Bearing Capacity of Offshore Wind Power Suction Bucket with Supports Under Extreme Wind and Waves

Abstract

In this study, we analyzed an offshore wind turbine suction bucket foundation with supports, referred to as the supported suction bucket foundation. A numerical model of a typical 3MW wind turbine suction bucket foundation was established, and the accuracy of the numerical modeling method was validated. The daily wind and wave extreme data for 2022 monitored by a station in the eastern China Sea were converted into loads and applied to the conventional and supported suction bucket foundations. The bearing capacity of the two bucket foundations was, thus, compared and analyzed. The results show that the supported suction bucket foundation reduces the foundation displacement and amplitude. Also, the support structure effectively reduces the soil displacement around the bucket. The maximum displacement of the outer soil of the bucket decreases by about 89.8%, while that of the inner soil of the bucket decreases by about 70.7%. Increasing the supports can reduce the separation of the top lid from the inner soil below it. Further, the supported suction bucket foundation can effectively reduce the plastic strain of the soil around the bucket and reduce the plastic strain difference between the inner and outer soil of the bucket.

1. Introduction

With growing concerns and interest in renewable energy and sustainability, wind power generation is increasing exponentially. Europe is the first region in the world to put the concept of offshore wind power generation to practice [1]. Denmark is the first country in the world to master offshore wind power technology, followed by other countries worldwide in exploring technological advancement and applications of offshore wind power generation [2]. Unlike the onshore environment, the working environment of the offshore wind turbine is more severe. The wind turbine foundation has to bear the self-weight of the superstructure, as well as the wind, waves, currents, and other environmental loads. Thus, the bearing performance of the wind turbine foundation has very high requirements.

The suction bucket foundation is one of the foundation forms for offshore wind energy. Researchers have extensively studied the characteristics of offshore wind power conventional suction bucket foundation bearings. Chen et al. [3] carried out a suction bucket model test under wave cyclic loading and concluded that when the wave period is stable, the sandy soil at the front and back sides of the foundation is suspected to liquefy. Wang et al. [4] conducted numerical simulations based on centrifugal tests to investigate the tensile and compressive properties of suction bucket foundations. Zhong et al. [5] proposed a reinforced bucket foundation and compared its bearing performance with the conventional bucket foundation by finite element simulation. Considering 11 years of wind and wave monitoring data, Qin et al. [6] investigated the bearing characteristics of shallow and deep suction bucket foundations under silty clay. Chen et al. [7] used numerical simulation to study the effects of the aspect ratio and soil conditions on suction bucket foundations. Wang et al. [8] investigated the response of the suction bucket foundation under horizontal dynamic loading in fine sand and concluded that the sandy soil around the suction bucket exhibits softening or even liquefaction characteristics. Gao et al. [9] investigated the response of suction bucket foundations under seismic and wind loads. Ma et al. [10] studied the horizontal response of a suction bucket foundation subjected to cyclic loading. They confirmed that the soil deformation around the bucket is proportional to the number of loading actions. In contrast, the stiffness is inversely proportional to the number of load actions, and the center of rotation of the bucket is upwardly shifted with the loading characteristics. Alp S Y et al. [11] studied the performance of the suction bucket foundation under cyclic axial pressure load based on the finite element model. Pouyan et al. [12] found that the displacement of the suction bucket foundation under horizontal loading leads to dense soil, and the displacement depends on the foundation shape and soil properties. Luo et al. [13] demonstrated that the periphery of the bucket wall and the lower part of the bucket end are the primary areas that develop an equivalent plastic strain distribution when the suction bucket foundation is subjected to inclined loading. Yang et al. [14] introduced Parkfield seismic waves and found that the foundation developed large deformation displacement after generating oscillations; the bearing characteristics of conventional suction bucket foundations in the relevant soil layers were also obtained. In harsh marine environments, recent research has focused on optimizing the structural configuration of suction bucket foundations to enhance their stability and load-bearing capacity. Zhang et al. [15] investigated the infiltration resistance of the suction bucket foundation with bulkheads during installation in sandy soil. Liang et al. [16] experimentally investigated the pullout bearing capacity of the suction bucket foundations with scaled walls and found that the bearing capacity is inversely proportional to the loading angle. Zhao et al. [17] investigated the force characteristics of umbrella suction bucket foundations under scouring action and suggested that the influence of scouring was concentrated in the area behind the anchor branch and bucket foundation. Li et al. [18] carried out model tests on skirt suction buckets and found that the pullout bearing capacity of skirt suction buckets increased with the length and width of the skirt structure after being subjected to inclined loads. Moreover, the bearing capacity and suction force in the bucket were found to be inversely and positively proportional to the angle of load application and the horizontal angle of clamping, respectively. Luo et al. [19] analyzed the load-bearing characteristics of conventional and new suction buckets using small-scale laboratory test models.

