L. Kellezi & S. Sundararajan, O. Koreta

1 SUBSTATION FOUNDATIONS – SOIL INTERACTION ANALYSES

1.1 Introduction and scope of work

The offshore substations are offshore platforms con-taining electrical components to connect an offshore energy generation project (for example, multiple wind turbines) to the onshore electricity network. Offshore substations build on the years of experi-ence of the offshore oil and gas industry, and the most common designs, use a platform consisting of a ‘topside’ in which the main equipment is housed, and a foundation structure, which is either a steel ‘jacket’, a ‘monopile’ or a gravity base structure.

The current substation is designed as a four - legs steel jacket structure (Fig. 1) temporarily founded on mudmats and finally on monopile foundations for permanently supporting the jacket and the topside structure. The mudmats are designed as skirted foundations equipped with an open pile sleeve, where the steel monopiles will be guided through during driving or installation.

The scope of work for the mudmat design in-cludes conventional penetration prediction and as-sessment of skirted mudmats yield capacity and sta-bility for combined (horizontal (H) – Vertical (V)) environmental load conditions. Regarding the monopiles, vertical and torsion capacities and de-formations, are investigated. Other aspects of the

monopile design will be presented in future work. Research on large skirted foundations is among oth-ers, presented by (Kellezi et al 2008, 2005), and on monopiles, by (Kellezi & Hansen 2003, 2002).

1.2 Location of substation and foundation types

The offshore wind farm comprises 140 wind tur-bines and one substation structure.

Figure 1. Offshore substation, steel jacket structure, mudmats and monopiles

Innovative design of foundations - offshore wind farm substation

L. Kellezi & S. Sundararajan, O. Koreta GEO, Lyngby, Denmark

ABSTRACT: As part of the efforts to develop more sustainable energy, wind power continues to be the pre-ferred solution, and offshore wind farms, equipped with transformer platforms / substations are constructed worldwide. Design of substation platform foundations for a wind farm offshore UK is the focus of this paper. The integrated structure-foundation design process, assuming initially simplified soil-structure interaction and then, the detailed design of the temporary jacket foundations (mudmats) and permanent substation founda-tions (monopiles), are presented. The conventional and 2D & 3D FE analyses of the skirted mudmats, for in-stallation and environmental loads, are carried out. Some aspects of the monopile design are also presented. Analyses applying different conventional methods in comparison to FE axisymmetric (for axial loading) and 3D FE (for torsion loading) modelling are elaborated and discussed. Important conclusions and recommenda-tions, applicable to the engineering design of these types of offshore foundations, are drawn.

L. Kellezi - Innovative design of foundations 1

According to bathymetry data, the water depth at the substation location is about 19 m and the seabed is generally flat.

1.3 Soil data and soil interpretation

The soil conditions at the site are assessed based on soil data elaborated from different soil investiga-tions, comprising four shallow cone penetration tests (CPTs), one at each leg location, respectively, few deep seabed CPTs and one deep borehole located near the jacket legs.

The seabed soil is described as a sediment layer consisting of very soft, becoming firm variably silty clay, underlain by loose to medium dense, becoming dense fine to coarse sands. Further down, mud faci-es, clay till and mudstone are found to larger depth. Considering the homogeneity of the soil conditions, the soil profile applicable to the design of the mud-mat and monopile foundations is generalized and represented by the CPT given in Figure 2.

