Structural Safety Evaluation of a Fairlead Chain Stopper for the Disconnectable Mooring System of Floating Offshore Wind Turbines

2.1 Floater Design

For the development of a disconnectable mooring system, the floating offshore wind turbine was based on a semi-submersible floating structure paired with the DTU 10 MW offshore wind turbine (Bak et al., 2013). The semi-submersible structure comprised three outer columns and a central column supporting the turbine tower, all interconnected by a lower pontoon and an upper deck, as shown in Fig. 1. The interior of the floating structure was divided into three ballast tanks, with the main specifications listed in Table 1. The system was designed for an installation depth of 150 m. using three mooring lines arranged in a catenary configuration.

Fig. 1

Semi-submersible floater design

Table 1

Principal dimensions of the FOWT

2.2 Mooring System

The mooring system used three lines with a catenary configuration, as shown in Fig. 2. The key components of the disconnectable mooring system included the FCS, which featured a bidirectional latch system for easy chain attachment and detachment, and the SMP installed underwater to facilitate tensioning operations. The newly developed FCS enabled safer and faster detachment compared to the existing systems used in floating platforms, significantly reducing the installation and decommissioning times for floating offshore wind turbines. The detachable mooring system developed in this project was designed to accommodate a 147 mm studless R4S grade chain capable of withstanding an MBL of 21,179 kN.

Fig. 2

Developed mooring system configuration (image by Korea Offshore Marine Solution)

3.1 Structural Analysis Model – FCS

Fig. 3 presents the FCS, a key component of the developed disconnectable mooring system. The FCS was mounted on the upper section of the outer column of the floating structure, securing the mooring lines to the floating body. One FCS unit weighed 32 t, with the materials and yield strengths for each component detailed in Table 2. The components and their respective functions are outlined below.

Fig. 3

Components of the FCS

Table 2

Material properties

3.2 Finite Element Analysis Model

Finite element analysis (FEA) was performed using solid elements (C3D8R), with the material properties defined by an elastic modulus of 206 GPa and a Poisson’s ratio of 0.3. The accuracy of the nonlinear contact analysis was ensured by convergence analysis conducted by varying the element size at the contact interfaces. The equivalent stress results were plotted as a function of the element size, showing that stress convergence occurs at 30 mm or smaller for the plate and 20 mm or smaller for the pin (Figs. 4 and 5). Based on these results, the element size at the contact interfaces was set to 15 mm or less. The final finite element model consisted of 313,621 elements and 418,921 nodes. Fig. 6 presents the generated finite element model, which includes the FCS and a portion of the floater’s support structure.

Fig. 4

Mesh convergence test model

Fig. 5

Mesh convergence test results

Fig. 6

FE model for structural analysis

3.3 Boundary Conditions

The numerical analysis model for strength assessment included the FCS and a portion of the supporting structure. The end of the supporting structure connected to the floater was defined with a fixed boundary condition (Fig. 7(a)). A half-model was used with the symmetry boundary conditions in Fig. 7(b) to minimize the computational cost. A nonlinear analysis incorporating the contact conditions was performed, with the contact surfaces defined at the pin joint and the wall interfacing with the stopper block, as shown in Fig. 8. Table 3 lists the applied friction coefficients. For the pin-connected regions, the friction coefficient between the steel and bushing was applied, while the friction coefficient between steel and steel was applied for the wall surfaces.

Fig. 7

Boundary conditions for structural analysis

Fig. 8

Contact locations for structural analysis

Table 3

Contact conditions

4.1 Load Application

The loads used for the structural strength evaluation were based on the minimum breaking load (MBL) of the mooring chain. The MBL for the 147 mm diameter R4S-grade chain was calculated using Eq. (2) in accordance with DNV-OS-E301 (DNV, 2021a). Fig. 9 shows the design angles at which the loads are applied. In this analysis, a horizontal design working range (DWR) angle of 0° and vertical design inlet angles (DIA) of 10° and 24.5° were considered. The 10° angle represents the minimum requirement set by DNV regulations, while the 24.5° angle, derived from integrated load analysis, is the angle at which the maximum mooring line tension occurs. MBL was applied using the bearing pressure to the chain contact surface. The bearing load was calculated using Eqs. (3)–(5) and applied at each node of the contact surface. The applied direction is the uniaxial load in the direction of chain movement, as shown in Fig. 10.

Fig. 9

Load application angles

Fig. 10

MBL load application area

4.2 Structural Analysis Results

Structural analysis was conducted using Abaqus 6.22, and the final displacements and equivalent stress values were evaluated (SIMULIA, 2022). Fig. 11 presents the resulting displacement distribution, where a maximum displacement of up to 8 mm is observed in the direction of the tensile force acting on the FCS arm. Fig. 12 shows the equivalent stress distribution under each analysis condition, with the maximum equivalent stress of 629 MPa occurring at the load-applied surface of the chain stopper. In addition, high equivalent stress values were observed at the main pin and the FCS support, which appeared to be caused by uneven contact at the support plate because of the bending deformation of the main pin. The FCS arm, where large displacements occurred, showed a relatively uniform distribution of equivalent stress. In contrast, high stress was observed at the intersection of the base plate and wall plate of the FCS housing, where little displacement occurred. The cause of this stress concentration was analyzed during the evaluation process.

