Encyclopedia of Ocean Engineering

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| Editors: Weicheng Cui, Shixiao Fu, Zhiqiang Hu

Capacity of Suction Anchor

  • Dengfeng FuEmail author
Living reference work entry
DOI: https://doi.org/10.1007/978-981-10-6963-5_204-1
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Definition

The suction anchors are usually cylindrical units with large diameter (D of 3–8 m), open at the bottom and closed at the top, and generally with an aspect ratio (length to diameter, L/D) of two to six. The anchoring line is connected to the padeye attached to the side of the caisson at a depth down the mudline. The capacity of suction anchor is therefore the ability to resist the permanent and environmental loading from the anchor line during the operational conditions.

Failure Mechanism

The capacity of suction anchor is highly dependence of soil plasticity flow around the foundation. Many influential factors, especially for the complicated boundary conditions, such as anchor geometry, type of mooring system (loading direction and magnitude), installation performance, soil conditions, time after installation, durations of applied sustained loading, cycles of loading, etc., should be explicitly accounted for to investigate the failure mechanism.

In the following subsections, the typical failure mechanisms under uniaxial (including mainly vertical and horizontal) and inclined loading are discussed.

Under Vertical Uplift Loading

Under conditions of pure uplift loading, three failure mechanisms have been identified by a number of experimental or numerical analyses: three important components of failure modes including reverse end bearing failure, sliding failure and tensile failure, as illustrated in Fig. 1.
Fig. 1

Failure mechanisms for vertical pullout resistance (a) reverse end bearing, under undrained conditions; (b) caisson pullout, under drained conditions; (c) caisson and plug pullout, under partially drained conditions

Reverse End Bearing Failure

Reverse end bearing failure can be relied upon for a suction anchor if the top cap is sealed and the soil response is undrained, typically when the anchor is subjected to short-term (transient) loading. Passive suction is generated in the anchor, so that the soil plug is pulled out together with the caisson. The failure mode at the bottom during uplift was similar to that during compression and was referred to as “a reverse end bearing mechanism.”

Sliding Failure

When suction anchor is subjected to a sustained (drained) loading or the anchor lid is vented, reverse end bearing cannot be generated, and the failure occurs along the internal and external skirt.

Tensile Failure

When partially drained condition prevails, such as when the suction anchor is pulled out at intermediate rates, passive suction may be generated partially, so that the caisson and the internal soil plug tends to detach from the soil below the caisson.

Under Optimal Horizontal Loading

The suction anchors under pure horizontal loading is rarely found offshore, while drag angle at the padeye is typically at a small angle of 10°–20° in the catenary mooring system. Meanwhile, the padeye position is often optimized in design, allowing the anchor to slide horizontally with minimal rotation.

As shown in Fig. 2, the corresponding failure mechanism for lateral capacity of the suction anchor is considered to be similar to those for laterally load piles (Murff and Hamilton 1993) that a passive/active conical wedge extending from the edge of the caisson at shallow depth to the seabed dominating the soil mechanism; and a confined flow region is assumed below the wedge with the center of rotation locating below the anchor tip (Martin and Randolph 2006). For caissons loaded above the optimal depth, an external scoop mechanism can replace the flow region, with the center of rotation locating within the anchor.
Fig. 2

Soil failure mechanism for suction anchor (a) conical wedge and flow region and (b) external base rotational scoop. (After Randolph and House 2001)

Under Inclined Loading

The failure mechanism under inclined loading depends on the position of the centerline intercept for the applied load (varying with padeye position and loading angle), the length of the suction anchor, and the soil properties (Supachawarote 2006). The impact of tensile crack on the failure mechanism is examined by Fu et al. (2020).

The maximum holding capacity is obtained if the chain is attached at a depth where the anchor moves largely in horizontal translation with minimal rotation. This is called the “optimal padeye depth.”

When the centerline intercept is located above the optimal padeye position, the anchor rotates forward, and combinations of wedge flow around and full circular rotational failure were observed. For the short anchor, the wedge and rotational flow mechanisms were observed, without the flow around region. For the longer anchor, the flow around region increasingly dominates.

While when the centerline intercept is located below the optimal padeye position, a backward rotation was observed, and the rotational flow failure extends to the mudline. And with the centerline loading point moves deeper, the full circular rotational failure eventually moves below the soil surface, which is evident for anchor with increasing length. For long anchor, the circular rotational flow failure is only observed below the tip of the anchors; and only combination of the wedge and flow around mechanisms was observed within the anchor length.