Figure 2, Cone resistance, pore pressure and friction ratio as a function of depth at the generalized CPT

Based on the correlation of the laboratory strength tests and in-situ / field tests, the characteristic soil parameters are derived as follows: Table 1. Generalized lower bound soil profile

Soil type Depth γ φ cu G

- M kN/m3 ° kN/m2 MPa

Clay / Silt 0.0–1.2 6.0 - 3.1 10

Sand, loose 1.2–2.8 9.0 28 - 10

Different dense sand layers (2.8 – 20.6)

Mud Facies 20.6–22.0 10.0 30 140 210

22.0–22.5 10.0 35 360 220

22.5–23.1 10.0 - 80 110

23.1–25.1 10.0 - 50 60

Glacial Till 25.1–27.1 11.5 - 110 220

27.1–28.0 13.0 - 510 430

Mudstone 28.0–40.0 13.0 - 700 600

The undrained shear strength (cu) for the clay / silt layers has been interpreted by applying cone factor Nkt = 22 for lower bound parameters and Nkt = 9 for upper bound ones, on the net cone, (qnet) resistance; The angles of internal friction (φ) for the sand layers are generally derived from the CPT cone, (qc) re-sistance; The characteristic lower bound soil profile, generally interpreted as given in Table 1, is used in the current design of the foundations. For the mud-mats, the upper part of the soil profile has been in-terpreted based also on other shallow CPTs, located at each jacket leg, respectively.

The design parameters are derived by applying (DNV 1992) material factors (γm=1.2 on the tan(φ) and γm=1.3 on the cu). Design parameters are applied for the design of the mudmats and characteristic pa-rameters for the design of the monopiles.

2 MUDMAT PENETRATION AND (V-H) BEARING CAPACITY

2.1 Mudmat geometry and loads

The four mudmats of the substation structure (with distance from center to center (30 m x 30 m), are, af-ter several attempts, designed fit to the soil condi-tions, as skirted footings with the same geometry, consisting of an outside diameter of 4.5 m and skirt length of 3.1 m. Details can be seen in Figure 3.

Figure 3, Cross section of the designed steel mudmats

The mudmats are expected to be open at the pile sleeve, (which have an inside diameter of 2.45 m) for the monopiles to be positioned, go through and be driven to the designed depth. Different configura-tions of the mudmats are investigated considering the soil conditions, particularly the thickness of the upper very soft silt layer, and the size of the opening at the pile sleeve.

In order to increase the vertical bearing capacity or limit large penetrations due to low capacity at full base contact with the very soft seabed soil, the sleeve was finally extended below the mudmat plate,


to the top of the loose sand layer (Fig. 3). After dif-ferent assessments, the pile sleeve and the mudmat outer skirt are designed to be extended below the bottom of the top mudmat plate, to 1.5 m and 3.1 m depth, respectively, expected to achieve at least 0.3 m embedment / penetration in the more competent sand layers.

2.2 Skirted mudmat penetration prediction

Skirted mudmat penetration analyses are carried out conventionally based on ((DNV 1992) approaches, modified due to the chosen mudmat geometry. The calculation of the skirt and pile sleeve extension re-sistances is based on the results of the available shal-low CPTs. An average qc at a chosen interval of 0.05 m is calculated for the upper 3.1 m from the seabed. No soil plugging is assumed.

Skirt penetration resistance is a function of aver-age qc and empirical coefficients kp and kf, which re-lates qc to skirt tip resistance and skin friction, re-spectively as given from equation (1)

dzzqzkAdqAdkR cfscpp )()()()( (1)

where: d = depth of the tip element; kp(z) and kf(z) = empirical coefficients relating qc to end resistance and skin friction respectively; qc(z) = average cone resistance at depth z; Ap= tip area of penetrating el-ement; As= side area of penetrating element.

The skirt & sleeve soil resistances versus penetration depth is shown in Figure 4. The installa-tion load to be applied on the mudmats from the steel jacket self-weight, is calculated equal to 2170 kN / mudmat as indicated in this figure.

