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J. Ocean Eng. Technol. > Volume 38(3); 2024 > Article
Yu, Park, and Cho: Collision-Damage Analysis of a Floating Offshore Wind Turbine Considering Ship-Collision Risk


As the number of offshore wind-power installations increases, collision accidents with vessels occur more frequently. This study investigates the risk of collision damage with operating vessels that may occur during the operation of an offshore wind turbine. The floater used in the collision study is a 15 MW UMaine VolturnUS-S (semi-submersible type), and the colliding ships are selected as multi-purpose vessels, service operation vessels, or anchor-handling tug ships based on their operational purpose. Collision analysis is performed using ABAQUS and substantiation is performed via a drop impact test. The collision analyses are conducted by varying the ship velocity, displacement, collision angle, and ship shape. By applying this numerical model, the extent of damage and deformation of the collision area is confirmed. The analysis results show that a vessel with a bulbous bow can cause flooding, depending on the collision conditions. For damage caused by collision, various collision angles must be considered based on the internal stiffener arrangement. Additionally, the floater can be flooded with relatively small collision energy when the colliding vessel has a bulbous bow.


σT: Tensile strength (MPa)
σY: Yield strength (MPa)
σtr: True stress (MPa)
σY,tr: True yield strength (MPa)
σT,tr: True tensile strength (MPa)
σHS,tr: True hardening start stress (MPa)
σYD: Dynamic yield stress (MPa)
σTD: Dynamic tensile stress (MPa)
T: Tensile strain
Y: Yield strain
tr: True strain
Y,tr: True yield strain
T,tr: True tensile strain
HS,tr: True hardening start stain
∊̇: Strain rate (s-1)

1. Introduction

Offshore wind turbines offer advantages such as securing large installation space and abundant wind resources compared with onshore ones. However, complex risk factors arise from the sea. Among them, ultimate limit states (ULS), such as typhoons, are considered in the design phase of offshore wind turbines, whereas the accidental limit state (ALS), which may occur during the operating period, is not sufficiently considered. The ALS must be considered in the design phase because collision accidents with ships may occur during the operation of marine structures. Although the occurrence frequency is low, the risk of such accidents is high.
In 2020, the fixed offshore wind turbine structure of the Borkum Rifgrund wind farm in Germany collided with a 26-m-long high-speed vessel (Renews, 2020). In 2021, a jack-up vessel mobilized for the installation of a wind farm in the waters of Huizhou City, China collided with a substructure (BBN, 2021). In 2022, the monopile substructure of a fixed offshore wind farm in the Netherlands North Sea collided with the bulk carrier “Julietta D,” and the installed substructure was dismantled (Durakovik, 2022). In the Netherlands’ Hallandse Kust Zuid offshore wind farm, a cargo ship collided with a marine substation, which damaged the “HKZ Beta” jacket substructure (Windfair, 2022). In 2023, a cargo ship with 1,500 tons of grains collided with a wind turbine in the Orsted’s offshore wind farm in Germany (Mandra, 2023). In addition to the cases mentioned above, several collision cases may be undisclosed, and such collision accidents are expected to increase in future if offshore wind farms are expanded. Since floating offshore wind turbines (FOWTs) are likely to be constructed in a large wind farm, the collision of one unit may cause a series of accidents. Therefore, these collision risk factors must be examined for safe wind-farm operation.
Research on ship collision was initiated by Minorsky (1959), whereas Cloughley (1978) and Donegan (1982) highlighted the necessity for research on the collision of marine structures. Many studies pertaining to the collision of FOWTs are currently being conducted. Recently, Do et al. (2020) performed a drop impact test and a verification via numerical analysis for cylinders stiffened with rings and stringers. They developed a collision-damage estimation formula based on the verified numerical analysis. Ren et al. (2023) investigated the collision of an FOWT using a 1:30 scale model and compared the experimental and numerical analysis results. Vandegar et al. (2023) investigated the collision of a cylindrical FOWT and performed an analysis using a simplified numerical-analysis method.
In this study, the results of drop impact tests and numerical analysis were compared for the verification of the collision numerical analysis employed. The numerical analysis technique that showed good agreement with the test results was further analyzed. Do et al. (2023) conducted a collision-damage analysis for a semi-submersible floater that applied a taut mooring system and developed a residual-strength estimation formula based on the damage incurred. The floater of a floating wind turbine can be flooded below the sea level when damaged by ship collision. If the floater is inclined beyond the allowable inclination angle due to flooding, power production will cease. Even within the allowable inclination angle range, power-generation efficiency can be degraded. To predict the economic loss and damage caused by a collision accident at sea, one must identify the ALS that may occur during operation. Therefore, in this study, a 15 MW semi-submersible FOWT was selected, and scenarios were established for collision accidents that may occur during its operating period. In addition, damage to the floater and its flooding caused by collision were examined.

