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Article

Stratum Responses and Disaster Mitigation Strategies During Pressurized Pipe Bursts: Role of Geotextile Reinforcement

1
Department of Geotechnical Engineering, College of Civil Engineering, Tongji University, 1239 Siping Road, Shanghai 200092, China
2
Shanghai Geotechnical Investigations & Design Institute Co., Ltd., 38 Shuifeng Road, Shanghai 200093, China
*
Author to whom correspondence should be addressed.
Buildings 2025, 15(15), 2696; https://doi.org/10.3390/buildings15152696
Submission received: 1 July 2025 / Revised: 24 July 2025 / Accepted: 28 July 2025 / Published: 30 July 2025
(This article belongs to the Section Building Energy, Physics, Environment, and Systems)

Abstract

Urban subsurface pipeline bursts can induce catastrophic cascading effects, including ground collapse, infrastructure failure, and socioeconomic losses. However, stratum responses during the erosion cavity expansion phase and corresponding disaster mitigation strategies have rarely been researched. In this study, a numerical model validated through experimental tests was employed to investigate the effects of internal water pressures, burial depths, and different geotextile-based disaster mitigation strategies. It was revealed that a burial depth-dependent critical internal water pressure governed the erosion cavity expansion, and a predictive equation was derived based on the limit equilibrium theory. Higher internal water pressure accelerated the erosion cavity expansion and amplified the stratum stress within a range of twice the diameter D. Increased burial depth d reduced peak ground heave but linearly expanded the heave zone range, concurrently elevating the overall stratum stress level and generating larger stress reduction zones (i.e., when d / D = 3.0, the range of the stress reduction zone was 8.0D). All geotextile layout configurations exhibited different disaster mitigation effects (the peak ground heave was reduced by at least 15%). The semi-circular closely fitted configuration (SCCF) optimally restricted the expansion of the erosion cavity, reduced the stratum displacement (i.e., 39% reduction in the peak ground heave), and avoided stress concentration. Comprehensive analysis indicated that SCCF was suited for low-pressure pipelines in deformation-sensitive stratum and semi-circular configuration (SC) was suitable for deformation-insensitive pipeline sections. These findings provide actionable insights for tailoring mitigation strategies to specific operational risks.

1. Introduction

The occurrence of frequent and abrupt roadway cave-in accidents significantly heightened the safety risks of traffic, pedestrians, and existing structures and facilities. To identify the triggering factors of roadway cave-in accidents, a number of statistical databases have been created [1,2,3,4,5], indicating that leakage pipelines buried below roadways were a significant cause. According to the official statistical result, 28.3% of water supply pipelines and 18.9% of drainage pipelines in China have been in operation for over 20 years [6], which has a high risk of material degradation. Aged pipelines were susceptible to disturbance, including adjacent construction, earthquakes, and freeze–thaw [7,8,9,10,11]. The disturbance can lead to pipeline defects and leakage, whose interaction with adjacent underground infrastructures (e.g., ongoing shield tunneling projects, foundation piles) exacerbates adverse effects on the urban environment and escalated economic losses [12,13].
The road cave-in accidents caused by pipeline leakage can be categorized into three types based on erosion mechanism: exfiltration, infiltration, and cyclic infiltration–exfiltration. Exfiltration erosion occurred in defective pipelines when the internal pressure head was higher than the phreatic water level [4,14,15,16]. Infiltration erosion occurred in defective pipelines when the internal pressure head was lower than the phreatic water level [17,18,19]. Cyclic infiltration–exfiltration erosion occurred in defective pipelines as the internal pressure head alternately rose above and dropped below the phreatic water level [20,21,22,23]. As a special form of exfiltration erosion, pipe bursts were typically characterized by sudden occurrence and higher hydraulic pressure, which led to the disturbance of soil particle distribution, decline in stratum stability, and localized large-scale deformation accompanied by stress redistribution (see Figure 1).
Using experimental model tests, many studies have been carried out to explore the failure mechanism of pipe bursts [24,25]. Researchers employed vertically upward jets to simulate stratum induced by pipe bursts, dividing the process into three distinct phases, i.e., seepage diffusion, erosion cavity expansion, and soil fluidization [25,26]. Alsaydalani and Clayton identified factors affecting the onset of fluidization and its development through data from small-scale experiments and simple analysis [27]. With the help of an experimental model test, Van Zyl et al. investigated the extent of the fluidized zone and the head drop within the bed [28]. Mena et al. explored the effects of particle sizes, burial depth, and defect open on critical flow rate via high-speed videos [29]. Schulz et al. presented the theoretical and experimental results for the evolution of cavities generated by upward vertical leakage jets in porous media [30]. Numerical simulation was commonly employed to analyze the phenomena observed in experimental tests, providing insight into the underlying mechanisms.
The Finite Element Method (FEM) demonstrated strong capabilities in modeling the deformation of continuous media, particularly in soft clay stratum [31,32,33]. However, FEM has limitations in simulating the response of discontinuous stratum induced by pipe bursts in sandy stratum [18]. In contrast to FEM, the Discrete Element Method (DEM) exceled in modeling the discrete nature of granular materials (e.g., sand, gravel, and cobble) [34,35]. DEM could reasonably simulate particles’ interaction forces, migration paths, and large deformation behavior during the process of sandy stratum instability [4,14,18,19]. Long and Tan established a numerical model to analyze the stress evolution and developed the stress arching model in granular soil layer [18]. With the help of DEM, Nguyen and Nguyen investigated the effects of burial depth, defect size, and particle size distribution on the critical jet velocity [36]. Several DEM-based studies implied that optimizing the particle gradation and utilizing dry sand can enhance stratum stability, thereby mitigating the losses [14,36]. Concurrently, the application of geotextiles can protect buried pipes and minimize the consequences of pipe bursts [37]. The existing studies related to pipe bursts primarily focused on phenomena and mechanisms in the soil fluidization phase, whereas the mechanism responses of the stratum in the phase of expansion of the erosion cavity have rarely been explored. Furthermore, limited research has been conducted on the positive effects of disaster mitigation strategies (e.g., different geotextile layout configurations) on pipe bursts.
In this study, a validated numerical model was developed to investigate the stratum responses to pipe bursts. Subsequently, a series of numerical simulations were conducted and the effects of internal water pressure, burial depth, and geotextile-based disaster mitigation strategies were examined from both macroscopic and mesoscopic perspectives. The findings could provide critical references for disaster mitigation in urban pipeline infrastructure design.