The aforementioned research studies on the load-carrying performance after improving the suction bucket base structure were mainly carried out on small-scale models in the laboratory. Findings on large-scale modeling for practical field applications and studies considering variations in the load-bearing properties of bucket foundations after actual extreme wind and wave action are sparse. Therefore, further integration of prototypes commonly applied in the field, field-monitored wind and wave data, and the load-bearing properties of different improved suction bucket structures must be carried out in depth. This study introduced a new type of supported suction bucket foundation, designed as the actual size of a 3 MW wind turbine, thus overcoming the limitations associated with small-scale experiments. By utilizing extreme wind and wave data collected in a Chinese sea area in 2022, the research realistically simulated the long-term cumulative effects of complex marine environmental loads. The study compared the bearing performances of both supported and conventional suction bucket foundations under these loads, providing valuable insights for the design and optimization of offshore wind turbine foundations in challenging marine environments.

2. Numerical Simulation

2.1. Model Dimensions and Parameters

The design of the supported suction bucket foundation is illustrated in Figure 1. The fixed transverse gears are actuated by machine manipulation, which then rotates the vertical gears. This rotation drives the support structure to align with a predetermined thread configuration until the support extends to a specified length. The purpose of this paper is to investigate how the addition of a supporting structure affects the load-carrying performance of the suction foundation. Therefore, the structural form of the new suction bucket foundation is simplified in the numerical model by adding four lateral supports in the horizontal direction to the conventional suction bucket foundation and setting these supports as a regular cylinder. To analyze the load-carrying performances of the two buckets, the model was established using the foundation dimensions of a classical 3 MW wind turbine. In this model, the foundation has a diameter of D = 18 m, a height of H = 12 m, a bucket wall thickness of d1 = 0.004 D, and a top cover thickness of d2 = 0.01 D. The height of the wind turbine tower is 90 m. Figure 2 illustrates the grid division for the conventional suction bucket, the supported suction bucket, and the combined suction bucket foundation and soil model, as well as the geometric relationships involved in the modeling process. The geometric relationships are the same for both types of bucket foundations. The soil and suction bucket models were simulated using integral three-dimensional solid elements. The soil was modeled as a cylinder using the Mohr–Coulomb constitutive model. To reduce potential boundary effects on the calculation results, the diameter of the soil was set to be ten times that of the bucket foundation, and its height was five times the height of the bucket. Both buckets were modeled as linearly elastic materials. The bucket bodies consist of linear elastic material, and the grid division for the buckets uses eight-node linear hexahedral elements (C3D8R), while the soil grid division employs eight-node linear hexahedral hybrid elements (C3D8RH). The grid density is the same for both types of bucket foundations. The tangential surface of the bucket and soil was a “penalized” contact with a coefficient of friction of 0.35, while the normal surface was a “hard” contact, allowing the bucket and soil to separate. The grid densities of the two bucket bases were convergent. The support added to the supported suction bucket was 2 m in diameter and 8 m in length and was added at 0.7 times the depth of the bucket. The parameter settings for the suction bucket foundation, support structure, and soil are given in

Table 1. Parameter settings of suction bucket foundation, supported structure, and soil.

Materials	Elastic Modulus/(Pa)	Poisson Ratio	Angle of Friction/(°)	Unit Weight γ/(kN/m3)
Suction bucket foundation	2.1 × 1011	0.3	-	78.5
Structural support	2.1 × 1011	0.3	-	78.5
Soil	6 × 106	0.495	18	6

.

