Open AccessArticle

Centrifugal Test Study on the Vertical Uplift Capacity of Single-Cylinder Foundation in High-Sensitivity Marine Soil

Mingzhe Wei

¹,

Yanghui Ye

²,

Wei Zhao

^1,*,

Zehao Wang

¹,

Fuhao Ge

¹ and

Tingkai Nian

^1,*

State Key Laboratory of Coastal and Offshore Engineering, Dalian University of Technology, Dalian 116024, China

Economic and Technical Research Institute, State Grid Jiangxi Electric Power Co., Ltd., Nanchang 330096, China

Authors to whom correspondence should be addressed.

J. Mar. Sci. Eng. 2024, 12(12), 2152; https://doi.org/10.3390/jmse12122152

Submission received: 29 October 2024 / Revised: 17 November 2024 / Accepted: 20 November 2024 / Published: 25 November 2024

(This article belongs to the Special Issue Advances in Marine Engineering: Geological Environment and Hazards—3rd Edition)

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Abstract

Offshore wind power is a new type of clean energy with broad development prospects. Accurate analysis of the uplift capacity of offshore wind turbine foundations is a crucial prerequisite for ensuring the safe operation of wind turbines under complex hydrodynamic conditions. However, current research on the uplift capacity of suction caissons often neglects the high-sensitivity characteristics of marine soils. Therefore, this paper first employs the freeze–thaw cycling procedure to prepare high-sensitivity saturated clay. Subsequently, a single−tube foundation for wind turbines is constructed within a centrifuge through a penetration approach. Ten sets of centrifuge model tests with vertical cyclic pullout are conducted. Through comparative analysis, this study explores the pullout capacity and its variation patterns of suction caisson foundations in clay with different sensitivities under cyclic loading. This research indicates the following: (1) The preparation of high-sensitivity soil through the freeze−thaw procedure is reliable; (2) the uplift capacity of suction caissons in high−sensitivity soil rapidly decreases with increasing numbers of cyclic loads and then tends to stabilize. The cumulative displacement rate of suction caissons in high-sensitivity soil is fast, and the total number of pressure–pullout cycles required to reach non-cumulative displacement is significantly smaller than that in low-sensitivity soil; (3) the vertical cyclic loading times and stiffness evolution patterns of single-tube foundations, considering the influence of sensitivity, have been analyzed. It was found that the secant stiffness exhibits a logarithmic function relationship with both the number of cycles and sensitivity. The findings of this study provide assistance and support for the design of suction caissons in high-sensitivity soils.

Keywords:

offshore wind turbines; freeze–thaw; high-sensitivity soil; centrifuge model test; cyclic loading

1. Introduction

Wind, solar, and hydropower are currently the most advanced renewable energy sources, with wind energy projected to experience the fastest growth in production among renewables [1]. Since 2020, offshore wind energy has gradually become a key strategic industry within the global wind sector [2]. As the energy landscape shifts from land-based to oceanic sources, seabed stability assessment has emerged as a critical research area [3]. For example, Nian et al. have explored earthquake-induced seabed instability and examined how temperature affects the rheological properties of mudflows [4,5]. Common offshore wind turbine foundations include monopile, gravity-based, jacket, and floating foundations [6,7]. However, as wind power moves toward deeper waters, monopile foundations face challenges such as escalating costs and construction complexities [8]. In contrast, cylindrical foundations, which feature lower production costs, simpler design, and recyclability, are more adaptable to complex marine hydrodynamic conditions and varying seabed strata [9], leading to extensive research and application in recent years [10,11,12].

The application of cylindrical foundations dates back to 2002, when Denmark first installed a cylindrical foundation with a diameter of 12 m and a height of 6 m [13]. A key design consideration for these foundations is their response under cyclic loading. Kou et al. conducted indoor model tests to investigate the displacement, rotation angle, and rotation center variation characteristics of cylindrical foundations with suction caissons under different horizontal cyclic loads [14]. Lu utilized centrifuge model tests to investigate the deformation and bearing characteristics of cylindrical foundations in sandy soil under unidirectional cyclic loading [15]. Jiao conducted centrifuge model tests to study the development of pore pressure in clayey soil foundations under cyclic loading, revealing that the depth of soil liquefaction increases with the amplitude of cyclic loading. Moreover, based on pore pressure data, the author further examined the effects of cyclic loading on the bearing capacity of cylindrical foundations [16]. All in all, the bearing capacity and cyclic displacement of single−cylinder foundations have been examined mainly by scaled model tests [17,18,19], finite element simulations [20,21,22,23], and limiting analyses [24,25]. Some have examined the effects of foundation geometry, soil conditions, and load direction on the bearing capacity of cylinder foundations [26,27,28]. There are also some scholars who have studied the effects of load magnitude, load mean, load direction, and other factors on the cyclic displacement of the cylinder base [29]. The sensitivity of marine soils may be high due to rapid sedimentation on the seafloor, where the superporous water pressure within the sediments does not dissipate in time [30,31,32]. Although numerous studies have examined cylindrical foundations, few have addressed the high sensitivity of marine soils. There has been limited research into the effects of soil softening in high-sensitivity soils under cyclic loading, and design methods often fail to account for the influence of sensitivity on soil strength [33,34]. The settlement and sliding failure of the Phase II regulation structures at the Yangtze River Estuary in Shanghai was attributed to inadequate consideration of the high sensitivity of marine soil [35]. Therefore, it is crucial to account for the high sensitivity of marine clay in research and design processes.