2. Establishment and Validation of Model

2.1. Experimental Test of Pipe Burst

2.1.1. Experimental Apparatus

As illustrated in Figure 2, the apparatus included a cuboid box, a PVC pipe (Haosheng Plastic Co., Ltd., Shantou, Guangdong, China), a constant-head water tank, and a camera. The box consisted of one soil chamber (600 mm × 200 mm × 600 mm) and two water tanks (100 mm × 200 mm × 600 mm), with the front plate being transparent to observe the experimental phenomena. Sands (Xiamen ISO Standard Sand Co., Ltd., Xiamen, Fujian, China) susceptible to seepage erosion were chosen as the stratum medium. Water tanks on both sides serve to simulate the boundary conditions of the phreatic level. Orifice plates were overlaid gauze, designed to permit the passage of water while retaining soil particles. A 50 mm diameter PVC pipe was utilized to simulate a burial pipeline. The pipe was positioned 1 mm away from the front plane to model a full-section pipe burst. The constant-head water tank was connected to the pipeline for water supply, and the internal water pressure was regulated by adjusting the tank’s height. The camera was set in front of the cuboid box to record the experimental phenomena.

2.1.2. Experimental Procedures

The experimental procedures can be divided into three stages.
(a)
Stage I—experimental preparation. The dried sand was layered into the soil chamber and each layer was compacted prior to adding the subsequent layer (30 mm thick per layer). When the total soil layer thickness reached 250 mm, a PVC pipe was installed at the designated position. To enhance the visibility of stratum movements, four layers of red-colored sand were positioned along the inner wall of the soil chamber at vertical of 25 mm. Thereafter, the constant-head water tank was positioned at the predetermined height. Finally, the water tanks on both sides were filled and the water level was maintained at 250 mm.
(b)
Stage II—experimental process. The camera was turned on first. Then, the valve was opened. The ground heave and soil fluidization were recorded during the process of pipe bursts. The test was terminated when the erosion cavity remained stable for a long time or soil fluidization occurred.
(c)
Stage III—result analysis. Based on the recording results, qualitative analyses of the evolution process of pipe burst were carried out.
Figure 2. Schematic diagram of experimental apparatus.
Figure 2. Schematic diagram of experimental apparatus.
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2.2. Establishment of Model

To systematically investigate the stratum responses to pipe bursts, numerical simulation was employed. By implementing DEM, a refined model presenting the stratum’s large deformation was developed. Compared to FEM, DEM showed distinct advantages in simulating discontinuous and large deformation behaviors of the stratum. Referring to Tupa and Palmeira’s approach [38], the erosion cavity formed by pipe bursts was idealized as an outward uniformly pressurized circular cavity. With the help of PFC 2D 5.0, a numerical model was developed to investigate the deformation and stress redistribution of the stratum during the erosion cavity expansion phase. The disaster mitigation effectiveness of geotextile reinforcement was also evaluated. To ensure the reliability of the model, the results of the self-designed experimental test and numerical simulation were compared. Based on the validated model, a series of variable-parameter analyses was carried out.

2.2.1. Model Parameters

The appropriate calibration of micro-parameters for both the soil and geotextile was critical to guarantee the reliability of the model and accuracy of the computational results.
The soil parameters employed in the numerical simulation were systematically summarized in Table 1. Soil parameters were determined by comparing the results from the biaxial compression tests shown in Figure 3a and conventional triaxial test data (Figure 3b). The soil sample used in the conventional triaxial test was the same soil that was used for the experimental model test in Section 2.1. Commonly used particle contact constitutive models included the linear contact model, the linear bond contact model, and the linear parallel bond contact model. Among them, the linear contact model could only transmit pressure, the linear bond contact model was capable of transmitting pressure and tensile forces, and the linear parallel bond contact model was able to transmit pressure, tensile forces, and moments. Considering these, the linear contact model was selected to simulate the interaction between soil particles [39,40]; the numerical model adopted local damping mechanism to dissipate particles’ kinetic energy and accelerate stratum’s equilibrium [41,42,43]; and the parallel bond model was employed for particle–particle interactions within the geotextile [44,45,46]. Based on the existing studies in the literature, to reduce the number of soil particles in the numerical model to improve computational efficiency, particle sizes were reasonably scaled up [14,18,19]. In this study, particle radii ranged from 4 to 26 mm. Particle size upscaling can lead to slight fluctuations in the results; nevertheless, the simulation outcomes remain valuable and accurate.
The geotextile parameters are presented in Table 1. To accurately simulate the interaction between the soil and geotextile, a 2D interface direct shear test model was simulated by DEM to determine the micro-parameters (Figure 3c, where σ is the normal stress and ν is the velocity of the wall). During the shear test, the upper wall was controlled by a servo system to maintain constant normal stress, side walls was fixed to restrict the soil’s lateral displacement, and the geotextile was set to displace together with the wall beneath it. After assigning a certain uniform velocity for the bottom wall, the interface direct shear test began and the relationship between stress and strain could be obtained. Feng et al. conducted a series of experimental model tests and DEM numerical simulations, obtaining the shear strain curves of the geotextile in sandy stratum [46]. As shown in Figure 3d, the results were compared against the existing study, confirming the validity of the parameters [46]. The observed fluctuations in the results primarily stemmed from particle upscaling effects.
Figure 3. Validation of model parameters: (a) biaxial test in 2D DEM model; (b) comparison between experimental and numerical results; (c) sand–geotextile interface direct shear test in 2D DEM model; (d) comparison between numerical results and existing literature [46].
Figure 3. Validation of model parameters: (a) biaxial test in 2D DEM model; (b) comparison between experimental and numerical results; (c) sand–geotextile interface direct shear test in 2D DEM model; (d) comparison between numerical results and existing literature [46].
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2.2.2. Internal Water Pressure Control Method