2.2. Model Validation

To verify the reliability of the above modeling method, a suction bucket foundation model with the same dimensions, parameters, contact, and other conditions as in the literature [20] was established. A horizontal load was applied to it as specified in the literature [20], horizontal displacements under the loads were obtained, and the resulting load–displacement curves were compared with the results in the literature [20]. The comparison results are shown in Figure 3. The displacement response of the suction bucket grows slowly until the load reaches 2500 kN, and when the load exceeds the modified critical value, the displacement development is characterized by a significant acceleration. The difference in the load–displacement curves between this paper and the literature [20] is about 4%. To analyze the reason for this, this paper follows the principle of “near dense far sparse” when dividing the grid cells, i.e., the closer the contact surface is between the bucket and the soil, the more densely the grid is divided, and in the direction away from the contact surface between the bucket and the soil, the regional grid is gradually changed from dense to sparse. Therefore, the specific meshing method is not given in the literature [20], and there are small differences in the meshing of the model, resulting in a small error within the acceptable range. This indicates the rationality and reliability of the modeling method used in this paper.

Xie et al. [21] conducted experimental research on the bearing capacity characteristics of suction bucket foundations under cyclic loading in sandy soil. To further validate the reliability of the modeling method, a conventional suction bucket foundation model with the same dimensions and material parameters was established in this study. The diameter (D) and height (L) of the suction bucket were both 100 mm, with a wall thickness (t) of 5 mm. According to Xie et al. [21], the conventional suction bucket foundation, as used in this paper, was subjected to a sinusoidal reciprocating horizontal load with an amplitude of 1 mm, and the sinusoidal reciprocating horizontal load cycle was applied fifteen times. The simulation results are compared with the findings of Xie et al. [21], as shown in Figure 4.

The red curves in Figure 4 represent the load–displacement curves of a suction bucket from the literature [21] under 15 cyclic loading cycles, while the blue curves correspond to the results obtained from the model developed in this study under identical conditions. As shown in Figure 4, comparisons between the proposed model and the literature [21] reveal a maximum deviation of approximately 9% due to factors such as boundary condition discrepancies (which may differ from real-world scenarios) and the spatial variability of parameters. This deviation remains within acceptable limits, and the trend of the simulation results aligns closely with that reported in the literature [21]. The simulation results are consistent with the test results, which further illustrates the reliability of the modeling approach adopted in this paper.

3. Calculation and Application of Wind and Wave Loads

3.1. Calculation of the Wave Load

In conjunction with the loading formulas specified in the literature [22], the wave-parallel positive-force speed component and the inertia component acting on the full cross-section of the column at a height z above the bottom surface of the water can be calculated according to the following formulas:

$$F_D={\displaystyle \frac{1}{2}}{\displaystyle \frac{\gamma }{g}}{C}_{D}Du\mid u\mid$$

(1)

$$F_I={\displaystyle \frac{\gamma }{g}}{C}_{M}A{\displaystyle \frac{{\partial }_u}{{\partial }_t}}$$

(2)

$$u={\displaystyle \frac{\pi H}{T}}{\displaystyle \frac{ch\frac{2\pi z}{L}}{Sh\frac{2\pi d}{L}}}cos\omega$$

(3)

$${\displaystyle \frac{{\partial }_u}{{\partial }_t}}=-{\displaystyle \frac{2\pi 2H}{T2}}{\displaystyle \frac{ch\frac{2\pi z}{L}}{Sh\frac{2\pi d}{L}}}sin\omega t$$

(4)

$$\omega ={\displaystyle \frac{2\pi }{T}}$$

(5)

where γ is the gravity of seawater, N/m³; g is the acceleration of gravity, m/s²; C_D is the speed force coefficient of the wave, which is taken as 1.2 for the circular cross-section; D is the diameter of the tower, subjected to wave load m; u is the horizontal speed of the wave mass, m/s; C_M is the coefficient of the inertial force of wave, which is taken the as 2.0 for the circular section; A is the windward area of the fan blade of the offshore wind turbine (m²); t is the time, s; H is the wave height in meters of the wave in which the building is located; T is the wave period, s; ch is a hyperbolic cosine function; sh is a hyperbolic sine function; z is the vertical coordinate of the water quality point, m; L is the wavelength in m; d is the depth of water in which the building is located (the depth of water adopted in this study is 20 m); and ω is the original frequency of the wave, rad/s.

According to the literature [23], the above equations can be simplified to finally obtain equations related to the wave height and wave period, and the maximum drag force F_DMAX and the maximum inertia force F_IMAX of the wave are expressed as Equations (6) and (7).

$$F_{\mathit{DMAX}}=\left({\displaystyle \frac{{\gamma DC}_D}{2}}{\displaystyle \frac{2kd+sh2kd}{8sh2kd}}\right)H^2$$

(6)

$$F_{\mathit{IMAX}}=\left({\displaystyle \frac{\gamma \pi D2C_M}{8}}thkd\right)H$$

(7)

where k is 2π/L and th is the hyperbolic tangent function.