Fig. 11

Displacement contour for Inlet angle of 10° (Deformation scale factor is 50)

Fig. 12

Equivalent stress contour of the entire model

4.3 Structural Strength Evaluation

Structural safety evaluation of the FCS was conducted using an allowable stress set to 90% of the yield strength of the material following the DNV-ST-0119 (DNV, 2021b). Table 4 lists the results of the structural analysis, including the maximum equivalent stress and its compliance with the allowable stress criteria. The analysis confirmed that all components satisfied the specified criteria, except for the wall plate of the FCS housing, where the maximum equivalent stress of 291 MPa exceeded the allowable limit of 279 MPa. Fig. 13 presents the stress distribution on the wall plate of the FCS housing. The elevated stress observed at the intersection of the base plate and wall plate was attributed to the use of a fine mesh (20 × 20 × 20 mm) and abrupt geometric transitions, resulting in localized peak stresses. Considering that peak stresses can reach up to 1.7 times the allowable stress, the overall structural strength requirements were still deemed to be satisfied (DNV, 2021c). Therefore, the designed FCS met the structural strength criteria.

Table 4

Strength evaluation results of the FCS

Fig. 13

Areas exceeding allowable stress

5.1 Fatigue Load

Loads with significant influence on fatigue were applied to evaluate the fatigue strength, with a particular focus on the tensile loads induced by the mooring lines. The range of these tensile loads and the number of cycles expected over the design life were derived from the integrated load analysis results provided by a collaborating research institution. The range of tensile loads and their corresponding cycle counts were extracted using the Rainflow counting method from the time series data of mooring line tensions. Fatigue loads corresponding to a 25-year service period were calculated by considering the occurrence probabilities of wind speeds under each design load case. The design load cases, based on IEC 61400-3-2 (IEC, 2019), were selected with reference to the LIFES 50+ project and are summarized in Table 5 (EU, 2015). Table 6 lists the mean tension and corresponding angles of the mooring lines under each design load case. Based on this result, the fatigue loads were calculated for mooring line 2, which exhibited the highest mean tension. Fig. 14 shows the fatigue loads applied to the FLS, where the tension range was divided by 10 kN intervals.

Table 5

Design load cases for integrated load analysis

Table 6

Mean pivot angle, wrap angle, and mean tension

Fig. 14

Fatigue loads of mooring line 2 (for 25 years)

5.2 Screening

The purpose of screening is to select areas relatively vulnerable to fatigue. For this, unit tension of 1kN was applied to the chain, and areas showing high principal stress were selected. The fatigue strength of these identified vulnerable areas was then assessed. Fig. 15 shows the screened zones along with maximum principal stress distributions, and Table 7 lists the maximum (tensile) principal stress and directionality of screened zones. Zone 11 was selected to evaluate fatigue failure at the welded joint of the FCS support structure.

Fig. 15

Floating system general arrangement

Table 7

Hot spot stress from unit tension for the screened zones

5.3 S-N Curve

The designed FCS was evaluated using the fatigue S-N curves applicable to air environments (DNV, 2020). Fig. 16 presents these applied fatigue curves, and Table 8 lists the applied fatigue S-N curves for each component.

Fig. 16

S-N curves

Table 8

Input parameters for fatigue life calculations

5.4 Fatigue Strength Evaluation

The fatigue strength of the FCS was evaluated to determine its fatigue life and compliance with the design fatigue factor (DFF) requirements. A design life of 25 years was assumed for this evaluation, comprising 20 years under normal operating conditions and five years under non-operating conditions, with a DFF of 10. The fatigue life was calculated using Palmgren–Miner’s linear cumulative damage method, as expressed in Eq. (6) (DNV, 2020). In the case of toe cracks, a stress level higher than the hot spot stress was applied to obtain conservative fatigue assessment results. The hot spot stress is typically determined by extrapolating the maximum principal stress obtained from FEA with a mesh size equal to the plate thickness. In this study, however, a higher stress level was used intentionally to make the results more conservative. A very fine solid mesh with an element size of approximately 15 × 15 × 15 mm, corresponding to approximately 25% of the minimum plate thickness (ranging from 60 to 130 mm), was used to achieve this. The principal stress at the intersection element was extracted and applied for the fatigue safety evaluation. In the case of root cracks, the stress value associated with potential crack initiation in the weld throat of the welded joints was calculated using Eq. (7). Table 9 lists the results of the fatigue strength evaluation. All screening areas exhibited a DFF greater than 100, satisfying the fatigue strength criteria. Nevertheless, the DFF may be reduced if the FCS were operated in a seawater environment.

Table 9

FCS fatigue analysis results

where the stress components are explained in Fig. 17.

Fig. 17

Stresses on the throat section of fillet weld (DNV, 2020)