As the interaction between fluid and soil particles was frequent and complex in the erosion cavity expansion phase, it was difficult to calculate the flow field. To satisfy the need to analyze the mechanism responses of the stratum and effects of disaster mitigation strategies, the erosion cavity was simplified as a deformable cylinder that applies uniform pressure in all directions [37]. Since it was difficult for wall elements to apply stress directly to particles, equivalent velocity was imposed at the wall vertices as a substitute approach. By adjusting the velocity of the wall vertices dynamically, the stress exerted by the erosion cavity could be controlled. The velocity calculation equation at the vertex of the wall under the target pressure is shown in Equation (1):
v p t = G σ t t σ c t
The calculation of the contact force increment induced by the movement of the wall vertex within one timestep could be expressed as follows:
Δ σ s p = k n v p t Δ t A
To maintain stability, a safety factor was employed to ensure that the increment of the average node pressure per timestep is less than the difference between the target value and the current stress.
Δ σ s p α σ t t σ c t
Substituting Equations (1) and (2) into Equation (3) and rearranging, the resulting expression was obtained as follows:
G < α A k n t
where v p t is the velocity of wall vertices to be solved; σ t t and σ c t are the target stress and current stress of the wall element, respectively; G is a constant parameter describing the relationship between v p t and stress difference σ t t σ c t ; Δ σ s p is the variation in average stress within per timestep Δ t ; k n is the stiffness of a particle–wall contact; A is the total particle–wall contact area; and α denotes the safety factor.
In this study, the value of α was chosen to be 0.2 to improve the accuracy of pressure controlling. Then, G can be determined by Equation (4).

2.2.3. Procedures of Numerical Simulation

The procedures of numerical simulation can be divided into three major stages:
(a)
Stage I—preparation of the numerical simulation. The dimension of the model as shown in Figure 4 (where d is the burial depth of pipelines and D is the diameter of the initial erosion cavity, D = 1.00 m). The specific height depended on burial depth. The soil particles were initially generated within the computational domain and subsequently compacted using the Multi-layer with Undercompaction Method to achieve a target porosity of 0.30 [47]. The process continued until soil particles fully occupied the designated simulation area. As for the geotextile, the following sequence was executed when the height of the generated stratum reached the target depth of the geotextile: (1) removal of soil particles which took up the geotextile’s position; (2) generation of geotextile particles; and (3) establishment of effective contact between geotextile particles and adjacent soil particles. Sequentially, the gravitational field was applied to all particles. To record the variation in stress during the simulation, the measurement regions were defined in advance.
(b)
Stage II—simulation of the erosion cavity expansion phase. Soil particles occupying the position of the erosion cavity were removed and a circle of diameter D was generated by wall elements. Following the method in Section 2.2.2, the vertex velocity of wall elements was automatically adjusted to simulate internal water pressure under different working conditions. The ground heave and the expansion of the erosion cavity could be simulated.
(c)
Stage III—data processing. Ground heave, stratum stress, and the microscopic properties of particles across different working conditions were extracted and summarized. Through comparative analysis, the effects of internal water pressure and burial depths during the pipe burst process were identified, and geotextile performance in disaster mitigation was evaluated.
Figure 4. PFC 2D numerical model. (The colors of the strata were to facilitate observation and they had no special meaning.)
Figure 4. PFC 2D numerical model. (The colors of the strata were to facilitate observation and they had no special meaning.)
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2.3. Model Validation

To further validate the applicability of the established numerical model for simulating stratum responses during the erosion cavity expansion phase of pipe bursts, a systematic comparative analysis was performed between numerical simulation data and experimental test results.
The comparative results are shown in Figure 5. During the preliminary erosion cavity expansion (Figure 5a,b) and the progressive expansion nearing collapsed (Figure 5c,d), the ground displacement patterns and area development of the erosion cavity exhibited remarkable consistency. As the erosion cavity collapsed (Figure 5e), discrepancies were observed between numerical simulation and the experimental test. The differences stemmed from the inability to simulate the process of soil particle detachment caused by water flow erosion, which led to a reduction in overburden thickness and subsequent failure. The removal of the wall elements resulted in a funnel-shaped soil collapse (Figure 5f), which was similar to the results reported by Long and Tan [17]. This study mainly focused on the mechanical response of the stratum before the erosion cavity collapsed. The consistency between the numerical simulation and the experimental test validated the effectiveness of the model.

2.4. Numerical Cases

To explore the effects of the internal water pressure, burial depth, and different geotextile layout configurations on the mechanical response of the stratum during the erosion cavity expansion phase, systematic numerical simulations were conducted. Table 2 summarizes the details of the numerical cases. Based on the relevant literature [48], the range of internal water pressure was 0.15–0.30 MPa. The burial depth was normalized by D. According to engineering experience, four geotextile layout configurations were implemented (Figure 6), i.e., horizontal (HOR), broken-line (BL), semi-circular (SC), and semi-circular closely fitted (SCCF). Cases 1–5 were designed to study the effects of internal water pressure through intra-group comparative analysis. Meanwhile the effects of burial depth can be evaluated by maintaining identical internal water pressure conditions. Together with cases 2 and 6–9, the disaster mitigation performance of different geotextile layout configurations can be investigated.