Therefore, this paper proposes the use of a freeze–thaw procedure to prepare high-sensitivity saturated clay and obtain the undisturbed strength c_ui and remolded strength c_ud through vane shear tests to calculate the sensitivity S_t = c_ui/c_ud. Subsequently, in a centrifuge, the actuator is controlled for the purpose of installing a single-cylinder foundation for wind turbines through penetration. Ten centrifuge model tests of vertical cyclic pullout were conducted. Through comparative analysis, this study explores the pullout capacity and its variation patterns of suction caisson foundations in clay with different sensitivities under cyclic loading. Additionally, a logarithmic function relationship is proposed between the secant stiffness of hysteresis loops—associated with cyclic loading and cumulative displacement—and clay sensitivity, as well as the number of loading cycles. The findings offer valuable insights to support future experimental research and engineering design applications.

2. Experimental Methods

2.1. Preparation of High-Sensitivity Soil for Centrifuge Testing

The raw material, consisting of ball clay or kaolin, is manually mixed with an appropriate amount of water until it reaches a specific consistency. It is then transferred into a vacuum mixing tank for thorough blending, ensuring that no large agglomerates remain in the soil slurry. Simultaneously, a vacuum pump evacuates air, maximizing the saturation of the slurry [36]. The well-mixed slurry is then carefully poured, layer by layer, into a standard laboratory mold lined with filter paper, geotextile, and a pressure plate. The top surface of the slurry is also covered with filter paper, geotextile, and a pressure plate to facilitate normal consolidation. Consolidation pressure is applied using standard weights placed on top of the mold. To maintain constant saturation throughout the process, a water layer is kept on the surface, and a dial indicator records soil settlement data to monitor consolidation.

Once the soil achieves normal consolidation strength, following the consolidation process, the method by Nian et al. on the effects of temperature changes on soil strength [37] is referenced. Freeze–thaw cycling is applied to regenerate the soil structure and enhance its sensitivity. This approach is used to prepare the high−sensitivity marine saturated soil needed for the experiment. A sensitivity range of 4 to 8 is considered as high sensitivity [38].

By subjecting the clay to different temperature variations and multiple freeze–thaw cycles, electric cruciform shear tests, and making corrections to the data [39], the undisturbed strength and remolded strength of the specimens were obtained; then, the sensitivity was calculated based on the results. A comparative analysis of the clay’s sensitivity before and after freeze–thaw cycling was conducted, as shown in Table 1. The results indicate a significant increase in sensitivity post-freeze–thaw treatment. The soil samples, both pre- and post-freeze–thaw, were sectioned, sublimation-dried, metal-coated, and scanned using an electron microscope to observe the microstructural changes. The analysis revealed that, following the freeze–thaw process, the number of small pores in the soil decreased, while the number of large pores increased, as illustrated in Figure 1.

The large pores observed after the freeze–thaw process are formed due to the suction force exerted by the freezing front on the water in the unfrozen region. The temperature differential creates matric suction in the unfrozen zone, causing pore water to migrate from warmer areas to colder ones. As the water accumulates at the freezing front, it freezes and expands into ice crystals, ultimately leading to the coalescence of smaller pores into larger ones. This transformation results in increased soil sensitivity.

Using this freeze–thaw cycling method, using kaolin clay, and with a fixed consolidation strength of 10 kPa, high-sensitivity marine saturated soil for centrifuge testing was prepared by setting different consolidation times and freeze–thaw cycles. Following the preparation, vane shear tests were conducted to measure the strength and sensitivity of the soil. The control group consisted of low-sensitivity soil, while the experimental group featured high-sensitivity soil produced via the freeze–thaw method. The relevant soil parameters are detailed in Table 2.

2.2. Design of the Centrifuge Test Scheme

In practical engineering, it is common to place multiple suction caisson foundations within a specific area. To ensure the safe implementation of such foundations in real-world projects, particular attention is given to investigating their impact on soil properties under relevant test conditions. According to relevant research [40], the following assumptions are typically made:

(1): Wind, wave, and other environmental loads are uniformly distributed across all suction caisson foundations within the same marine area.
(2): The overall bending moment acting on the foundation is resisted by the combined uplift and compression capacities of each suction caisson.
(3): Regarding the foundation of a multi-bucket structure, the bending moment loads on each bucket are borne by the underlying soil. The magnitude of these loads depends on factors such as the footprint of the jacket structure, the structural weight, and environmental loads. When investigating the vertical extraction of suction buckets, the bending moment loads of the individual buckets are neglected.

The prototype monopole selected for the test has a diameter of 4 m, a length of 4 m, and a thickness of 0.04 m for both the top and sidewalls. In the centrifuge model test, an acceleration value of 100 g was used, which allowed for the calculation of the model monopole dimensions: a diameter of 40 mm, a length of 40 mm, and a thickness of 4 mm. The top of the monopole is equipped with a 5 mm diameter drainage valve and is connected to the centrifuge loading rod using a threaded rod with a length of 15 mm and a diameter of 8 mm. The centrifuge device is equipped with an actuator capable of radial movement up to 600 mm and circumferential movement up to 180°, as shown in Figure 2a. The purpose of compression, extraction, and cyclic loading is achieved by externally setting relevant force control parameters and controlling the radial movement of the actuator while the centrifuge is rotating at high speed. The position of the centrifuge test model box, suction cylinder, and sensors is shown in Figure 2b. The arrangement of the vane shear test and suction cylinder placement points is shown in Figure 2c. The penetration, monotonic, and cyclic loading processes are illustrated in the schematic diagram in Figure 2d.