The maximum value of the wave load is calculated by Equation (8).

F_{wave MAX} = F_DMAX + F_IMAX

(8)

Considering that the wave load increases with water depth, the total wave load is taken as shown in Equation (9).

F_{wave, total} = F _{wave MAX} d/2

(9)

3.2. Wind Load Calculation

The wind load was calculated using Equation (10) [24].

ω_k = β_zμ_sμ_zω₀

(10)

where w_k is the standard value of wind load; βz is the wind vibration coefficient at height z, which is a dynamic value factor, and the specific value is calculated by Equation (11); μ_s is the coefficient of the wind load for the body type (taken as 1.2 for the most unfavorable body type coefficient when considering the whole structure according to the circular cross-section); μ_z is the coefficient for the variation in wind pressure due to height (for the offshore wind turbine in the environment for a Class A area, this paper takes the fan blade installation location from the sea level height of 70 m and the wind pressure height variation coefficient as 2.05); w₀ is the basic wind pressure, and according to the wind speed–wind pressure relationship derived from Bernoulli’s equation, it can be seen that w₀ = 0.5 ρv², where ρ is the air density kg/m³; and v is the measured wind speed m/s at that point.

The wind vibration coefficient at height z was taken as given in Equation (11).

$${\beta }_z\mathrm{ }=\mathrm{ }1\mathrm{ }+\mathrm{ }2g{I}_{10}{B}_Z\sqrt{1+R^2}$$

(11)

where g is the peaking factor, taken as 2.5; I₁₀ is the nominal turbulence intensity at a height of 10 m and is taken as 0.12 because the environment in which the offshore wind turbine operates is a Class A area; R is the resonant component factor of the pulsating wind load; B_Z is the background component factor for pulsating wind loads.

The resonance component factor R of the pulsating wind load was calculated using Equation (12).

$$R=\mathrm{ }\sqrt{{\displaystyle \frac{\pi }{6{\zeta }_1}}{\displaystyle \frac{X_1^2}{{(1+X_1^2)}^{4/3}}}},\mathrm{ }{X}_1\mathrm{ }={\displaystyle \frac{30}{T_1\sqrt{k_w{w}_0}}}\mathrm{ },X_1>5$$

(12)

where $T_1$ is the first-order self-oscillation period of the structure, calculated according to structural dynamics and taken as 0.013 H; $k_w$ is the ground roughness correction factor, which is taken as 1.28 for Class A environments; and ${\zeta }_1$ is the structural damping ratio, taken as 0.01; when the value of X₁ ≤ 5, it is taken as 5.01.

The background component factor B_Z for pulsating wind loads was calculated by Equation (13).

$$B_Z\mathrm{ }=\mathrm{ }k{H_B}^{a_1}{\rho }_x{\rho }_z{\displaystyle \frac{{\Phi }_1\left(z\right)}{{\mu }_z}}$$

(13)

where ${\Phi }_1\left(z\right)$ is the first-order vibration coefficient of the structure, which takes a value of 1 according to the location of the turbine when the height of the tower structure is 100%; $H_B$ is the total building height, which according to the specification requirements of a Class A environment cannot exceed 300 m (in this study, the tower height was 90 m, which is lower than this value, in line with the requirements); k and $a_1$ are the correlation coefficients, which are taken as 1.276 and 0.186, respectively, for Class A areas; and ρ_x and ρ_z are the horizontal and vertical correlation coefficients of the pulsating wind, respectively.

$${\rho }_x={\displaystyle \frac{10\sqrt{B+50e^{-B/50}-50}}{B}}$$

$${\rho }_z\mathrm{ }=\mathrm{ }{\displaystyle \frac{10\sqrt{H_B+60e^{-H_B/60}-60}}{H_B}}$$

where $H_B$ is the total building height (m) and B is the windward width of the structure (m).

The total wind load was calculated using Equation (14).

F_wind = w_k·A

(14)

where A is the windward area of the fan blade of the offshore wind turbine (m²).