2. Analysis Technique Validation

2.1 Impact Test Model

Prior to performing collision-damage analysis, a comparative study was conducted with the results of the impact test for a stiffened cylinder model conducted by Do et al. (2018) to verify the numerical analysis technique. The impact test was conducted at the Structural Research Laboratory of the Department of Naval Architecture and Ocean Engineering, University of Ulsan. Fig. 1 shows the test facility. The test was conducted on a cylindrical model with both longitudinal stiffeners and ring stiffeners. This stiffener arrangement is common to most cylindrical floaters. The knife-edge geometry was adopted for to the striker, and the impact-test conditions are summarized in Table 1. Numerical analysis for the verification was conducted using Abaqus 6.22 software (2022). As for the constraints, fixed boundary conditions were imposed on the bottom of the cylinder end plate in the same manner as the impact test. In the study of Do et al. (2018), the element size of the finite-element model was determined via element-convergence verification, and the element size in the collision area was approximately 1.6 times the thickness. In this study, a finite-element model was constructed based on the above as shown in Fig. 2. The fine-element size (5 mm × 5 mm) was applied to the collision area, whereas the 20 mm × 20 mm size was applied to other areas.

2.2 Material Properties

The material properties obtained by conducting a tensile test on the specimen used in the experiment were secured from the literature (Do et al., 2018). The input data that exerted the most significant effect on the results of collision analysis were the material properties. However, the material properties may differ between the steel material used in the experimental model and that used in the floater design. Therefore, the results of the static and dynamic tensile tests conducted using tensile specimens from the steel material used in the actual structure are required to obtain accurate results. Normally, the yield strength can be known in the initial design phase of the actual structure. However, in this study, the ultimate strength and tensile strain were estimated from the yield strength of the material using Eqs. (1) and (2) proposed by Park (2017). Park derived the equations by the regression analyses of 7,737 mill sheets of mild and high-tensile steels. To calculate the plastic deformation caused by the impact load via numerical analysis, a true stress-strain curve that can consider the strain-hardening effect of the material is required. Thus, a true stress-strain curve was calculated using Eqs. (3)(5). In this instance, 0.0218 was applied from the estimation formula of Park for the strain at the start of the strain hardening. In the numerical analysis, both the actual material properties and the material properties calculated from the estimation formulas were applied. Fig. 3 shows the true stress-strain curve.
In addition, the stress-strain curve based on the strain rate was estimated by applying Eqs. (6)(9) proposed by Park (2017) to consider the strain-rate effect of the material with instantaneous deformation. In this instance, 0.001, 1, 10, 30, 50, 80, 100, 200, and 1000 s−1 were considered for the strain rate. Fig. 4 shows the stress-strain curve with consideration of the strain rate.

2.3 Validation Results

Collision analysis was conducted by applying the material properties defined above, and the displacement values at the point with the maximum displacement were compared. Fig. 5 shows the deformation history at the collision point of the numerical-analysis results. The average value of the section from 0.03 to 0.045 s after the spring-back was regarded as the final permanent deformation. Subsequently, the test and numerical-analysis results were compared, as shown in Table 2. The largest error of 22% occurred when the Cowper-Symonds equation was applied, and an error of 19% was observed even when the actual material properties were applied. Meanwhile, applying Park’s dynamic equation resulted in an error of 8% with respect to the test results. In addition, the numerical-analysis results obtained by applying Park’s estimation formula and the numerical-analysis results obtained based on the actual material properties derived from the tensile test were almost identical, that is, the difference in maximum deformation was only 0.2 mm. Fig. 6 shows the deformed model after the impact test and the numerical- analysis results. As shown in the figure, the deformed shape was almost identical. As for the collision analysis for the actual structure, numerical analysis was conducted in this study by applying Park’s material-property estimation formula and dynamic stress-strain curve equations. This is expected to yield more conservative results as compared with the damage to the actual structure.