3. Results and Discussion

3.1. Effects of Internal Water Pressures

3.1.1. Development of Stratum Movements

To investigate the evolution of stratum movements during the erosion cavity expansion phase, a series of numerical simulations was conducted. A comparative analysis of cases 1–5 revealed that internal water pressures primarily influenced the expansion rate and final dimensions of the erosion cavity but had little influence on the evolution patterns of the stratum movements.
Figure 7a illustrated the relationship of ground heave with different erosion cavity area growth ratios. The data was extracted from the case where d /D = 2.0, P = 0.20 MPa and no geotextile was laid out. As the area of the erosion cavity increased, the peak value of ground heave presented significant growth. The ground heave curve presented a roughly symmetric unimodal distribution on both sides, conforming to the characteristics of the Gaussian curve. Moreover, the Gaussian equation was characterized by extensive applicability and facile interpretability in fitting the distribution of the ground displacements [18,49,50]. To better characterize the deformation pattern, a Gaussian function (Equation (5)) was employed to describe the ground heave distribution, with the fitted curve and inflection points shown in Figure 7b.
s x = α e x 2 2 β 2
where α is the maximum ground heave and β is the horizontal distance from the origin to inflection point. The fitting parameters are listed in Table 3, demonstrating satisfactory fitting performance with the coefficient of determination R 2   > 0.85. Figure 7b illustrates the variation of β   under different growth ratios of the erosion cavity area. As the growth ratio increased, parameter β first decreased and then increased, reaching its minimum value when the growth ratio was 80%. This transition occurred because the erosion cavity initially grew in the vertical direction due to the horizontal resistance from the stratum. However, when the erosion cavity became large enough, the horizontal resistance would become ineffective and the cavity would begin to grow in all directions.
Figure 8a–d illustrates the development of stratum displacement contour with P = 0.20 MPa and d /D = 2.0. With the increase in timesteps, the erosion cavity continued to expand, which resulted in sustained increases in both the wedge-shaped displacement zone’s area and the slip band angle. To explore the relationship between particle movement characteristics and positions, typical particles located in the main displacement zone were extracted to analyze the variation in displacement increments and vertical velocities. Figure 8e,f shows the variation in soil particle displacement increments and vertical velocities with timesteps. It was revealed that the displacement increments were predominantly vertical. The horizontal displacement increments increased as the particles’ position approached the edge of the wedge-shaped zone. Similarly, the vertical velocities of the particles depended on the relative position to the center of the wedge-shaped zone.
Table 3. Fitted results of ground heave for different growth rates of erosion cavity area.
Table 3. Fitted results of ground heave for different growth rates of erosion cavity area.
Growth Rate of Erosion Cavity Area (%)Fitted Equation α β R 2
20 s x = 0.11 e x 2 2 × 1.496 2 0.1101.4960.869
40 s x = 0.146 e x 2 2 × 1.416 2 0.1461.4160.911
60 s x = 0.233 e x 2 2 × 1.271 2 0.2331.2710.962
80 s x = 0.294 e x 2 2 × 1.253 2 0.2941.2530.959
100 s x = 0.344 e x 2 2 × 1.266 2 0.3441.2660.960
120 s x = 0.394 e x 2 2 × 1.309 2 0.3941.3090.972
140 s x = 0.446 e x 2 2 × 1.344 2 0.4461.3440.980
160 s x = 0.488 e x 2 2 × 1.358 2 0.4881.3580.983
180 s x = 0.533 e x 2 2 × 1.384 2 0.5331.3840.984
200 s x = 0.576 e x 2 2 × 1.403 2 0.5761.4030.987
Figure 7. Ground heave under various working conditions: (a,b) different growth ratios of erosion cavity area; (c,d) different burial depths; (e,f) different geotextile layout configurations.
Figure 7. Ground heave under various working conditions: (a,b) different growth ratios of erosion cavity area; (c,d) different burial depths; (e,f) different geotextile layout configurations.
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3.1.2. Variation in Stratum Stress

To analyze the effects of internal water pressures on the stratum stress distribution, stress along both vertical and horizontal central axes were extracted at the same timestep. Burial depth was normalized by D and stress was normalized by γ and D, as shown in Figure 9a,b and Figure 10a,b ( d /D = 2.0, no geotextile).
Figure 9a,b shows the vertical and horizontal stress variations along the vertical central axis, respectively. It could be found that increasing internal water pressures led to a marked enhancement of vertical stress peaks in both overlying and underlying soils surrounding the erosion cavity. As shown in Table 4, the maximum vertical stress in the overlying soil increased to 3 times the initial stress, while the underlying stratum reached 2.5 times the initial stress. The regions experiencing increased stress extended from 1D above the erosion cavity to 2D below it. The horizontal stress distribution showed similar responses to increasing pressures as the vertical stress increased, but with lower stress magnitudes and a narrower stress-elevated zone. An analysis of Figure 10a,b revealed that the vertical stress exhibited insignificant variation along the horizontal direction. However, localized zones of vertical stress reduction were observed and marked in gray, which was caused by unloading effects due to the detachment of the overlying stratum. As for the distribution of horizontal stress, a dramatic enhancement throughout the computational domain was induced by the internal water pressure. Compared to the initial state, the peak horizontal stress increased by a factor of 6.5. This phenomenon arose primarily from lateral soil compression driven by erosion cavity expansion.
Figure 8. Stratum displacement contour and typical particles movement: (ad) stratum displacement contour at different timesteps; (e) particle displacement increments; (f) particle velocities development (1 timestep =   5 × 10 5   s ).
Figure 8. Stratum displacement contour and typical particles movement: (ad) stratum displacement contour at different timesteps; (e) particle displacement increments; (f) particle velocities development (1 timestep =   5 × 10 5   s ).
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3.2. Effects of Burial Depths