Due to the randomness and unpredictability of horizontal cyclic loads, such as those caused by wind, waves, and currents, these loads are typically characterized by their amplitude and mean value. Firstly, the direction and symmetry characteristics of the load are determined by analyzing the relative magnitudes of the amplitude and mean value. Next, the load level is evaluated by comparing the ratio of the amplitude and mean value to the static bearing capacity. To account for the randomness of cyclic loads, they are often simplified into a series of load sequences with varying amplitudes in design. These sequences are then equivalently transformed, based on the principle of strain accumulation, into loads corresponding to the maximum amplitude within the sequence, along with an equivalent number of cycles [41]. Cyclic loads are classified into four basic types based on their direction and symmetry, with bidirectional symmetric cyclic loads being considered the most critical loading condition [42,43]. Therefore, bidirectional symmetric cyclic loading was selected for the centrifuge simulation test.

In the centrifuge model tests, to minimize boundary effects and avoid interactions between different points, the distance between the suction caisson and the edges of the model box was kept at least 1.8D, and the distance between suction caissons was ensured to be at least 2D (where D is the diameter of the caisson). Each soil box had four test locations, with the arrangement of suction caisson locations and vane shear test locations shown in Figure 2c. Out of the four test locations in each soil box, one was designated to determine the ultimate bearing capacity, while the other three were used for cyclic penetration tests.

2.3. Centrifuge Test Procedures

The centrifuge model test was conducted in the 450 g·t drum-type centrifuge at Dalian University of Technology, as shown in Figure 2a. The test acceleration used in this paper was 100 g, and the data acquisition frequency was 40–50 Hz [44,45]. The main test procedures were divided into the following 6 steps.

2.3.1. Preliminary Inspection and Equipment Installation

The suction caisson was installed, and all sensors and the data acquisition system were checked to ensure proper functioning. The actuator’s direction was adjusted to confirm that the suction caisson was perpendicular to the soil surface.

2.3.2. Suction Caisson Foundation Installation and Positioning

The centrifuge was gradually accelerated to 100 g in increments of 20 g, maintaining each step for at least one minute. Upon reaching 100 g, the actuator was controlled to penetrate the suction caisson radially at a speed of 10 mm/s (given the suction cylinder’s length of 40 mm, to ensure a more precise and controlled penetration to the top of the suction cylinder during the installation process, a speed of 10 mm/s is employed for a duration of 4 s) until the top of the foundation was level with the soil surface. During penetration, the valve at the top of the suction caisson was kept open to release air and water from the caisson. After penetration, the centrifuge was stopped.

2.3.3. Determination of Ultimate Bearing Capacity

After the initial penetration, the suction caisson’s valve was closed, and the surrounding soil was inspected for gaps. The centrifuge was then re-accelerated to 100 g in the same incremental manner. Once stabilized, the actuator continued penetrating the soil at 1 mm/s for an additional 10 mm (0.25 times the length of the suction caisson). The resistance at this depth was considered the vertical bearing capacity V₀.

2.3.4. Determination of Cyclic Frequency

The cyclic frequency was chosen based on the load application rate. Undrained conditions were more easily achieved when the cycle duration was less than 2 s. The cyclic test frequency generally ranged from 0.2 to 0.7 Hz, representing different loading rates. Given the limitations of the centrifuge’s loading precision, a frequency range between 0.4 and 0.8 was applied, with cycle durations between 1.25 and 2.5 s. Setting the test frequency between 0.4 and 0.5 Hz met both the undrained condition and equipment precision requirements.

2.3.5. Vertical Cyclic Loading Test

After determining the ultimate bearing capacity, the centrifuge was gradually slowed and stopped. The suction caisson was then positioned at the cyclic test location, and the centrifuge was accelerated uniformly back to 100 g. Once stabilized, the ultimate bearing capacity V₀, cyclic frequency (0.3–0.5 Hz), maximum and minimum cyclic loads, and the force-controlled loading frequency were input into the control built-in program for cyclic press–pull testing. The test was terminated once the preset cyclic termination conditions were met, and the process was repeated at other locations.

2.3.6. Termination Conditions for Cyclic Loading

There are two ways to determine the termination conditions for cyclic loading:

(1): The test could be terminated when the amplitude and average value of the vertical cyclic displacement of the suction caisson stabilized within a certain range, indicating no further displacement accumulation.
(2): If the sum of the amplitude and average displacement reached one-fourth of the caisson’s length, the suction caisson was considered to have been pulled out, and the test was terminated.

The maximum cyclic load was calculated as the sum of the average load and the amplitude, while the minimum load was the difference between the average and the amplitude. In each test, the load amplitude ratio V_C/V₀ remained constant as the number of cycles increased (where V₀ is the ultimate bearing capacity, V_C is the load amplitude, and V_a is the average load value). The ratio V_C/V₀ ranged from 0.37 to 0.64, while V_a/V₀ ranged from 0.01 to 0.05, considering the inherent errors in force-controlled cycling.

3. Experimental Results and Analysis

The entire centrifuge test primarily focuses on the response in terms of load and displacement. The force sensor is placed between the suction bucket and the actuator, serving to connect them. The laser displacement sensor is positioned below the force sensor and is also connected to the actuator. Throughout the test, force sensors are used to record the cyclic force conditions, and displacement sensors are used to record displacement data. Finally, a relationship diagram between cyclic load and displacement is compiled. By defining the secant stiffness of the hysteresis loop, which represents the relationship between cyclic load and displacement changes, and plotting its scatter plot, a functional relationship between the secant stiffness, the number of cycles, and the soil sensitivity is fitted.