Design Method of Evaluating Capacity

The load capacity of a suction caisson anchor, assuming a sealed cap, is derived from bearing resistance between the soil and the projected area of the caisson (on a vertical plane for horizontal resistance and the cross-sectional area for reverse end bearing), aided by frictional resistance along the outside of the caisson shaft. Reverse end bearing relies on passive suctions developed within the soil plug, and so consideration needs to be given to the time over which these can be sustained.

Vertical Uplift Capacity

Equation

The vertical pullout capacity is equal to one of the following depending on the drainage conditions (essentially whether the passive suction is sustained during loading):
  1. (a)

    For undrained conditions, with a nondimensional diameter Tk (= cv/vD, where cv is the consolidation coefficient, v is the loading velocity, D is the diameter of the anchor) recommended to be less than 0.002 (Deng and Carter 2000), the passive suction is thought to be sustained during loading

     
$$ {V}_{\mathrm{ult}}={W}^{\prime }+{A}_{se}{\alpha}_e{\overline{s}}_{u(t)}+{N}_c{s}_u{A}_e $$
(1)
  1. (b)

    For drained condition, (Tk > 0.6), the passive suction is not generated during loading

     
$$ {V}_{\mathrm{ult}}={W}^{\prime }+{A}_{se}{\alpha}_e{\overline{s}}_{u(t)}+{A}_{si}{\alpha}_i{\overline{s}}_{u(t)} $$
(2)
  1. (c)

    For partially drained condition (0.002 < Tk < 0.6), passive suction is partially generated

     
$$ {V}_{\mathrm{ult}}={W}^{\prime }+{A}_{se}{\alpha}_e{\overline{s}}_{u(t)}+{W}_{plug}^{\prime } $$
(3)
where Ase and Asi are the external and internal shaft surface area, Ae is the external cross-sectional area, αe and αi are the coefficient of external and internal shaft friction, Nc is the reverse end bearing factor, su is the undrained shear strength at tip level, \( {\overline{s}}_{u(t)} \)is the average undrained shear strength over penetrated depth at time t after installation, Wplug is the effective weight of the soil plug, and W′ is the submerged caisson weight.

Evaluation of Parameters α and Nc

The soil shear strength around the anchor is thought to be remold due to disturbance during skirt penetration during self-weight penetration. And this strength recovers later with time following installation, due to a combination of thixotropy and consolidation (Andersen et al. 2005), which however remains below the intact undrained shear strength. This process is referred as “setup” and has been expressed conveniently with αsu, with α < 1. A base value of α of 0.65 was proposed for suction caisson design (Anderson and Jostad 2004), subjected to their specific soil properties. The shaft friction depends on the anchor surface roughness, soil type, and over-consolidation ratio. And Chen et al. (2009) suggests method of installation (referring jacking or suction installation) leads to negligible difference in α based on the centrifuge and LDFE numerical studies. And their results also suggest α of 15–20% lower than those recommended by the American Petroleum Institute (API) for driven piles in normally consolidated clays. It should also be noted that lower values of friction coefficient for internal shaft friction compared with external friction were reported in the model tests conducted on double-walled caisson (Jeanjean 2006).

The reverse end bearing Nc is a function of soil property and the aspect ratio (L/D) of the anchor. For an anchor with L/D of 2 in the over-consolidated kaolin clay, Fuglsang and Steensen-Bach (1991) reported the centrifuge and laboratory tests, suggesting the reverse bearing capacity factors varying between 6.5 and 8.5. This is lower than the theoretical lower bound results of 9.2 for the caisson with L/D of 2 (Martin 2001). Other centrifuge tests for the anchor with the same aspect ratio (Clukey and Morrison 1993) suggested that reverse end bearing is around 11. This high magnitude of Nc may be related to the vane shear strength adopted for the interpretation, which leads to an overestimation of 25% (Watson et al. 2000). Considering the loading mode (monotonic transient loading, cyclic loading, and long-term sustained loading), the reverse end bearing falls in a wider range of 9.1–14.6 (Randolph and House 2001). And a conservative Nc value of 9 is suggested to be taken, due to the strain-softening nature of the response as the caisson is extracted. A bearing capacity factor of Nc = 9 irrespective of the aspect ratio for suction anchor is also suggested by El-Gharbawy and Olson based on the results of the laboratory tests in kaolin clay. And by contrast, an Nc value of 12 is found to be mobilized at large displacements (Jeanjean 2006), although values of around 9 were mobilized at the displacement where peak external shaft friction was achieved. Hence it is rational to take Nc as of 9 is appropriate.