3. Collision Analysis

3.1 Collision Scenarios

The appropriate collision scenarios must be selected for the collision analysis of a floater. Dai et al. (2013) examined the risk of ship collision that may occur during the life cycle of an offshore wind turbine. They considered the ships that may approach the floater during its operation. In this study, installation waters were not selected, and thus, ships that operate along the routes under ALS scenarios were excluded. Among the ships that can be utilized during the life cycle, only working vessels with relatively high potential for collision damage were defined as targets. Table 3 summarizes the working vessels mobilized based on the operational stage and purpose of FOWTs. Vessels such as crew-transfer vessels (CTVs) were considered in ULS design. However, they were excluded from collision analysis owing to their insignificant impact on structures due to small displacement. As for colliding ships, multipurpose vessels (MPVs), service-operation vessels (SOVs), and anchor-handling tug ships (AHTSs), which are ships with different bow shapes, were considered (see Fig. 7). Additionally, 1,500 and 4,000 tons of displacement were applied considering the cargo loads. The hydrodynamic effect by fluid viscosity only considered the added mass and 0.25 was applied as the added-mass coefficient in accordance with DNV standards (2021). The DNV standards recommend collision velocities of 2 m/s or higher for the ALS. Therefore, collision velocities of 2 and 5 m/s were considered in this study. Analysis cases that do not involve fracture at a velocity of 2 m/s were excluded from the analysis cases for a velocity of 5 m/s. At a velocity of 2 m/s, only the cases with the highest collision energy were considered for each collision ship type. In this study, the collision between the bow of the ship and the outer column of the floater was considered, and collision angles of 0°, 30°, 60°, and 90° examined. In particular, angles of 0° and 90° represent collision cases with vertical girder members in the outer column, whereas angles of 30° and 60° represent collision cases with longitudinal stiffeners. Thus, 0° and 60° were selected as the final collision angles. Table 4 summarizes the collision-analysis conditions. The energy of the striker was calculated using Eq. (10).

3.2 Material Properties

In the collision analysis, a yield strength of 355 MPa was assumed for high-tensile steels frequently used in the marine structural design. Material properties that consider the true stress-strain curve and the strain-rate effect were estimated using the estimation formula of Park (2019). The same material properties were applied to both the shell and stiffener. Fig. 8 shows the stress-strain curves of the material applied to the numerical-analysis model. In addition, a shear-fracture model was applied to consider fracture deformation. Meanwhile, the shear criterion was set based on the studies of Ringsberg et al. (2018) and Park et al. (2023), who specified 0.23 to 0.27 as the shear criterion based on a parametric analysis for validation. In this study, a value of 0.25 was applied as the criterion for the shear-fracture model.

3.3 Analysis Setup

A 15MW semi-submersible type of UMaine VolturnUS-S was applied as the floater (Christopher et al., 2020). The internal stiffener of the floater was designed in accordance with DNV standards, and watertight bulkheads were installed at 5 m above the still water level (SWL) and 3 m below the SWL. Vertical displacement was restrained at the bottom of the floater. Except for the collision column, constrains for horizontal displacement were imposed on both sides of the column at the bottom. Collision ships were modeled as rigid bodies. A mesh size of 50 mm × 50 mm, which is 1.6 times the thickness of the column shell, was applied to the collision area, whereas 400 mm × 400 mm was applied to the other areas to reduce the analysis time. Fig. 9 shows the applied boundary conditions and the arrangements of the floater and collision ship. Fig. 10 shows the created finite-element model. Yoon and Choung (2023) demonstrated that collision occurs instantaneously. Thus, the effect of hydrodynamic force on damage is limited. In the collision analysis of this study, hydrodynamic force was considered only as the added mass of kinetic energy, whereas the damage range and the occurrence of fracture were examined based on the collision-area geometry.