3.1. Experimental Relationship Between Normalized Vertical Displacement and Number of Cycles

By defining one cylinder length as the unit length, the normalized vertical displacement is calculated as the ratio of the cylinder’s vertical displacement to the unit length L. During the test, the position changes in the suction caisson were recorded each time the set pressure and tension values were reached. Based on these recorded values, an experimental relationship between normalized vertical displacement and the number of cycles was plotted. In the diagram, tension is designated as positive and compression as negative, as illustrated in all parts of Figure 3.

As shown in Figure 3, the relationship depicted is between the normalized vertical displacement w/L and the number of cycles N. The cycling process stops when the displacement reaches one-quarter of the cylinder length, at which point the cylinder is considered to have been pulled out. The upward bearing capacity of the soil during cycling is generally lower than its downward bearing capacity. As a result, under the same vertical force, the upward and downward bearing capacities of the cylinder change due to soil softening and soil structural modifications during continuous cycling. This leads to the upward displacement being greater than the downward displacement, ultimately causing the extraction of the cylinder.

In high-sensitivity soils, where the pullout bearing capacity is lower than the compression bearing capacity, the outward displacement of the cylinder continues to accumulate with each cycle. This accumulation causes the cylinder to progressively move outward, and the compression position changes constantly, making it impossible to maintain a stable compression range. In contrast, normally consolidated low-sensitivity soils show a more stable compression position during cycling and may even experience further penetration in some cases. From the perspective of forces acting on the cylinder during pullout, the primary forces involved are end resistance and frictional resistance of the soil. Since the frictional resistance remains the same during both compression and pullout, the end resistance is closely tied to the soil’s strength. During cyclic loading, the soil undergoes disturbance and softening, leading to a reduction in strength. The sensitivity of the soil reflects the degree of structural integrity and the rate of strength decay after being disturbed. Therefore, under cyclic loading, high-sensitivity soils experience a rapid decrease in strength, and the soil at the cylinder location is unable to provide adequate bearing capacity. Consequently, the accumulation of upward and downward displacements is more pronounced in high-sensitivity soils compared to the control group. Furthermore, the rapid strength degradation in high-sensitivity soils after disturbance results in a pullout rate that far exceeds that of low-sensitivity soil. This is why the number of cycles for high-sensitivity soils is generally lower than that for low-sensitivity soils, and the softening rate is much faster. In summary, sensitivity is a critical factor in the response to cyclic loading, significantly influencing both displacement accumulation and bearing capacity.

3.2. Relationship Between Normalized Cyclic Loading and Normalized Displacement

By recording the position changes and vertical load magnitudes of the suction cylinder at each instance when the set pressure and tension values were reached during the experiments, a relationship between normalized vertical load and normalized vertical displacement was established (as shown in Figure 4).

Figure 4 shows the relationship between normalized cyclic loading and normalized vertical displacement for low-sensitivity soil, also known as the hysteresis loop. When the load amplitude is high, the maximum displacement of the cylinder reaches 0.25L within 100 cycles, as shown in Figure 4c,f. When the load amplitude is lower, the maximum displacement does not reach 0.25L, even after 400–500 cycles, as seen in Figure 4b,d. To clearly demonstrate the change in the hysteresis loop with the number of cycles, for groups with a larger number of experimental cycles, typical hysteresis loops at different loading stages are provided in the figures. From Figure 4a,b, it can be observed that the upward displacement of the single cylinder accumulates significantly with increasing cycle number, while the downward displacement decreases slightly with increasing cycle number. In contrast to Figure 4c, the downward displacement of the single cylinder also accumulates, but the accumulation is less than that of the upward displacement. This is because the average load V_a/V₀ in Figure 4c, which is 0.05, is slightly greater than the average load V_a/V₀ in Figure 4a,b, which is 0.01. The slightly higher pressure compared to tension exacerbates the accumulation of downward displacement.

Figure 5 illustrates the relationship between normalized cyclic loading and normalized vertical displacement for high-sensitivity soil. The slope of the line connecting the highest and lowest points of the Nth hysteresis loop is defined as the secant stiffness K_N of the Nth hysteresis loop for the single cylinder, as shown in Figure 6. This secant stiffness can be obtained by calculating the ratio of the change in vertical load to the change in vertical displacement within that cycle:

K_{N} = \frac{V_{\max} - V_{\min}}{W_{\max} - W_{\min}}

(1)

where V_max and V_min are maximum and minimum values of vertical loads, respectively; W_max and W_min are maximum and minimum values of vertical displacement, respectively. By further segmenting the hysteresis loops representing the cyclic loading–displacement relationship for both low-sensitivity soil and high−sensitivity soil, their secant stiffnesses are calculated and plotted as scatter plots, as shown in Figure 7.

Based on Figure 7, it can be observed that the hysteresis loop stiffness K_N decreases as V_C/V₀ increases, and K_N also decreases as the number of cycles N increases [46]. The relationship between the cyclic secant stiffness K_N and the number of cycles N can be fitted as follows:

K_{N} = K_{1} + A_{k} \ln N

(2)

where K₁ and A_k are the coefficient variables. Next, by fitting the load magnitude and initial secant stiffness for both low-sensitivity soil and high-sensitivity soil, the following parameters can be obtained:

K_{1} = (78.06 {lnS}_{t} - 381.91) ζ_{b} + 213.37 {S_{t}}^{- 0.206}

(3)

where

ζ_{b}

is load magnitude,

S_{t}

is level of sensitivity, the R-squared (R²) value for the data fit with low-sensitivity soil is 94.9%, and the R-squared (R²) value for the data fit with high-sensitivity soil is 98.3%. Furthermore, the parameter

A_{k}

can be estimated and fitted with the load magnitude to obtain the following relationship:

A_{k} = (54.33 {S_{t}}^{- 0.688}) ζ_{b} + 6.045 {lnS}_{t} - 22.212

(4)

The R-squared (R²) value for the data fit with low-sensitivity soil is 92.3%, and the R-squared (R²) value for the data fit with high-sensitivity soil is 83.7%. Substituting Equations (3) and (4) into Equation (2), the relationship between the secant stiffness

K_{N}

, the number of cycles N, and the sensitivity

S_{t}

can be derived; ultimately, we arrive at Equation (5).