3.2.1. Development of Stratum Movements

Figure 7c illustrates the distribution of ground heave at varying burial depths with P = 0.20 MPa and no geotextile. The curves were plotted under the condition of the same erosion cavity area. The maximum ground heave demonstrated progressive attenuation with increasing burial depths. To quantify the variation in the ground heave, all the ground heave profiles were fitted using Equation (5). The fitted curves and inflection points are shown in Figure 7d, with the corresponding mathematical expressions and parameters detailed in Table 5, No. 1–5. All regression analyses yielded R 2 > 0.95, indicating excellent fitting accuracy. Increased burial depths significantly reduced the gradient of the fitted curves, exhibiting a decrease in peak ground heave and an increase in the width of the influenced range. As presented in Figure 7d, β showed a linear relationship with burial depth. Quantitative analysis revealed a linear correlation coefficient of 0.969 between these parameters.
Figure 11 presented the effects of burial depth on the development of the erosion cavity and stratum stability. The growth of the cavity can be effectively restricted by a greater burial depth, due to the need to overcome higher vertical stratum stress. As marked in Figure 11a–d, an increase in burial depth led to an increase in the angle of the wedge-shaped zones, which had positive effects on reducing the extent of the stratum deformation. It is important to note that the stratum utilized in this study was a homogeneous sandy formation. In heterogeneous or layered soils, the characteristics of the wedge zone could exhibit variations.
Figure 9. Stress variation in vertical direction under various working conditions: (a) vertical stress variation under different internal water pressures; (b) horizontal stress variation under different internal water pressures; (c) vertical stress variation under different burial depths; (d) horizontal stress variation under different burial depths.
Figure 9. Stress variation in vertical direction under various working conditions: (a) vertical stress variation under different internal water pressures; (b) horizontal stress variation under different internal water pressures; (c) vertical stress variation under different burial depths; (d) horizontal stress variation under different burial depths.
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Figure 10. Stress variation in horizontal direction under different working conditions: (a) vertical stress variation under different internal water pressures (localized zones of vertical stress reduction were marked in gray); (b) horizontal stress variation under different internal water pressures; (c) vertical stress variation under different burial depths; (d) horizontal stress variation under different burial depths.
Figure 10. Stress variation in horizontal direction under different working conditions: (a) vertical stress variation under different internal water pressures (localized zones of vertical stress reduction were marked in gray); (b) horizontal stress variation under different internal water pressures; (c) vertical stress variation under different burial depths; (d) horizontal stress variation under different burial depths.
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3.2.2. Variation in Stratum Stress

Figure 9c,d illustrates the vertical and horizontal stress distributions along the central vertical axis under various burial depths with P = 0.20 MPa and no geotextile. As presented in Table 4, the expansion of the erosion cavity triggered abrupt vertical stress increases, with peak magnitude reaching 200% of the initial stress state. The stress intensification phenomenon was extraordinary in the region from 1D above the erosion cavity to 1.5D below it. In contrast, the effects of burial depth on horizontal stress in the vertical direction were negligible, inducing only localized stress elevation within confined zones.
Increased burial depths substantially expanded the spatial domain where vertical stress became lower than the initial states (i.e., shadow region in Figure 10c), which resulted from the combined effects of localized unloading and enhanced soil arching. The horizontal stress distribution along the horizontal center axis showed consistent trends and similar magnitude ranges compared to the initial state across different burial depths, indicating a weak dependence on burial depth (Figure 10d).
Table 4. Summary of stress distribution under different working conditions.
Table 4. Summary of stress distribution under different working conditions.
Position of Measuring
Region
Stress
Direction
Internal Water Pressure, P (MPa)Burial Depth,
d/D
Normalized Peak
Value
Normalized
Affected Zone
Vertical central axisVertical0.102.01.31.0
0.152.01.71.8
0.202.02.31.2
0.252.03.04.0
0.302.03.34.4
0.201.52.13.2
0.202.02.43.5
0.202.52.12.4
0.203.01.22.2
0.203.51.81.8
Horizontal0.102.01.21.0
0.152.01.21.2
0.202.01.61.4
0.252.02.31.8
0.302.02.12.0
0.201.52.11.4
0.202.01.31.0
0.202.51.21.8
0.203.00.81.0
0.203.51.11.2
Horizontal central axisVertical0.102.00.96.0
0.152.01.04.6
0.202.01.44.4
0.252.01.43.6
0.302.02.43.2
0.201.52.50.6
0.202.01.42.3
0.202.50.85.0
0.203.01.26.5
0.203.50.88.0
Horizontal0.102.02.12.0
0.152.03.84.6
0.202.05.47.4
0.252.06.210.0
0.302.06.210.0
0.201.57.610.0
0.202.06.210.0
0.202.55.410.0
0.203.03.46.0
0.203.54.34.2
Note: The affected zone was defined as the region where the stress either doubles or drops below the initial stress. The peak value was normalized with the initial stress. The affected zone was normalized with diameter, D, of the initial erosion cavity.
Table 5. Fitted results of ground heave for different burial depths and geotextile layout configurations.
Table 5. Fitted results of ground heave for different burial depths and geotextile layout configurations.
No.Burial Depth (d/D)Geotextile Layout ConfigurationFitted Equation α β R 2
11.5No geotextile s x = 0.399 e x 2 2 × 1.058 2 0.3991.0580.969
22.0No geotextile s x = 0.354 e x 2 2 × 1.298 2 0.3541.2980.956
32.5No geotextile s x = 0.338 e x 2 2 × 1.345 2 0.3381.3450.962
43.0No geotextile s x = 0.322 e x 2 2 × 1.525 2 0.3221.5250.960
53.5No geotextile s x = 0.300 e x 2 2 × 1.735 2 0.3001.7350.952
62.0No geotextile s x = 0.415 e x 2 2 × 1.127 2 0.4151.1270.989
72.0HOR s x = 0.363 e x 2 2 × 1.351 2 0.3631.3510.987
82.0BL s x = 0.355 e x 2 2 × 1.400 2 0.3551.4000.981
92.0SC s x = 0.273 e x 2 2 × 1.537 2 0.2731.5370.971
102.0SCCF s x = 0.229 e x 2 2 × 1.300 2 0.2291.3000.965