3.3. Application of Wind and Wave Loads

Based on the literature [25], the maximum daily wind and wave monitoring data for 2022 monitored by a station in the eastern China Sea were collected. The data were converted to extreme wind and wave loads and applied to the suction bucket foundations following the calculations shown in Section 3.1 and Section 3.2 to investigate the corresponding load-bearing performance. The wind and wave loads listed in this paper are only for 353 days because 12 days of the monitoring data were missing. The extreme wind and wave monitoring data are shown in Figure 5 and Figure 6. Wind and wave loads were superimposed in chronological order to generate the load distribution curves shown in Figure 7. Figure 8 shows the directional distribution of wind and wave loads throughout the year. The wind and wave load directions in Figure 8 are obtained from the data. Firstly, the daily maximum wave height, wind speed, and corresponding wind direction monitored by a station are measured to determine the extreme value data set; secondly, the annual wind and wave data in the same direction are superimposed to obtain Figure 8.

The wind and wave loadings were realized using the magnitude function in ABAQUS finite element software. The load was assumed to be applied at the center point of the lower plane of the bucket top. Considering the consistency of the coordinate system between ABAQUS 2020 software and the one shown in Figure 8, the right-hand rule was used to establish the coordinate system [26]; i.e., the direction of the positive x-axis starts at 0°. The loading direction was rotated clockwise through 360°, as illustrated in Figure 9.

4. Analysis of the Carrying Performance

4.1. Characteristics of the Bucket Displacement

Figure 10 shows the time–history curves of the displacements in the x-axis and y-axis of the two types of bucket foundations after the extreme wind and wave loads throughout the year. As shown in Figure 10, the amplitude of the base displacement of the supported suction bucket is small and smooth. Conventional suction bucket foundations have large and unsteady displacement amplitudes. The positive and negative vibration extremes on the x-axis are 122.2 mm for the conventional suction bucket and 47.2 mm for the supported suction bucket, demonstrating a 61.37% amplitude reduction compared to the conventional design. In terms of the y-axis, the vibrational extremes extend to 168.0 mm for the base of the conventional suction bucket, whereas the supported suction bucket exhibits a mere 55.4 mm, reflecting a remarkable 67.02% decrease in vibrational amplitude. Supported suction bucket foundations can effectively reduce bucket foundation displacement. The main reason for this is that the introduction of the supporting structure markedly alters the load transfer path and the bucket–soil interaction mechanism. Under the action of wind and wave loads, conventional suction bucket foundations primarily rely on the frictional resistance between the bucket wall and the soil, along with the soil resistance generated by the negative pressure within the bucket, which tends to result in the bucket’s overall translation and rotation. However, the braced suction bucket, through rigid bracing rods, forms a multi-point anchoring mechanism with the external soil, enhancing the reactive forces against the bucket body, thus reducing bucket displacement and effectively increasing the stability of the suction bucket foundation.

Figure 11 shows the z-axis displacements of the two suction bucket foundations after extreme wind and wave loading throughout the year. Since the z-axis direction is downward, positive values in the z-axis are specified in this paper to represent the sinking of the suction bucket foundation after wind and wave loads, while negative values are specified to describe the uplift of the suction bucket foundation after wind and wave loads. As shown in Figure 11, the vertical displacement of the conventional suction bucket foundation fluctuates greatly, mainly upward. The vertical displacement of the supported suction bucket foundation fluctuates little, is less likely to be uplifted, and is primarily characterized by a lesser degree of subsidence. The maximum uplift displacement of the conventional suction bucket foundation is 44.3 mm. In comparison, the supported suction bucket foundation is 4.7 mm, and the uplift reduction of the supported suction bucket foundation over the conventional suction bucket foundation is about 89.49%. The main reason is that the support structure increases the contact area between the bucket and the soil. By extending the load distribution range, the lateral frictional resistance is enhanced, and the stress concentration in localized areas of the bucket is reduced. Moreover, the bracing rods create additional bending resistance nodes on the outer side of the bucket, effectively utilizing the weight of the soil above the support to significantly increase the bucket’s resistance to overturning moments. These factors impede the bucket from pulling up, effectively improving the load-bearing capacity of the bucket.

Figure 11 also shows that on 17 March 2022, the pullout of both bucket bases reached a maximum value. Considering that the bucket is a rigid body, the lid was selected for analysis. Figure 12 is the morphological map of the lid position of the two bucket bases extracted on 17 March 2022, with the circles representing the position of the outer contour of the top cover.