K_{N} = [(78.06 {lnS}_{t} - 381.91) ζ_{b} + 213.37 {S_{t}}^{- 0.206}] + [(54.33 {S_{t}}^{- 0.688}) ζ_{b} + 6.045 {lnS}_{t} - 22.212] \times \ln N

(5)

Finally, the data were substituted into the formula to obtain the dashed straight line in Figure 7. Comparing it with the original data scatter plot, it can be seen that the error is not significant.

4. Conclusions

This study utilized the freeze–thaw method to prepare high-sensitivity soil and conducted centrifuge model tests on vertically symmetrical cyclic loading of offshore wind turbine suction caissons. The high-sensitivity soil served as the experimental group, while low-sensitivity soil was used as the control group. The primary goal was to compare and investigate the cyclic load–displacement response under vertically symmetrical loading, focusing on the effect of soil sensitivity. The following conclusions were drawn from the experiments:

Under symmetric cyclic loading, the uplift bearing capacity of the suction caisson was found to be lower than its compressive bearing capacity. As the number of cycles increased, the caisson continued to be pulled upward. The average load value was influenced by both the cyclic amplitude and soil sensitivity. Higher cyclic amplitude and greater soil sensitivity resulted in larger average load values. Soils with high sensitivity exhibited stronger structural properties, contributing to higher average load values.
During the experiments, high-sensitivity soils showed a rapid reduction in strength under the disturbance of cyclic loading, causing a swift decrease in both uplift and compressive bearing capacities. As a result, the suction caisson experienced greater displacement to reach the set load value. In contrast, the low-sensitivity soil group exhibited minimal strength degradation, resulting in a higher number of cycles and a longer displacement accumulation period. The compressive position remained stable within a certain range, and a near-steady state was observed under low cyclic amplitudes.
Secant stiffness, defined as the secant slope of the hysteresis loop relating normalized cyclic load to normalized displacement, was investigated. The vertical cumulative displacement and stiffness evolution of the suction caisson were examined, taking into account the influence of soil sensitivity. A logarithmic function relationship between secant stiffness, number of cycles, and sensitivity was established.
The centrifuge test was accelerated to a maximum of 100 g, simulating long-duration cyclic loading according to the time-scaling law. However, in real-world conditions, the intervals between uplift and compression cycles are shorter, and various factors, such as temperature fluctuations and reconsolidation, may affect the response. These factors were not accounted for in the experiment due to inherent limitations and errors in model testing. Due to limitations such as the environmental conditions in the laboratory and errors in experimental equipment, using the same parameters may lead to some variations in results, resulting in non-reproducibility of the experiments. We kindly seek the readers’ understanding for this.

Author Contributions

Conceptualization, W.Z. and T.N.; methodology, M.W. and F.G.; experimental, M.W., F.G., and Y.Y.; formal analysis, Z.W.; investigation, W.Z. and T.N.; resources, W.Z.; data curation, M.W. and Y.Y.; writing—original draft preparation, M.W. and Z.W.; writing—review and editing, Z.W., T.N., and W.Z.; visualization, Z.W.; supervision, T.N.; project administration, T.N.; funding acquisition, T.N. All authors have read and agreed to the published version of the manuscript.

Funding

This research was funded by National Natural Science Foundation of China, grant numbers 52079020 and 42377185, and the open research fund program of Zhoushan Field Scientific Observation and Research Station for Marine Geo-hazards.

Institutional Review Board Statement

Not applicable.

Informed Consent Statement

Not applicable.

Data Availability Statement

Data can be obtained from the first author upon request.

Acknowledgments

The authors gratefully acknowledge the support from the funding listed above. We also appreciate the anonymous reviewers who gave comments to revise the paper.

Conflicts of Interest

Author Yanghui Ye was employed by the company State Grid Jiangxi Electric Power 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.