Optimal Horizontal Capacity

Equation

Assuming the padeye is positioned at the optimum level and pure translation failure occurs, the maximum horizontal resistance is
$$ {H}_{\mathrm{max}}={LD}_e{N}_p{\overline{s}}_u $$
(4)
where L is the embedded length of caisson, De is the external diameter of the caisson, Np is the lateral bearing capacity factor, and \( {\overline{s}}_u \)is the average undrained shear strength over penetrated depth.

Evaluation of Parameters Np

Zhang et al. (2016) recommended the following approach to calculate the lateral bearing capacity factor
$$ {N}_{\mathrm{p}0}={N}_1-\left({N}_1-{N}_2\right){\left[1-{\left(\frac{z/D}{Z}\right)}^{0.6}\right]}^{1.35}-\left(1-\alpha \right)\le {N}_{\mathrm{p}\mathrm{d}} $$
(5)
where
$$ {N}_1=11.94 $$
(6)
$$ {N}_2=3.22 $$
(7)
$$ Z=16.8-2.3\lg \lambda \ge 14.5 $$
(8)
$$ \lambda =\frac{s_{um}}{kD} $$
(9)
$$ {N}_{\mathrm{pd}}=9.14+2.8\alpha $$
(10)

In Eq. 8, Z stands for the normalized depth (z/D) at which the soil failure mechanism transits from the wedge failure to the localized flow around failure for a fully rough soil-pile interface in idealized weightless soil. The value of Z is related to the normalized strength homogeneity parameter λ. Npd stands for the limiting bearing capacity factor mobilized in a localized flow-around mechanism, expressed as a function of the interface roughness factor α (0 ≤ α ≤ 1), according to the plasticity solution.

Inclined Load Capacity

In the presence of both vertical and horizontal loading, a reduction occurs in the pure vertical and horizontal capacities, as the caisson is simultaneously displaced vertically and laterally (or rotated). Clukey et al. (2003) suggested for the loading angles applied by the catenary mooring line system, which are generally less than 20° from the horizontal, the caisson capacity is dominated by the horizontal capacity, and the capacity may be estimated by the capacity under pure translation divided by the cosine of the loading angle at the padeye. Conversely for high loading angles from the horizontal applied by taut and semi-taut mooring systems (generally in excess of 30°), the caisson capacity is essentially governed by the vertical capacity of the caisson and maybe approximately taken as the vertical capacity divided by the sine of the loading angle.

For the more general cases, the interaction between vertical and horizontal loading may be conveniently modelled as a failure envelope in the combined vertical-horizontal load space, which was found to vary with the aspect ratio (L/D), location, and direction of the applied load. The shape of the failure envelopes for suction anchor may be modelled by an elliptical relationship
$$ {\left(\frac{H}{H_{\mathrm{ult}}}\right)}^a+{\left(\frac{V}{V_{\mathrm{ult}}}\right)}^b=1 $$
(8)
where Hult and Vult are the uniaxial horizontal and vertical capacities, respectively, and the exponents a and b vary with caisson aspect ratio L/D according to Supachawarote (2006).
$$ a=\frac{L}{D}+0.5 $$
(9)
$$ b=-\frac{L}{3D}+4.5 $$
(10)
The sensitivity of caisson capacity to the changes in padeye position or loading angle was also investigated using the three-dimensional finite element analysis, limit equilibrium solutions upper bound limit analyses, profiles of lateral resistance based on Murff and Hamilton (1993), and centrifuge modelling (Clukey et al. 2003). It was found that the variation of centerline intercept for the applied load (varying with padeye position and loading angle) results in the large discrepancy of the capacity. As shown in Fig. 3, the centerline loading depth around 0.7 gives the largest capacity. With the shift of the padeye position and loading angle, the capacity reduces gradually within the depth of ±0.15 L from the optimal centerline location and reduces sharply beyond this range.
Fig. 3

Effect of padeye depth on capacity. (After Andersen et al. 2005)

Consideration in Design

The Effect of Crack on the Suction Anchor Capacity

The inclined load capacity of a suction anchor will also depend on whether a crack develops along the trailing edge of the caisson. A crack is normally formed for suction anchors with a long aspect ratio in over-consolidated soils. The finite element results (Fu et al. 2020) show a reduction of up to 50% for any loading angle (the reduction is mainly found at the optimal padeye depth and maintained for loading angles of up to about 45° for L/D = 1.5, 2, and 3 in homogeneous soil).

Considering Site Conditions

Some changes in site conditions should be explicitly accounted for in the design of geotechnical capacity of suction anchor. The soil scour around a suction foundation is an important scenario in design, where the soil migration around foundation might occur under the wave and current loading. The volume of soil mobilized by foundation is therefore changed, resulting in a significant decrease in capacity. In some cases, the shallow gas accumulation inside a suction foundation also should be given attention (Gylland and de Vries 2008). Additionally, the trenching is another issue and most likely to occur particularly when using semi-taut to taut mooring configurations with caisson employed in the soft deposits. Considerable motion of ground chain could lead to the remolded softening and erosion of soil particles, forming a curved trench in the vicinity of foundation. O’Neill et al. (2018) presented a method to interpret and estimate the primary mechanism of trenching development.