Both suction bucket foundations tilt after being subjected to the wind and wave loads, as shown in Figure 12. The maximum pullout at the top of the conventional suction bucket foundation is 118.4 mm, the maximum displacement of the sinking is 8.3 mm, and the conventional suction bucket is tilted by 0.4° compared to its initial position. The maximum pullout of the supported suction bucket foundation is 28.7 mm, the maximum displacement of the subsidence is 19.4 mm, and the supported suction bucket is tilted by 0.15° compared to its initial position. The supported suction bucket foundation demonstrated a 75.76% reduction in uplift resistance, a 133.73% increase in settlement, and a 62.5% decrease in tilt compared to the conventional design. Further, the supported suction bucket exhibits reduced pull and tilt under wind and wave loads; thus, this foundation is more stable.

Further considering the positional relationship between the bucket lid and the internal soil surface under uplift and tilting conditions, Figure 13 is obtained. As shown in Figure 13, the conventional bucket foundation lid exhibits an uplift difference of 26.45 mm compared to its internal soil plane, resulting in detachment between the lid and the topsoil layer. In contrast, the supported bucket foundation demonstrates a significantly smaller uplift differential of 3.45 mm, indicating essentially no separation from the internal topsoil. The internal soil surface of the conventional suction bucket is inclined by 0.37° compared to the initial plane. Meanwhile, the supported suction bucket is inclined by 0.13°, representing a 64.86% decrease in tilt magnitude. These results demonstrate that the supported suction bucket foundation enables coordinated load-bearing behavior between the bucket structure and soil mass, thereby significantly enhancing anti-overturning performance.

4.2. Displacement and Plastic Characteristics of the Soil Around the Bucket

Figure 14 and Figure 15 show the displacement clouds of the soil around the buckets for the two types of bucket foundations after being subjected to extreme wind and wave loads throughout the year. The displacement of the soil around the perimeter of the conventional suction bucket is affected over a larger area, whereas that around the perimeter of the supported suction bucket is affected over a smaller area. Combined with Figure 8, the distribution and magnitude of wind and wave extreme loads throughout the year are regional, with a significant concentration in the positive direction of the y-axis (near 90°) and the 18–162° interval. This indicates that the extreme loads in this sea area have a clear directional preference. This directional feature directly affects the displacement distribution of the soil around the foundation of the conventional suction bucket, resulting in the soil displacement around the conventional suction bucket foundation being local, i.e., large displacement occurs in some areas and small displacement occurs in others. The regional characteristics depend on the direction of wind and wave action in the offshore region where the wind power is located. The overall displacement of the soil around the perimeter of the supported suction bucket is more uniformly distributed.

Further extraction of the maximum soil displacement around and inside the perimeter of the two bucket foundations is shown in Figure 16. The conventional bucket foundation exhibits a large displacement of both the inner and outer soil of the bucket after one year of extreme wind and wave loading. The maximum displacement of the soil on the outer side of the bucket occurs at 81°, which is about 97.7 mm, and the overall displacement is irregularly distributed. The displacement in the direction of 18° to 162° is much larger than the displacement in the other directions, and the soil distribution on the inside of the bucket is more regular compared to that on the outside. A ‘peanut’ shape on an axis from 63° to 243° is formed with a maximum displacement of the inner soil of about 50.2 mm. In conjunction with Figure 8, it is observed that the direction of maximum displacement is essentially the same as the direction in which wind and wave loads frequently act in the region during the year. The supported suction bucket foundation experiences a year of extreme wind and wave loading with minor variations in displacement for both the inner and outer bucket soils. This phenomenon can be attributed to the active constraint put on the soil displacement direction by the supporting structure: when wind and wave loads act in a particular direction, the opposing support counters the displacement trend through compressive reactive forces, thus mitigating the directional impact of loads on soil deformation. The lateral soil outside of the bucket foundation exhibits a maximum displacement (14.7 mm) in the 90° direction, while the internal soil displacement remains relatively consistent across all angular positions at approximately 10 mm, demonstrating a circular distribution pattern. The maximum soil displacement of the external soil in the supported bucket foundation shows an 84.95% reduction compared to conventional bucket foundations, with an 80.08% decrease observed in the maximum internal soil displacement. These results indicate that the added support structure effectively improves the stability of the bucket foundation.