References

Cozzi, M.E.L.; Goodson, T. Global Energy Review; IEA: Paris, France, 2021; Available online: https://www.iea.org/reports/global-energy-review-2021 (accessed on 1 April 2021).
Wang, S.; Liu, J.; Liu, Z.; Li, N. Analysis of Current Situation and Future Development of Offshore Wind Power Industry. South Energy Constr. 2023, 10, 103–112. [Google Scholar]
Guo, X.; Fan, N.; Zheng, D.; Fu, C.; Wu, H.; Zhang, Y.; Song, X.; Nian, T. Predicting impact forces on pipelines from deep-sea fluidized slides: A comprehensive review of key factors. Int. J. Min. Sci. Technol. 2024, 34, 211–225. [Google Scholar] [CrossRef]
Nian, T.; Guo, X.; Zheng, D.; Xiu, Z.; Jiang, Z. Susceptibility assessment of regional submarine landslides triggered by seismic actions. Appl. Ocean Res. 2019, 93, 101964. [Google Scholar] [CrossRef]
Nian, T.; Guo, X.; Fan, N.; Jiao, H.; Li, D. Impact forces of submarine landslides on suspended pipelines considering the low-temperature environment. Appl. Ocean Res. 2018, 81, 116–125. [Google Scholar] [CrossRef]
Byrne, B.W.; Houlsby, G.T. Foundations for offshore wind turbines. Philos. Trans. R. Soc. A 2003, 361, 2909–2930. [Google Scholar] [CrossRef]
Liu, R.; Zhou, L.; Lian, J.; Ding, H. Behavior of monopile foundations for offshore wind farms in sand. J. Waterw. Port. C ASCE 2016, 142, 04015010. [Google Scholar] [CrossRef]
Bae, K.T.; Choo, Y.W.; Youn, H.J.; Kim, J.Y.; Kwon, O.S. Geotechnical perspective on offshore wind plan, strategy, projects and research in Korea. In Proceeding of the TC 209 Workshop, Seoul, Republic of Korea, 20 September 2017; Volume 19, pp. 13–20. [Google Scholar]
Wang, K.; Zhu, Y.; Yu, M.; Peng, M. Study on the Compressive Bearing Capacity Mode of Suction Caisson Foundations for Offshore Wind Turbines. J. Eng. Geol. 2022, 30, 1753–1761. (In Chinese) [Google Scholar]
Bienen, B.; Gaudin, C.; Cassidy, M.J.; Rausch, L.; Purwana, O.A.; Krisdani, H. Numerical modelling of a hybrid skirted foundation under combined loading. Comput. Geotech. 2012, 45, 127–139. [Google Scholar] [CrossRef]
Kim, D.J.; Choo, Y.W.; Kim, J.H.; Kim, S.; Kim, D.S. Investigation of monotonic and cyclic behavior of tripod suction bucket foundations for offshore wind towers using centrifuge modeling. J. Geotech. Geoenviron. 2014, 140, 04014008. [Google Scholar] [CrossRef]
Wang, X.; Yang, X.; Zeng, X. Centrifuge modeling of lateral bearing behavior of offshore wind turbine with suction bucket foundation in sand. Ocean Eng. 2017, 139, 140–151. [Google Scholar] [CrossRef]
Zhang, Y.; Wang, C.; Li, D.; Qi, Y. Numerical Simulation of Composite Bearing Characteristics of Skirted Suction Foundations in Sandy Soil. Sci. Technol. Eng. 2021, 21, 1515–1521. (In Chinese) [Google Scholar]
Kou, H.; Zhou, N.; Yang, D.; Feng, J.; Tian, H. Experimental Study on Horizontal Bearing Characteristics of Suction Bucket Foundations for Offshore Wind Turbines in Sandy Soil Foundations. J. Eng. Geol. 2021, 29, 1752–1758. (In Chinese) [Google Scholar]
Lu, X.; Wang, Y.; Zhang, J.; Sun, G.; Shi, Z. Centrifuge Tests on Deformation of Bucket Foundations under Horizontal Dynamic Loads. Chin. J. Geotech. Eng. 2005, 27, 789–791. (In Chinese) [Google Scholar]
Jiao, B.; Shi, Z.; Lu, X.; Zhang, J. Centrifuge Experimental Study on the Horizontal Dynamic Response of Bucket Foundations in Silty Soil. Eng. Mech. 2010, 27, 131–141. (In Chinese) [Google Scholar]
Choo, Y.W.; Kim, D.; Park, J.H.; Kwak, K.; Kim, J.H.; Kim, D.S. Lateral response of large-diameter monopiles for offshore wind turbines from centrifuge model tests. Geotech. Test. J. 2014, 37, 107–120. [Google Scholar] [CrossRef]
Clukey, E.C.; Aubeny, C.P.; Murff, J.D. Comparison of analytical and centrifuge model tests for suction caissons subjected to combined loads. Offshore Mech. Arct. Eng. 2004, 126, 364–367. [Google Scholar] [CrossRef]
Ding, H.; Liu, Y.; Zhang, P.; Le, C. Model tests on the bearing capacity of wide-shallow composite bucket foundations for offshore wind turbines in clay. Ocean Eng. 2015, 103, 114–122. [Google Scholar] [CrossRef]
Hu, Y.; Randolph, M.F. Bearing capacity of caisson foundations on normally consolidated clay. Soils Found. 2002, 42, 71–77. [Google Scholar] [CrossRef]
Bransby, M.F.; Yun, G.J. The undrained capacity of skirted strip foundations under combined loading. Géotechnique 2009, 59, 115–125. [Google Scholar] [CrossRef]
Lee, S.H.; Vicent, S.; Kim, S.R. An experimental investigation of the cyclic response of bucket foundations in soft clay under one-way cyclic horizontal loads. Appl. Ocean Res. 2018, 71, 59–68. [Google Scholar]