Figure 4, Skirt & sleeve-soil resistance versus penetration depth for the generalized CPT profile

The current results show that the skirt bearing capacities follow the same trend as the CPTs. For the installation load, most probable (MP) resistance, the mudmats will penetrate to full skirt penetration of 3.1 m, reaching full base contact with the seabed. From the variation of the soil strength with depth, and the maximum loads expected to act on the

mudmats during installation, until the monopiles are driven, it was assessed that the bearing capacity is not affected from the soil layers below (9-10) m depth. For the highest expected (HE) resistance, the mudmat penetrations are limited to about 2.8 m, not reaching full base contact with the seabed. This maens that a gap of about 0.3 m between the underside of the mudmat top plate and the seabed soil can exist, which could reduce the capacity of the mudmat under environmental loads with vertical components larger than the available installation load. This could result in additional skirt & sleeve penetrations, not necessary uniform at all four mudmats.

2.3 FE analyses of mudmat capacity / stability for (V-H) combined loads

Considering the seabed soil conditions and the mudmat geometry, 3D FE analyses, for environmen-tal loads, are found applicable. First 2D axisymmet-ric FE modelling (for (H=0 and V=yield) are carried out in order to validate the 3D modelling (Fig. 5). Both analyses are based on (Plaxis 2D V. 9.01) & (Plaxis 3D Foundation V. 2.2), respectively.

Figure 5. 2D axisymmetric FE model of mudmat-soil interac-tion under vertical compression loading Considering the soil parameters available, Mohr Coulomb (MC) soil model is initially chosen and applied for different soil layers, assuming undrained conditions for the clay layers and drained conditions for the sand layers. This means that linear elasticity is applied in the elastic range, combined with perfect plasticity in the plastic / non reversible range. Dila-tancy is included for dense sand applying dilation angle ψ = (φ – 30) degree. Hardening and NGI-ADP soil models are also attempted to investigate the ef-fects of stiffness hardening and anisotropy as alter-native to MC isotropic behavior. Updated mesh


(large deformation) analyses are generally per-formed.

The current 2D FE axisymmetric and 3D FE anal-yses are performed for mudmats (without piles posi-tioned / stabbed in the mudmats), in-placed at full base contact with the seabed, or penetrations (refer-ring to the skirt tip), equal to the skirt length of 3.1 m, (the assumption of mudmat penetrations corre-sponding to full base contact is based on observa-tions from installations of similar structures in the neighboring sites), and: - Without drain holes on the top mudmat plate; - With equivalent drain holes on the top mudmat plate (the total area of 144 circular drain holes of di-ameter 0.035 m, (which will allow the sea water to drain out of 12 mudmat compartments) is for numer-ical reasons, approximated to one ring of thick-ness=0.0125 m, and 12 holes of diameter=0.12 m for axisymmetric and 3D FE modellings, respectively);

As the vertical components of some of the envi-ronmental footings reactions are calculated to be larger than the vertical installation load, design soil parameters are applied in order to reduce the conse-quences of possible natural / virtual preloading be-ing the cause of additional mudmat penetrations.

In the 2D axisymmetric FE analyses (Fig. 5) (without drain hole), solid elements are applied for both, the top mudmat plate, the skirt and the pile sleeve extension. Plate elements are only applied at the upper part of the pile sleeve. In this model the skirt and the pile sleeve extension are modelled with the real thickness as shown in Figure 3, while the in-terface is modelled as another soil layer with re-duced strength (50 % for the clay cu and 33% for the sand tan(φ)), surrounding the skirt and the pile sleeve extension. The reason for this kind of model-ling is the large sensitivity of the skirt and sleeve tip bearing capacity on the mesh density. Different meshes are applied until a converged solution achieved.

The 3D FE model is constructed assuming sym-metry conditions, which mean half of the mudmat and surrounding soil is considered and half of the loads applied, (Figs. 6, 7).

Figure 6, 3D FE model of mudmat without (right) & with 12 equivalent drain holes (left)

A model ‘borehole’, with soil layers to 9.3 m depth as given in Table 1, and design soil parameters, is applied at the center of the footing. The interface el-ements associate the wall structural elements. In or-der not to overestimate the skirt and sleeve tip re-sistances the wall elements are applied 0.5 m deeper, but this part is not activated in the calculation phase, which means only the extended interface is active.