Considering Cyclic Loading Effect

Offshore foundations are often cyclically loaded in both calm sea condition and extreme events (i.e., storm or hurricane event). For suction anchor, the cyclic wind, wave, current, and structural loadings are transferred to the foundation by mooring lines. It gives rise to pore pressure accumulation for soil around foundation, which is one of the principle concerns in design of foundation’s capacity. Regarding clay, the strain softening may occur as the pore pressure accumulates, and both soil stiffness and strength can significantly reduce. Regarding sands, the earlier densification may occur during cyclic loading, then pore pressure starts accumulating, and finally a state of liquefaction may be present with the effective stress of soil almost being zero. Therefore, the continual buildup of excess pore pressure in soil under cyclic loading potentially results in a significant decrease in capacity of foundation.

Based on laboratory cyclic triaxial (TA) and direct simple shear (DSS) tests, Andersen (2015) presented a method to assess the accumulation of pore pressure, which is dependence of cyclic shear ratio τcy/σ′vc. τcy is the cyclic shear stress (kPa), and σ′vc is the vertical consolidation pressure (kPa). The corresponding cyclic shear strength τcy,f at failure was indicated using a strain contour diagram, with the equivalent number of cycles to failure captured at the permanent cyclic shear strain of 15%. Besides, many soil models based on the critical soil mechanics theory were numerically developed in the last three decades, to describe and interpret the cyclic behavior of soil. However, no constitutive model has been developed to date, enabling to represent all the key characteristics of cyclic soil response, as many factors may influence the cyclic performance of soil, i.e., cyclic stress level, loading frequency, over-consolidation ratio, static pre-shearing, etc. In practice, the alternative option available for engineers is choosing a soil material factor (ISO 2016) or considering the uncertainty by a partial factor of capacity (ISO 2016), when the cyclic loading effect is necessary to be considered.

Considering Reconsolidation Effect

For suction anchor, an allowable duration is often found after its installation, e.g., 3 to 6 months. This gives an opportunity for consolidation under self-weight prior to commencing operations. During consolidation, the excess pore pressures generated by the self-weight preloading dissipate reducing the void ratio of soil and enhancing the soil strength, hence increasing the foundation capacity. It is notable that the dissipation of pore pressure also occurs during the cyclic loading condition, as the pore pressures in fact generated during a cyclic duration and will dissipate in the subsequent cycles. The appropriate consideration of reconsolidation in service design is necessary to develop a reliable and economical method.

From some centrifuge model tests (Bienen et al. 2010; Fu et al. 2015), the significant increase in bearing capacity of a preloaded foundation in clay was found, for instance, a gain in bearing capacity of 30% after 10% of consolidation and a doubling after 80% of consolidation under a vertical preload of 50% of the ultimate vertical load.

Notation

a Parameter depending on L/D

Ae External cross-sectional area

Ase External shaft surface area

Asi Internal shaft surface area

b Parameter depending on L/D

cv Consolidation coefficient

D Diameter of suction anchor

De External diameter of suction anchor

Hmax Optimal horizontal capacity

Hult Uniaxial horizontal capacity

L Length of suction anchor

Nc Reverse end bearing factor

Np Lateral bearing capacity factor

su Undrained shear strength at tip level

\( {\overline{s}}_u \) Average undrained shear strength over penetrated depth

\( {\overline{s}}_{u(t)} \) Average undrained shear strength over penetrated depth at time t after installation