4. Collision Results

4.1 Results of Displacement and Deformation

The shell deformation results for the collision column were examined. Fig. 11 shows the maximum deformation distribution in the collision area under a collision angle of 0°. The maximum deformation of the shell was approximately 694 mm, which occurred during the collision of an SOV with a bulbous bow. Fig. 12 shows the x-direction displacement by case for a collision angle of 0°. As shown, the displacement distribution of the collision ship is consistent with the bow shape. Fig. 13 shows the maximum displacement at a collision velocity of 5 m/s. As presented, the displacement increased with the increase in the displacement of collision ship. This appears to be a valid result considering kinetic energy, and displacement more than doubled at a collision angle of 60°. Fig. 14 shows the displacement of the collision column in the xy-plane for Cases 3 and 4. The results indicate that a large deformation occurred at a collision angle of 60° due to the displacement of the entire column arising from the displacement of the floater pontoon. Table 5 summarizes the damage and fracture results for all collision cases. For the MPV and AHTS, the risk of flooding was low even after collision because the collision area and maximum deformation positions were above the sea level. In the case of SOV, the collision area and maximum deformation position were near the sea level, and damage occurred above and below the watertight bulkhead positions, as shown in Fig. 15. Therefore, when collision occurs, it is necessary to examine the occurrence of fracture in the shell and the flooding of the column.

4.2 Occurrence of Fracture

The occurrence of fracture in the shell of the collision column was examined. If fracture occurs in the shell, then flooding occurs in the area below the SWL or in the splash zone, thus jeopardizing the floater’s stability. In accordance with the standards, the fracture occurrence was similarly examined at 2 m/s, which is the lowest collision velocity of the ALS. Fig. 16 shows the equivalent plastic strain results of SOV-collision analysis. For a collision angle of 0°, no fracture occurred even at a high displacement of 4,000 t. In the case of 60°, fracture occurred even at a low displacement of 1,500 t. Fine fracture occurred at a collision velocity of 2 m/s with the smallest kinetic energy. Furthermore, the occurrence of fracture differed depending on the collision angle, because the longitudinal girders attached at 90° intervals from 0° prevented the plastic deformation of the shell (Fig. 17). These results confirmed that the shell deformation caused by collision is proportional to the collision energy, whereas the shell can be fractured even by small collision energy, depending on the collision angle and internal stiffener arrangement. Fig. 18 shows the extent of damage. Because fracture occurs between the sea level and the watertight bulkhead at 3 m below the SWL, the flooding section caused by collision is the VOID section near the SWL. The results confirmed that collision with ships featuring a bulbous bow, including the SOV that was applied as a collision ship, may cause the flooding problem. Thus, the damage stability must be further examined to secure the safety of FOWT systems after collision occurs. If the damage stability cannot be secured, it can be made possible by designing the minimum flooding section of the VOID tank between watertight bulkheads or by applying foam-filling technology to prevent flooding of the damaged area (Shin et al.,2019).

5. Conclusions

In this study, collision-damage analysis was performed on a 15 MW FOWT with a semi-submersible floater. First, the numerical-analysis results were compared with the impact-test results to verify the material properties applied in the numerical simulation. The results showed that the error of the numerical calculation was approximately 8% comparing with the experimental value. Collision scenarios for the accidental limit state were investigated for multi-purpose vessels, service operation vessels, and anchor handling tug ships, which can operate during the life cycle of offshore wind turbines. The extent of damage to the floater was examined under different collision conditions, such as the bow shape, displacement, collision velocity, and collision angle.
Analysis results based on the shape and displacement of the colliding ship confirmed that the extent of damage caused by collision was proportional to the kinetic energy of the striker. The maximum deformation of the shell (approximately 694 mm) occurred during collision with a ship featuring a bulbous bow (Case 7). The local deformation of the collision area was significant at a collision angle of 0°, whereas the total displacement of the floater column was significant at 60°. The results based on a collision angle of 0° showed that the displacement increased, and the shell deformation increased by approximately three times for barge-type ships with a large collision area. However, the shell deformation increased less (approximately 1.5 times) despite an increase in displacement for ships with a bulbous bow and knife-edge geometry. This suggests that displacement contributes significantly to the extent of damage for damage resulting from collision with a ship featuring a large collision area.
At a collision angle of 60° where only the small stiffeners were involved, fracture with a maximum area of 0.25 m2 occurred in the shell. Even at a collision velocity of 2 m/s with the lowest kinetic energy, fracture with an area of 0.01 m2 occurred. This shows that the occurrence of fracture is more significantly affected by the internal stiffener than by the maximum shell deformation. This indicates that an appropriate collision angle must be selected by considering the internal stiffener of the floater when establishing collision-analysis scenarios. In addition, collision with a ship featuring a bulbous bow may cause fracture in the collision area below the SWL, resulting in flooding in the floater. When collision analysis is conducted for a ship with a bulbous bow, damage can be evaluated conservatively. In the future, design improvements should be conducted to satisfy damage-stability criteria under flooding conditions in some compartments of the floater columns.