3.2.3. Critical Internal Water Pressure

Numerical simulations of cases 1 to 5 revealed that the erosion cavity would expand into a stable state under low internal water pressure. However, when the pressure exceeded a certain threshold, the erosion continued to grow until it reached the ground surface. Based on this phenomenon, the critical internal water pressure ( P c r ) was defined as the minimum pressure required for the erosion cavity to expand and break through the ground surface, ultimately leading to stratum instability. It was observed that P c r had a strong correlation with burial depth. To determine the P c r at different burial depths, the pressure listed in Table 1 was refined and increased in increments of 0.01 MPa. The stratum was considered to have reached a stable state when the following criteria were met: (1) the stratum displacements were minimal; (2) all the wall element velocities fell below 0.001 m/s.
Figure 12a presents the influence of internal water pressures on the expansion of the erosion cavity at various burial depths. The analysis demonstrated that increased burial depth was shown to elevate P c r . This behavior can be attributed to the fact that thicker overburden layers generated higher stratum stress and greater soil compaction in the vicinity of the erosion cavity. These combined effects effectively constrained upward cavity expansion while inducing stronger lateral confinement pressure on the erosion cavity. As shown in Figure 12b, the relationship between P c r and burial depths could be fitted by a quadratic equation
According to the limit equilibrium theory [51,52,53], the derivation process involved determining the limit stress at any point on the slip surface (Equation (6)), integrating the points to obtain the total shear resistance along the slip surface, and applying vertical force equilibrium (Figure 13a) to derive the critical internal water pressure formula (Equation (7)).
τ = γ z t a n ϕ 1 + K 0 2 1 K 0 c o s 2 θ 2
      P c r = γ H D + γ H 2 t a n ψ + γ H 2 t a n ϕ t a n θ 1 + K 0 2 1 K 0 c o s 2 θ 2
Figure 11. Variation in wedge-shaped zones under different burial depths: (ad) variation in wedge-shaped zones; (e) statistical analysis of wedge-shaped zone angles.
Figure 11. Variation in wedge-shaped zones under different burial depths: (ad) variation in wedge-shaped zones; (e) statistical analysis of wedge-shaped zone angles.
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A parameter sensitivity analysis was conducted with d / D = 2.0 as an example. Based on engineering experience, the values of parameters were chosen as follows: γ = 20.0   k N / m , ψ = 20 ° , ϕ = 35 ° , and K 0 = 0.25 . As shown in Figure 11b, θ was determined to be 40 ° . When ψ = 25 ° and ψ = 30 ° , the predicted results increased by 15% and 30%, respectively. In contrast, when K 0 was 0.275 and 0.3, the results changed by only 0.3% and 0.6%. Accordingly, ψ was regarded as a highly sensitive parameter, and the rationality of this parameter should be comprehensively considered in engineering applications.

3.3. Effects of Geotextile Layout Configurations

3.3.1. Development of the Stratum Movements

Figure 7e,f illustrates the ground heave under different geotextile layout configurations with P = 0.20 MPa and d /D = 2.0. All geotextile layout configurations effectively reduced the maximum ground heave. Specifically, SC and SCCF demonstrated more significant effects, achieving 33% and 39% reductions in peak heave, respectively, compared to the scenario with no geotextile scenario. This indicated that optimized geotextile placement can effectively constrain ground deformation caused by the erosion cavity expansion, thereby mitigating the severe consequences of pipe bursts. By fitting the curves (Table 5 No. 6–10) using Equation (5) and extracting β (Figure 7f), it was observed that geotextile installation increased the influence range of ground heave. SCCF exhibited the smallest increase in the influenced range (i.e., a 15% increase compared to the no geotextile case), whereas SC resulted in a more significant increase (i.e., a 36% increase). The preceding mechanism demonstrated that the geotextile could induce stress redistribution in the stratum, leading to the extension of the ground heave range.
Figure 12. Critical water pressure ( P c r ) under different burial depths and geotextile layout configurations: (a) method to determine P c r ; (b) relationship between P c r and burial depths; (c) P c r under different geotextile layout configurations.
Figure 12. Critical water pressure ( P c r ) under different burial depths and geotextile layout configurations: (a) method to determine P c r ; (b) relationship between P c r and burial depths; (c) P c r under different geotextile layout configurations.
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Figure 13. Schematic diagrams of stratum bearing capacity calculation: (a) vertical equilibrium calculation; (b) critical stress state on shear plane.
Figure 13. Schematic diagrams of stratum bearing capacity calculation: (a) vertical equilibrium calculation; (b) critical stress state on shear plane.
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Figure 14 showed that the stratum displacement contour under various working conditions exhibited a distinct evolution process. Compared to no geotextile configuration, all four different geotextile layout configurations could effectively reduce the development of the erosion cavity and stratum movements, with SC and SSCF demonstrating better disaster mitigation performance. However, in the early stage, HOR, BL, and SC led to larger stratum displacement zones, possibly because the geotextile spread the stress over a wider area. To evaluate the disaster mitigation effects of the geotextile from a mesoscopic perspective, specific soil particles were selected to analyze the variation in vertical velocities over timesteps, as shown in Figure 15. The arrangement of the geotextile can significantly decelerate the vertical displacement of particles, thereby maintaining the stability of the stratum. Among the four geotextile arrangements mentioned above, SCCF presented the best performance in reducing particle vertical velocity and mitigating disasters.

3.3.2. Variation in Stratum Stress

Figure 16 shows the particle displacement contour, contact force chains, magnitude, and orientation of stratum stress at the same timestep under various working conditions. The thicknesses of the force chain curves were proportional to the magnitudes of contact forces. The cross symbol denoted the stress distribution in the stratum. The magnitude of stresses was associated with both the length and the color of the lines. The orientation of the cross indicated the direction of the principal stresses.
In the absence of a geotextile (Figure 16a), force chains primarily concentrated around the erosion cavity, with sparser distribution at a greater distance from it. When laying the geotextile (Figure 16b–d), stress concentration was significantly alleviated, which could be attributed to the formation of a composite structure between the geotextile and the larger scope of the stratum, thereby achieving stress dispersion. Force chains above the geotextile were significantly sparser, indicating that the geotextile could interrupt the transfer of vertical stresses generated by cavity expansion. As shown in Figure 16b–e, thick force chains appeared at the geotextile’s position, which resulted from the tensile stresses borne by the geotextile due to stratum deformation. SCCF exhibited a distinct mechanical pattern, where there was no large-scale stratum stress increase. Instead, the geotextile resisted the stresses generated by cavity expansion through its own strength (Figure 16e).
Higher stratum stresses were distributed on both sides and below the cavity. The geotextile hardly affected stress distribution. However, HOR, BL, and SC led to an increase in stratum stress within the zone from the upper cavity to the geotextile (Figure 16b–d). Except for SCCF, the other three geotextile layout configurations all led to a large-scale stress increase below the cavity, among which HOR caused the greatest impact and SC the least.