A uniformly forced displacement of maximum 0.3 m is applied on the mudmat in 2D axisymmetric analyses while footing reaction measured, defining the yield capacity to vertical load. Point load appli-cations are used in 3D for vertical and horizontal loads, where the vertical loads are applied on the mudmat plate and horizontal load at a certain depth, chosen at 2.6 m below seabed, (considering the level of skirt thickness reduction and soil conditions) as-suming possible location of footing rotation centre.

2.3.1 2D axisymmetric and 3D FE analyses results and interpretations

For the mudmat without & with equivalent drain holes on the mudmat plate, the vertical capacities from the 2D axisymmetric and the 3D FE models, respectively, comply well. In Figure 7 are shown the vertical displacements (uy) and failure pattern for the mudmat-soil 3D FE model with 12 equivalent drain holes on the top mudmat plate.

Figure 7. 3D FE model with 12 equivalent drain holes, V=3600 kN, H=200 kN, (applied after installation load). Vertical dis-placement (m) and failure pattern (plastic points)

Figure 8, 3D (V-H) yield envelopes (Mudmat with and without drain holes on the top mudmat plate)


The soil plug at the compartments created from the mudmat skirt and the pile sleeve extension is fully / not fully mobilized for without / with drains holes, respectively.

The vertical capacity for the mudmat with drain holes is as expected lower than without drain holes, due to evacuation of the very soft seabed clay through the holes. However, the vertical and (V-H) 3D FE capacities calculated from the safety analyses (loads*factor of safety) are generally less (and even less for the mudmat with drain holes) than calculated by increasing the loads, following the load vector, to yield achieved. The reduction of the (V-H) capacity for mudmat with drain holes is as expected particu-larly large for (V-H) loads with larger vertical com-ponents than the installation load. The results sum-marised in Figure 8 indicate however sufficient mudmat (V-H) capacities to support environmental reaction combined loads.

3 MONOPILE FOUNDATIONS – BEARING CAPACITY ANALYSES

3.1 Introduction and scope of work

Two types of monopile geometries with varying outer diameters OD=1.829 m (72’’) and OD=2.134 m (84’’), are initially considered. The thickness of the monopile tube decreases with depth from 60 mm to 45 mm. The monopiles length below mudline are calculated to be 28.0 m in order to insure some em-bedment in the mudstone. All ultimate bearing ca-pacity analyses, including compression and tension, are carried out for characteristic soil parameters (Table 1). The design axial capacities are derived by applying safety factors on the characteristic values.

3.2 Monopile axial capacities - DNV approaches

The calculations according to (DNV 1992) are car-ried out based on (α-β) method, where α-method is used for the cohesive soil assuming undrained be-havior and β-method for the cohesive soil assuming drained behavior.

For cohesionless soils only one method is de-scribed assuming drained behavior as most relevant. For the pile tip resistance, the capacity of the full section is compared with the capacity of soil plug, which is the sum of internal shaft and pile tip (ring) resistances, choosing the lowest one. The results of the axial design bearing capacities for the monopile with OD=84”, in compression and tension, respec-tively are shown in Figure 9 and in Table 2 as a function of the monopile driving depth to full pene-tration. Slightly different capacity is derived for drain or undrain parameters of the mud facies layers, which are investigated in both, drained and un-drained conditions (Table 2).

Figure 9. DNV axial bearing capacities (compression and ten-sion) OD = 84''

Table 2. Pile bearing capacities – DNV approaches

Monopile geometry (OD) Compression Tension

(mm) (kN) (kN)

1829 15816 6173

2134 18520 7297

3.3 Monopile axial capacities, FE approaches

2D axisymmetric FE modelling of monopile-soil in-teraction, subjected to axial loading or displacement is carried out using (Plaxis 2D V. 9.01).