Skau, K.S.; Chen, Y.; Jostad, H.P. A numerical study of capacity and stiffness of circular skirted foundations in clay subjected to combined static and cyclic general loading. Géotechnique 2018, 68, 205–220. [Google Scholar] [CrossRef]
Aubeny, C.; Han, S.W.; Murff, J.D. Suction caisson capacity in anisotropic, purely cohesive soil. Int. J. Geomech. 2003, 3, 225–235. [Google Scholar] [CrossRef]
Aubeny, C.; Murff, J.D. Simplified limit solutions for the capacity of suction anchors under undrained conditions. Ocean Eng. 2005, 32, 864–877. [Google Scholar] [CrossRef]
Kim, S.R. Evaluation of combined horizontal-moment bearing capacities of tripod bucket foundations in undrained clay. Ocean Eng. 2014, 85, 100–109. [Google Scholar]
Bagheri, P.; Kim, J.M. Evaluation of cyclic and monotonic loading behavior of bucket foundations used for offshore wind turbines. Appl. Ocean Res. 2019, 91, 101865. [Google Scholar] [CrossRef]
Barari, A.; Glitrup, K.; Christiansen, L.R.; Ibsen, L.B.; Choo, Y.W. Tripod suction caisson foundations for offshore wind energy and their monotonic and cyclic responses in silty sand: Numerical predictions for centrifuge model tests. Soil Dyn. Earthq. Eng. 2021, 149, 106813. [Google Scholar] [CrossRef]
Jeong, Y.H.; Kim, J.H.; Ko, K.W.; Park, H.J. Simplified estimation of rotational stiffness of tripod foundation for offshore wind turbine under cyclic loadings. Appl. Ocean Res. 2021, 112, 102697. [Google Scholar] [CrossRef]
Wang, Z.; Zheng, D.; Guo, X.; Gu, Z.; Yan, C.; Nian, T. A review of submarine landslides associated with rapid sedimentation. Mar. Georesour. Geotechnol. 2024, 1–25. [Google Scholar] [CrossRef]
Jiao, H.; Guo, X.; Fan, N.; Wu, H.; Nian, T. An undrained dynamic strain-pore pressure model for deep-water soft clays from the South China Sea. Front. Mar. Sci. 2024, 11, 1377474. [Google Scholar] [CrossRef]
Wang, Z.; Zheng, D.; Gu, Z.; Guo, X.; Nian, T. A Methodology to Evaluate the Real-Time Stability of Submarine Slopes under Rapid Sedimentation. J. Mar. Sci. Eng. 2024, 12, 823. [Google Scholar] [CrossRef]
Khemakhem, M.; Chenaf, N.; Garnier, J.; Favraud, C.; Gaudicheau, P. Development of degradation laws for describing the cyclic lateral response of piles in clay. In Proceedings of the SUT Offshore Site Investigation and Geotechnics, London, UK, 12–14 September 2012; p. SUT–OSIG-12-27. [Google Scholar]
Liao, W.; Zhang, J.; Wu, J.; Yan, K. Response of flexible monopile in marine clay under cyclic lateral load. Ocean Eng. 2018, 147, 89–106. [Google Scholar] [CrossRef]
Fan, Q.; Li, N. Causes and Countermeasures of Local Damage to the Northern Guide Bank of the Yangtze Estuary Phase II Project. China Harbour and Waterfront Engineering: Beijing, China, 2004. (In Chinese) [Google Scholar]
Anagnostopoulos, C.; Chrysanidis, T.; Anagnostopoulou, M. Experimental data of cement grouting in coarse soils with different superplasticisers. Data Brief 2020, 30, 105612. [Google Scholar] [CrossRef] [PubMed]
Nian, T.; Gu, Z.; Guo, X.; Zhang, H.; Jia, Y. Effect of temperature rise on the mechanical behaviour of deep-sea clay surrounding oil and gas pipelines. Ocean. Eng. 2024, 302, 117533. [Google Scholar] [CrossRef]
JGJ 83–2011; Code for Geotechnical Investigation in Soft Soil Areas. Architecture and Building Press: Beijing, China, 2011. (In Chinese)
Morris, P.H.; Williams, D.J. A new model of vane shear strength testing in soils. Géotechnique 1994, 44, 771–774. [Google Scholar] [CrossRef]
Carbon Trust. Suction Installed Caisson Foundations for Offshore Wind; Design Guidelines: London, UK, 2019. [Google Scholar]
Andersen, K.H. Cyclic soil parameters for offshore foundation design. Front. Offshore Geotech. III 2015, 5, 5–82. [Google Scholar]
Kou, H.; Fang, W.; Zhou, N.; Huang, J.; Zhang, X. Dynamic response of single-bucket foundation in clay under vertical variable amplitude cyclic loadings. Ocean Eng. 2023, 273, 113973. [Google Scholar] [CrossRef]
Randolph, M.; Gourvenec, S. Offshore Geotechnical Engineering; CRC Press: Boca Raton, FL, USA, 2017. [Google Scholar]
Guo, X.; Nian, T.; Zhao, W.; Gu, Z.; Liu, C.; Liu, X.; Jia, Y. Centrifuge experiment on the penetration test for evaluating undrained strength of deep-sea surface soils. Int. J. Min. Sci. Technol. 2022, 32, 363–373. [Google Scholar] [CrossRef]
Wang, L.; Wang, Z.; Li, S.; Wang, Z.; Zhu, H. Study on the settlement behavior of underwater protective caissons based on centrifuge model tests. Waterw. Harb. 2024, 45, 448–454. (In Chinese) [Google Scholar]
Leblanc, C.; Byrne, B.W.; Houlsby, G.T. Response of stiff piles to random two-way lateral loading. Géotechnique 2010, 60, 715–721. [Google Scholar] [CrossRef]