Figure 17 displays nephograms of the plastic strain distribution in the peri-bucket soil for the two suction bucket foundation types. Analyzed in conjunction with Figure 8, the positive y-axis direction exhibits a higher annual frequency of wind–wave load applications compared to the negative y-axis direction. Therefore, the conventional suction bucket foundation in Figure 17 shows a larger plastic strain region in the positive y-axis direction with a maximum plastic deformation of 0.06 in the soil and a smaller plastic region in the negative y-axis direction. The supported suction bucket foundation shows only a small plastic strain zone in the positive y-axis direction, with the maximum plastic deformation of the soil decreasing to 0.0079 and no plastic region in the negative y-axis direction. Moreover, increasing the support can effectively reduce the plastic strain of the soil around the bucket.

Figure 18 shows the plastic strain distribution in different directions in the soil inside and outside the bucket for both bucket foundations. The large figure on the left shows the plastic strain distribution of the soil inside and outside the bucket for both types of buckets. Since the plastic strain distribution area of the supported suction bucket is small in the left-side figure (pink and black lines in the figure), to see the plastic strain distribution characteristics of the supported suction bucket more intuitively, the plastic strain distribution of the supported suction bucket is enlarged to obtain the small figure on the right side. Therefore, the small figure on the right side depicts the distribution for the supported suction bucket foundation. Figure 8 shows that the loading is primarily concentrated at specific angles (such as 18–162° and along the positive y-axis). This directionality is highly correlated with the plastic strain distribution in Figure 18: it is evident that the regions with greater plastic strain inside the soil of the conventional bucket foundation (e.g., along the positive y-axis) correspond to the directions of high-frequency loading in Figure 8. The frequent and intense loading in this direction leads to significant plastic deformation of the soil around the bucket. For the conventional bucket foundation, the plastic strain curve area of the inner soil of the bucket is larger than that of the outer soil. The maximum plastic strain of the outer soil is about 0.075, and the maximum plastic strain of the inner soil is about 0.209. Due to the large load in the positive direction of the y-axis, the internal soil of the bucket in the negative direction of the y-axis is subjected to a large strain after being loaded. The plastic strain of the soil around the supported suction bucket foundation is significantly reduced. The plastic strain curve area of the outer soil of the bucket foundation for the supported suction bucket is slightly larger than the area of the plastic strain curve of the inner soil, with a maximum plastic strain of approximately 0.016 and 0.009 in the outer and inner soils, respectively. The analysis indicates that the added support structure disperses the load transfer path, resulting in more uniform loading of the soil. Although the loading direction in Figure 8 is still concentrated in specific areas, the support structure significantly reduces the localized concentration of plastic strain by enhancing the soil–structure interaction. Moreover, the supported bucket foundation makes more use of the bearing properties of the wider range of soil on the outside to help the inner soil share the energy applied. Consequently, the disparity in the plastic strain curve areas between the soil inside and outside of the perimeter of the bucket is significantly reduced, and the soil is more uniformly loaded, which is more conducive to improving the stability of the bucket foundation.

5. Conclusions

By taking the daily wind and wave monitoring extreme value data from a measurement station in the eastern China Sea area in 2022 and applying them as loads to the conventional and supported suction buckets, the load-bearing characteristics of the two kinds of bucket foundations were analyzed using numerical simulations. The following conclusions are drawn from the obtained results.

(1)

Compared to the conventional suction bucket foundation, the supported suction bucket foundation has a positive and negative extreme value reduction of 61.37% in x-axis displacement vibration and 67.02% in y-axis displacement. On the z-axis, the reduction in bucket uprooting is about 89.49%.

(2)

The supported suction bucket foundation has a pullout reduction of 75.76% and a tilt reduction of 62.5% compared with the conventional suction bucket foundation, effectively reducing the pullout and tilt of the bucket under extreme wind and wave loads. The supported suction buckets reduce lid detachment from the topsoil by about 86.96% more than conventional suction buckets. The soil plane within the bucket is inclined by 64.86% compared to the initial plane, enabling the bucket and soil to be loaded in tandem, thus increasing the resistance against overturning.

(3)

Considering the year-round directionality of the wind and waves in the monitored area, the displacement of the outside soil of the conventional suction bucket foundation is larger in the range of 18° to 162°, while the displacement of the inside soil has a ‘peanut’ shape on the axis at 162°. The displacement of the inner soil of the supported suction bucket foundation is about the same at all angles, whereas the displacement of the outer soil is slightly larger than that of the inner soil and is accompanied by slight fluctuations.

(4)

The plastic strain area in the inner soil of the conventional suction bucket foundation is greater than that in the outer soil. Moreover, the plastic strain area in the inner soil of the supported suction bucket foundation is slightly less than in the outer soil.