from a landslide dam in Tibet. Landslide News. Int. Newslett. Jpn. Landslide Soc. 14 (15), 23–25.