Figure 1. Microscopic images of kaolin and ball clay before and after freeze–thaw cycles: (a) microscopic image of ball clay before freeze–thaw; (b) microscopic image of kaolin before freeze–thaw; (c) microscopic image of ball clay after freeze–thaw; (d) microscopic image of kaolin after freeze–thaw.

Figure 2. Centrifuge test: (a) centrifugal model test device; (b) photo of test model box and suction cylinder; (c) point layout drawing; (d) schematic diagram of penetration, monotonic, and cyclic loading.

Figure 3. The experimental relationship between the normalized vertical displacement w/L and the cycle number N: (a) normalized vertical displacements versus number of cycles for low-sensitivity soil at V_C /V₀ = 0.39, V_a/V₀ = 0.01; V_C /V₀ = 0.425, V_a/V₀ = 0.03; V_C /V₀ = 0.45 and V_a/V₀ = 0.05 condition; (b) normalized vertical displacements versus number of cycles for low-sensitivity soil at V_C/V₀ = 0.35, V_a/V₀ = 0.01; V_C/V₀ = 0.325, V_a/V₀ = 0.02; V_C/V₀ = 0.3 and V_a/V₀ = 0.01 condition; (c) normalized vertical displacements versus number of cycles for highly sensitive soils at V_C/V₀ = 0.4, V_a/V₀ = 0.01; V_C/V₀ = 0.425, V_a/V₀ = 0.01 conditions; (d) normalized vertical displacements versus number of cycles for highly sensitive soils at V_C/V₀ = 0.3, V_a/V₀ = 0.01; V_C/V₀ = 0.32, V_a/V₀ = 0.01 conditions.

Figure 4. Normalized cyclic load and normalized displacement relationship diagram of low-sensitivity soil: (a) normalized cyclic load versus normalized vertical displacement for V_C/V₀ = 0.3 and V_a/V₀ = 0.01; (b) normalized cyclic load versus normalized vertical displacement for V_C/V₀ = 0.35 and V_a/V₀ = 0.02; (c) normalized cyclic load versus normalized vertical displacement for V_C/V₀ = 0.45 and V_a/V₀ = 0.05; (d) normalized cyclic load versus normalized vertical displacement for V_C/V₀ = 0.325 and V_a/V₀ = 0.01; (e) normalized cyclic load versus normalized vertical displacement for V_C/V₀ = 0.39 and V_a/V₀ = 0.01; (f) normalized cyclic load versus normalized vertical displacement for V_C/V₀ = 0.425 and V_a/V₀ = 0.03.

Figure 5. Normalized cyclic load and normalized displacement relationship diagram of high-sensitivity soil: (a) normalized cyclic load versus normalized vertical displacement for V_C/V₀ = 0.3; V_a/V₀ = 0.01; (b) normalized cyclic load versus normalized vertical displacement for V_C/V₀ = 0.32; V_a/V₀ = 0.01; (c) normalized cyclic load versus normalized vertical displacement for V_C/V₀ = 0.4; V_a/V₀ = 0.01; (d) normalized cyclic load versus normalized vertical displacement for V_C/V₀ = 0.425; V_a/V₀ = 0.01.

Figure 6. Definition of secant stiffness.

Figure 7. Scatter plot of secant stiffness: (a) scatter plot of cut-line stiffness of low-sensitivity soil; (b) scatter plot of stiffness of high-sensitivity earth cut line.

Table 1. Table of Sensitivity Changes in Clay Before and After Freeze–Thaw Cycles.

Freezing Temperature	Number of Cycle	S_t (Before Freezing and Thawing)	S_t (After Freezing and Thawing)	Growth Rate of Sensitivity Before and After Freeze–Thaw
−10 °C	1	1.86	5.44	+192.5%
−20 °C	1	1.48	4.97	+235.8%
−10 °C	5	1.94	7.98	+311.3%
−20 °C	5	2.07	8.16	+294.2%

Table 2. Soil Strength Parameter Table.

Category	c_ui (kPa)	c_ud (kPa)	S_t
High-sensitivity soil A₁	8.08	1.11	7.31
High-sensitivity soil B₁	7.68	1.14	6.73
High-sensitivity soil C₁	8.45	1.24	6.81
High-sensitivity soil D₁	8.83	1.26	7.01
Low-sensitivity soil A₂	9.01	4.29	2.10
Low-sensitivity soil B₂	7.69	4.07	1.89
Low-sensitivity soil C₂	9.18	3.96	2.32
Low-sensitivity soil D₂	8.46	3.72	2.27

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MDPI and ACS Style

Wei, M.; Ye, Y.; Zhao, W.; Wang, Z.; Ge, F.; Nian, T. Centrifugal Test Study on the Vertical Uplift Capacity of Single-Cylinder Foundation in High-Sensitivity Marine Soil. J. Mar. Sci. Eng. 2024, 12, 2152. https://doi.org/10.3390/jmse12122152

AMA Style

Wei M, Ye Y, Zhao W, Wang Z, Ge F, Nian T. Centrifugal Test Study on the Vertical Uplift Capacity of Single-Cylinder Foundation in High-Sensitivity Marine Soil. Journal of Marine Science and Engineering. 2024; 12(12):2152. https://doi.org/10.3390/jmse12122152

Chicago/Turabian Style

Wei, Mingzhe, Yanghui Ye, Wei Zhao, Zehao Wang, Fuhao Ge, and Tingkai Nian. 2024. "Centrifugal Test Study on the Vertical Uplift Capacity of Single-Cylinder Foundation in High-Sensitivity Marine Soil" Journal of Marine Science and Engineering 12, no. 12: 2152. https://doi.org/10.3390/jmse12122152

APA Style

Wei, M., Ye, Y., Zhao, W., Wang, Z., Ge, F., & Nian, T. (2024). Centrifugal Test Study on the Vertical Uplift Capacity of Single-Cylinder Foundation in High-Sensitivity Marine Soil. Journal of Marine Science and Engineering, 12(12), 2152. https://doi.org/10.3390/jmse12122152

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