3.3.3. Variation in Critical Internal Water Pressure

For cases 6–9, the internal water pressure was increased at 0.01 MPa intervals to determine P c r , as illustrated in Figure 12c. Among four configurations, the HOR and BL configurations showed no significant increase in P c r , whereas SC and SCCF showed increases of 17% and 25% compared to the baseline, respectively. As indicated by Figure 14, the expansion of the wedge-shaped zone area was the primary factor driving critical pressure elevation in SC. For SCCF, the enhanced P c r resulted from the combined effects of zone expansion and confining action of the geotextile.

3.3.4. Internal Force and Deformation of Geotextile

The maximum internal force of the geotextiles helped practitioners determine the appropriate geotextile materials. By extracting distributions of the geotextile’s internal force, the variation in the maximum internal force with the timestep was analyzed, as shown in Figure 17a. Under the HOR and BL configurations, the overall internal force magnitudes remained relatively low. This resulted from the fact that upward-transmitted stress was insufficient to form a firmly interlocked geotextile–stratum composite. In contrast, SC presented higher stress levels, indicating its sustained role in restricting stratum displacement and redistributing stress during the erosion cavity expansion. SCCF showed the highest internal forces, suggesting stringent requirements for the geotextile’s tensile capacity.
Figure 16. Soil stress and contact force chains under various working conditions: (a) no geotextile; (b) HOR; (c) BL; (d) SC; (e) SCCF.
Figure 16. Soil stress and contact force chains under various working conditions: (a) no geotextile; (b) HOR; (c) BL; (d) SC; (e) SCCF.
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To analyze the deformation of the geotextile from a mesoscopic perspective, the vertical velocities of the geotextile particles at specific positions under different geotextile layout configurations were plotted against timesteps (Figure 17b). The vertical velocity under the HOR configuration was the highest, and it started to increase rapidly at 4 × 104 timesteps. In contrast, the vertical velocities under the other cases were generally lower and remained stable throughout, with the minimum vertical velocity observed under SCCF.
From the perspective of the geotextile’s mechanical performance and pipe burst disaster mitigation, SC presented enhanced suitability in geologically stable zones with relatively high permissible ground displacement (Table 6). In contrast, SCCF proved higher effectiveness in displacement-sensitive pipelines with low internal water pressures (Table 6). This distinction originated from fundamental differences in soil–geotextile interaction mechanisms: SC significantly reduced the peak ground heave (i.e., a 33% decrease compared to the no geotextile case) but resulted in a more extensive ground heave range (i.e., a 36% increase). Meanwhile, the internal force of the geotextile kept a stable and acceptable range under this configuration. SCCF increased the stratum’s P c r by 25% compared to the no geotextile case, which effectively restricted further expansion of the erosion cavity. This configuration leveraged the geotextile’s tensile strength to reduce the value of stratum displacement, indicating its optimal suitability for pipelines operating under lower internal water pressure.
Figure 17. Variation in geotextile’s maximum internal force and particle vertical velocities under different geotextile layout configurations: (a) maximum internal force; (b) vertical velocities of particle (1 timestep =   5 × 10 5     s ).
Figure 17. Variation in geotextile’s maximum internal force and particle vertical velocities under different geotextile layout configurations: (a) maximum internal force; (b) vertical velocities of particle (1 timestep =   5 × 10 5     s ).
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Table 6. Summary of disaster mitigation effect evaluation for different geotextile layout configurations.
Table 6. Summary of disaster mitigation effect evaluation for different geotextile layout configurations.
Geotextile
Layout
Configuration
Applicable Geological ConditionsDisaster
Mitigation Effect
Internal Water PressureSensitivity to Deformation
HORhighlowmoderate
BLhighlowmoderate
SChighlowexcellent
SCCFlowhighexcellent

4. Conclusions

This study investigated the phenomena of stratum movements and stratum stress variation during the erosion cavity expansion phase induced by pipe bursts through laboratory-validated 2D numerical simulations, proposing disaster mitigation strategies based on different geotextile layout configurations. The main conclusions were as follows:
(1)
Stratum movements were affected by internal water pressures, burial depths, and geotextile layout configurations. Higher pressure accelerated the cavity propagation rate and increased the final erosion cavity area. Increased burial depth reduced peak ground heave but expanded the heave zone range linearly with the burial depths. Geotextile layout configurations can effectively restrict erosion cavity expansion and reduce stratum displacement (the peak ground heave was reduced by at least 15%). Specifically, SCCF excelled in decelerating the disaster occurrence (i.e., 39% reduction in the peak ground heave).
(2)
The stratum stress distribution depended on internal water pressures, burial depths, and geotextile layout configurations. An increase in internal water pressure led to stress amplification within a 2D range around the erosion cavity, with pronounced effects on peak stress values. Larger burial depths increased the overall stress level and generated larger stress reduction zones induced by the unloading effect (i.e., when d / D = 3.0, the range of the stress reduction zone was 8.0D). The presence of the geotextile can disperse the concentrated stress around the cavity to a larger zone.
(3)
Critical internal water pressure ( P c r ) governed the erosion cavity expansion behavior during pipe bursts. When the internal water pressure was less than P c r , erosion cavity development finally stabilized due to the resistance of the overlying stratum. Concerning the limit equilibrium theory, this study derived a predictive equation for P c r and validated its accuracy. Compared to the no geotextile case, SCCF can increase P c r by 25%, thus preventing ultimate erosion cavity expansion and achieving the purpose of disaster mitigation.
(4)
A comparative analysis of stratum displacement contour and geotextile deformation under different working conditions revealed that SCCF and SC presented better disaster mitigation effects in controlling stratum movements. SC maintained higher internal force levels, making it suitable for pipeline sections with low stratum deformation sensitivity. For the low-pressure pipelines in deformation-sensitive stratum, SCCF proved to be an optimal choice.
For pressurized pipeline burst failures, this study investigated stratum movements and stress distribution characteristics through validated numerical simulations and demonstrated the effectiveness of geotextile mitigation strategies. Future studies should be conducted to explore the stratum of diverse or layered strata to pipe bursts, as well as the disaster mitigation effects of geotextiles composed of various materials, with the dual objectives of enhancing the disaster resilience of pressurized pipelines and minimizing associated socioeconomic losses.