Conflict of Interest

The authors declare that they have no conflicts of interest.


This study was supported by the Korea Institute of Energy Technology Evaluation and Planning (KETEP) grant funded by the Korea government (MOTIE) (No.20213000000030).

Fig. 1.
Drop-test facility in University of Ulsan
Fig. 2.
Finite-element model for impact-test analysis
Fig. 3.
True stress-strain curve of materials
Fig. 4.
Stress-strain rate curves of cylinder shell
Fig. 5.
Schemes follow the same formatting. If there are multiple panels, they should be listed as
Fig. 6.
Comparison of the deformed model of the experiment and numerical prediction
Fig. 7.
Collision ships for analysis
Fig. 8.
Stress-strain rate curves of material
Fig. 9.
Model setup for collision analysis
Fig. 10.
Finite-element model for collision analysis
Fig. 11.
Deformation result of collision analysis for 0° collision angle
Fig. 12.
Displacement contour of collision analysis
Fig. 13.
Displacement results of collision analysis for 5 m/s velocity
Fig. 14.
Displacement and deformation results
Fig. 15.
Displacement results of Case 7
Fig. 16.
Plastic equivalent strain from collision analysis
Fig. 17.
Equivalent plastic strain of collision analysis
Fig. 18.
Extent of damage from collision analysis
Table 1.
Impact-test condition
Item SS-Striker-I
Drop height, H (mm) 1,600.0
Collision velocity, v (m/s) 5.6
Striker mass, M(kg) 500.0
Kinetic energy, Ek (J) 7848.0
Table 2.
Displacement results of impact test and numerical analysis
Item Impact test result Numerical analysis

Exp.1)-CS2) Emp.3)-CS Exp.-DE4) Emp.-DE
Displacement (mm) 36.8 43.8 44.8 39.7 39.9
Difference (%) - 19 22 8 8

1) Material properties form tensile test

2) Cowper–Symonds equation

3) Material properties obtained using empirical formula (Park, 2017)

4) Dynamic equation (Park, 2017)

Table 3.
Utilized vessels during FOWT life cycle
Phase Pre-Installation Installation Operation Decommission
Tasks Geotechnical and metocean survey FOWT Anchor Maintenance FOWT Anchor
Grid Substation Grid Substation
Table 4.
Collision-analysis cases
Case Vessel Displacement (t) Velocity (m/s) Collision angle (°) Kinetic energy (kJ)
1 MPV 1,500 5 0 23,438
2 MPV 1,500 5 60 23,438
3 MPV 4,000 5 0 62,500
4 MPV 4,000 5 60 62,500
5 SOV 1,500 5 0 23,438
6 SOV 1,500 5 60 23,438
7 SOV 4,000 5 0 62,500
8 SOV 4,000 5 60 62,500
9 AHTS 1,500 5 0 23,438
10 AHTS 1,500 5 60 23,438
11 AHTS 4,000 5 0 62,500
12 AHTS- 4,000 5 60 62,500
13 MPV 4,000 2 60 10,000
14 SOV 4,000 2 60 10,000
15 AHTS- 4,000 2 60 10,000
Table 5.
Collision-analysis results
Case Vessel Kinetic energy (kJ) Maximum displacement (mm) Maximum deformation (mm) Fracture
1 MPV 23,438 135.9 111.0 Not
2 MPV 23,438 698.7 689.0 Not
3 MPV 62,500 405.8 60.7 Not
4 MPV 62,500 2,557.0 217.0 Not
5 SOV 23,438 504.1 499.6 Not
6 SOV 23,438 833.0 824.5 Not
7 SOV 62,500 1,142.0 232.6 Occur
8 SOV 62,500 1,789.0 596.0 Occur
9 AHTS 23,438 144.1 129.8 Not
10 AHTS 23,438 825.4 824.5 Occur
11 AHTS 62,500 229.5 74.1 Not
12 AHTS 62,500 2,082.0 218.0 Not
13 MPV 10,000 388.8 53.1 Not
14 SOV 10,000 484.6 156.0 Occur
15 AHTS 10,000 367.2 55.5 Not


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