Figure 10. 2D axisymmetric FE results (deformed mesh, plastic points, vertical displacements), mud facies drained, OD = 84''

The monopile is modelled as a weightless rigid body. To not overestimate the tip resistance, the in-terface is extended below the depth of the monopile tip. Soil plug is assumed to 90% of the pile length. A mesh sensitivity analysis is carried out until conver-gence achieved. The results of the design capacities from FE analyses are given in Table 3 for both monopile diameters and in Figure 10 for OD=84”. Table 3. Pile bearing capacities – FE approaches

Monopile geometry (OD) Compression Tension

(mm) (kN) (kN)

1829 16420 6000

2134 21250 6800


It can be seen that FE and DNV bearing capacities are very close. The FE compression capacities are slightly larger and tension capacities slightly small-er.

In Figure 11, the FE results are shown as reaction forces Fy (at the monopile tip) as function of pene-tration / pull-out for OD=84”. The monopile capaci-ties are derived as Fy*2π. The capacities for different mud behaviors are similar to DNV approaches.

Figure 11. Plaxis FE results, axial reactions (compression and tension), OD=84’’

3.4 Monopile torsion capacity, FE approach

Torsion capacity is calculated only for the finally chosen monopile with OD=84’’ and drained condi-tions for the mud layers based on the 3D FE analysis (Plaxis 3D Foundation V. 2.2) (Fig. 12). The same soil profile and parameters as applied for the 2D ax-isymmetric analyses are considered.

Figure 12, Failure pattern in torsion, total & vertical displace-ments, mud facies drained, OD=84”

A torsion moment equal to 2.0 MNm is assumed to act on the top of monopile. After this moment is successfully completed, a safety analysis is carried out. The torsion capacity is then calculated as the product of the applied torsion moment, times the safety factor. Alternatively, a large torsion moment

of up to 30 MNm is applied to failure. Based on both methods, the characteristic torsion capacity of the large monopile is derived being up to 4 times larger than the torsion moment expected to be developed during the life operation as shown in Figure 13, (Failure is assumed to occur at the turning point of the curve for the calculation phase with large torsion moment applied.

As the failure is at the monopile tip and top, the applied safety factor on the characteristic capacity of approximately 8 MNm, takes into account the cyclic effect which could lead to soil strength degradation.

Figure 13. Torsion bearing capacity, mud facies drained, OD=84”

4 CONCLUSIONS

The jacket structure is safely installed on mudmats, which reached more or less full penetrations equal to (3.0-3.2) m. The monopiles of OD=84’’ were suc-cessfully driven to the design depth and the topside is successfully operating.

5 REFERENCES

Det Norske Veritas (DNV). Foundations Classification Notes No. 30.4. February 1992.

Kellezi. L, G. Kudsk & H. Hofstede. 2008. Skirted Footings Capacity for Combined Loads and Layered Soil Condi-tions. 2nd BGA International Conf. on Foundations, ICOF 2008, June, Dundee, Scotland, Proc. Volume 1 page 923-935.

Kellezi. L, H. W. L. Hofstede & P. B. Hansen. 2005. Jack-up Footing Penetration and Fixity Analyses. Proceedings, Frontiers in Offshore Geotechnics, ISFOG 2005, Septem-ber, Perth, Australia, page 559 – 565.

Kellezi. L & P. B. Hansen. 2003. Static and Dynamic Analysis of an Offshore Monopile Windmill Foundation. Proceeding of BGA International Conference on Foundations, ICOF2003, Dundee, Scotland, 401-410.

Kellezi. L & P.B. Hansen. 2002. Dynamic Transient Analysis of a Monopile Windmill Foundation. Structural Dynamics, EURODYN’2002, Munich, Germany, 1315-1320.

Plaxis 2D Version 9.01. Finite Element Code for Soil and Rock Analysis, Delft University of Technology & Plaxis b.v. The Netherlands.

Plaxis 3D Foundation Version 2.2. Finite Element Code for Soil and Rock Analysis, Delft University of Technology & Plaxis b.v. The Netherlands.


L. Kellezi & S. Sundararajan, O. Koreta

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