Author Contributions

Conceptualization, Z.H.; methodology, Z.S.; software, Z.H., H.C. and Z.W.; validation, Z.S. and X.L.; formal analysis, Z.H. and X.L.; investigation, Z.S.; resources, Y.T.; data curation, H.C. and Z.H.; writing—original draft preparation, Z.H.; writing—review and editing, H.C., Z.H. and Y.T.; visualization, H.C.; supervision, Y.T.; project administration, Y.T.; funding acquisition, Y.T. All authors have read and agreed to the published version of the manuscript.

Funding

This research received no external funding.

Data Availability Statement

The data presented in this study are available upon request from the corresponding author.

Conflicts of Interest

Author Zekun Su was employed by the company Shanghai Geotechnical Investigations & Design Institute Co., Ltd. The remaining authors declare that the research was conducted in the absence of any commercial or financial relationships that could be construed as a potential conflict of interest.

Abbreviation

List of Symbols
d Burial depth of the pipeline
DDiameter of the initial erosion cavity
P Internal water pressure
P c r Critical water pressure
HORThe geotextile arranged in a horizontal arrangement
BLThe geotextile arranged in a broken-line arrangement
SCThe geotextile arranged in a semi-circular arrangement
SCCFThe geotextile arranged in a semi-circular closely fitted arrangement

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Figure 1. Schematic diagram of disaster induced by pipe bursts.
Figure 1. Schematic diagram of disaster induced by pipe bursts.
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Figure 5. Comparison between the full-section experimental test and numerical simulation (The colors of the strata were to facilitate observation and they had no special meaning).
Figure 5. Comparison between the full-section experimental test and numerical simulation (The colors of the strata were to facilitate observation and they had no special meaning).
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Figure 6. Schematic diagrams of geotextile layout configurations: (a) horizontal (HOR); (b) broken-line (BL); (c) semi-circular (SC); (d) semi-circular closely fitted (SCCF).
Figure 6. Schematic diagrams of geotextile layout configurations: (a) horizontal (HOR); (b) broken-line (BL); (c) semi-circular (SC); (d) semi-circular closely fitted (SCCF).
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Figure 14. Development of stratum displacement contour under various working conditions (1 timestep =   5 × 10 5     s ).
Figure 14. Development of stratum displacement contour under various working conditions (1 timestep =   5 × 10 5     s ).
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Figure 15. Development of particle vertical velocities under various working conditions: (a) distance of 1.5D at angle of 0°; (b) distance of 1.5D at angle of 15° (1 timestep =   5 × 10 5   s ).
Figure 15. Development of particle vertical velocities under various working conditions: (a) distance of 1.5D at angle of 0°; (b) distance of 1.5D at angle of 15° (1 timestep =   5 × 10 5   s ).
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Table 1. Parameters adopted in numerical simulation.
Table 1. Parameters adopted in numerical simulation.
Parameters
SoilParticle density (kg/m3)2600
Friction coefficient0.50
Cohesion (kPa)0.00
Particle normal stiffness (N/m)1.00 × 108
Particle shear stiffness (N/m)5.00 × 107
Damping0.70
Normal critical damping ratio0.20
Porosity of granular material0.30
Wall normal stiffness (N/m)1.00 × 108
Wall friction0.50
GeotextileParticle normal stiffness (N/m)1.00 × 108
Particle shear stiffness (N/m)5.00 × 107
Normal critical damping ratio0.20
Parallel bond normal stiffness (N/m)3.60 × 108
Parallel bond shear stiffness (N/m)5.00 × 107
Parallel bond tensile strength (N/m)1.00 × 106
Table 2. Summary of numerical cases.
Table 2. Summary of numerical cases.
Case No.Burial Depth, d/DInternal Water Pressure,
P (MPa)
Geotextile Layout
Configuration
11.50.10/0.15/0.20/0.25/0.30No geotextile
22.00.10/0.15/0.20/0.25/0.30No geotextile
32.50.10/0.15/0.20/0.25/0.30No geotextile
43.00.10/0.15/0.20/0.25/0.30No geotextile
53.50.10/0.15/0.20/0.25/0.30No geotextile
62.00.20HOR
72.00.20BL
82.00.20SC
92.00.20SCCF
Note: D is the diameter of the initial erosion cavity (D = 1.00 m).
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MDPI and ACS Style

Hao, Z.; Chao, H.; Tan, Y.; Wang, Z.; Su, Z.; Li, X. Stratum Responses and Disaster Mitigation Strategies During Pressurized Pipe Bursts: Role of Geotextile Reinforcement. Buildings 2025, 15, 2696. https://doi.org/10.3390/buildings15152696

AMA Style

Hao Z, Chao H, Tan Y, Wang Z, Su Z, Li X. Stratum Responses and Disaster Mitigation Strategies During Pressurized Pipe Bursts: Role of Geotextile Reinforcement. Buildings. 2025; 15(15):2696. https://doi.org/10.3390/buildings15152696

Chicago/Turabian Style

Hao, Zhongjie, Hui Chao, Yong Tan, Ziye Wang, Zekun Su, and Xuecong Li. 2025. "Stratum Responses and Disaster Mitigation Strategies During Pressurized Pipe Bursts: Role of Geotextile Reinforcement" Buildings 15, no. 15: 2696. https://doi.org/10.3390/buildings15152696

APA Style

Hao, Z., Chao, H., Tan, Y., Wang, Z., Su, Z., & Li, X. (2025). Stratum Responses and Disaster Mitigation Strategies During Pressurized Pipe Bursts: Role of Geotextile Reinforcement. Buildings, 15(15), 2696. https://doi.org/10.3390/buildings15152696

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