Tk Nondimensional diameter

W′ Submerged weight of suction anchor

W′plug Submerged weight of soil plug

v Loading velocity

Vult Vertical pullout capacity

zcl Centerline loading depth

α Coefficient of shaft friction

α Parameter depending on L/D

αe Coefficient of external shaft friction

αi Coefficient of internal shaft friction

β Parameter depending on L/D

τcy Cyclic shear stress

τcy,f Cyclic shear stress at failure

σ′vc Vertical consolidation pressure

References

  1. Anderson KH, Jostad HP (2004) Shear strength along inside of suction anchors skirt wall in clay. In: Proceedings of. annual offshore technology conference, Houston, OTC 16844Google Scholar
  2. Andersen KH, Murff JD, Randolph MF, Cluckey EC, Erbrich CT, Jostad HJ, Hansen B, Aubeny CP, Sharma P, Supachawarote C (2005) Suction anchors for deepwater applications. Frontiers in offshore geotechnics: ISFOG - Gourvenec & Cassidy. London: Taylor & Francis Group 0 415 39063 XGoogle Scholar
  3. Bienen B, Gaudin C, Cassidy MJ (2010) Centrifuge study of the bearing capacity increase of a shallow footing due to preloading. In: Proceedings of the 7th Internationale Conference on Physical Modelling in Geotechnics (ICPMG). 2:1019–1024Google Scholar
  4. Chen W, Zhou H, Randolph MF (2009) Effect of installation methods on external shaft friction of caissons in soft clay. J Geotech Geoenv Eng ASCE 135(5):605–615CrossRefGoogle Scholar
  5. Clukey EC, Morrison J (1993) A centrifuge and analytical study to evaluate suction caissons for TLP applications in the Gulf of Mexico. ASCE Geotechnical Special Publication 38:141–156Google Scholar
  6. Clukey EC, Aubeny CP, Murff JD (2003) Comparison of analytical and centrifuge model tests for suction caissons subjected to combined loads. In: Proceedings of the international conference on offshore mechanics and arctic engineering, OMAE2003–37503Google Scholar
  7. Deng W, Carter JP (2000) A theoretical study of the vertical uplift capacity of suction caissons. In: Proceedings of the 10th international offshore and polar engineering conference, pp 342–349Google Scholar
  8. Fu D, Gaudin C, Tian Y, Bienen B, Cassidy MJ (2015) Effects of preloading with consolidation on undrained bearing capacity of skirted circular footings. Géotechnique 65(3):231–246CrossRefGoogle Scholar
  9. Fu D, Zhang Y, Yan Y, Jostad HP (2020) Effects of tension gap on the holding capacity of suction anchors. Mar Struct 69(2020) 102679:1–14.  https://doi.org/10.1016/j.marstruc.2019.102679
  10. Fuglsang LD, Steensen-Bach JO (1991) Breakout resistance of suction piles in clay. In: Proceedings of international conference: Centrifuge 91. A.A. Balkema, Rotterdam, The Netherlands, pp 153–159Google Scholar
  11. Gylland AS, de Vries MH (2008) The effect of gas blow-out on shallow offshore foundations. In: Proceedings of the Second British Geotechnical Association International conference on foundations (ICOF 2008), volume 1: Piles, Excavations and Offshore Foundations, pp 885–896Google Scholar
  12. ISO (2016) ISO 19901-4:2016(en). Petroleum and natural gas industries specific requirements for offshore structures – part 4: geotechnical and foundation design considerations. International Standards Organisation, GenevaGoogle Scholar
  13. Jeanjean P (2006) Set-up characteristics of suction anchors for soft Gulf of Mexico clays: experience from field installation and retrieval. In: Proceedings of the. Annual Offshore Technical Conference, Houston, Texas, Paper OTC 18005Google Scholar
  14. Martin, C.M., (2001). Vertical bearing capacity of skirted circular foundations on Tresca soil. In: Proceedings of 15th international conference on soil mechanics and geotechnical engineering, 1, pp 743–746Google Scholar
  15. Martin CM, Randolph MF (2006) Upper bound analysis of lateral pile capacity in cohesive soil. Géotechnique 56(2):123–132CrossRefGoogle Scholar
  16. Murff JD, Hamilton JM (1993) P-ultimate of undrained analysis of laterally loaded piles. J Geotech Eng ASCE 119(1):91–107CrossRefGoogle Scholar
  17. O’Neil M, Erbrich C, McNamara A (2018) Prediction of seabed trench formation induced by anchor chain motions. In: Offshore technology conference, OTC Paper No. OTC-29068-MSGoogle Scholar
  18. Randolph MF, House AR (2001) Analysis of suction caisson capacity in clay. In: Proceedings of the annual offshore technology conference, Houston, OTC 14236Google Scholar
  19. Supachawarote C (2006) Inclined load capacity of suction caisson in clay. PhD thesis. The university of Western AustraliaGoogle Scholar
  20. Watson PG, Randolph MF, Bransby MF (2000) Combined lateral and vertical loading of caisson foundations. In: Proceedings annual offshore technology conference, Houston, OTC 12195, pp 797–808Google Scholar
  21. Zhang Y, Andersen KH, Tedesco G (2016) Ultimate bearing capacity of laterally loaded piles in clay – some practical considerations. Mar Struct 50(2016):260–275CrossRefGoogle Scholar

Copyright information

© Springer Nature Singapore Pte Ltd. 2020

Authors and Affiliations

  1. 1.State Key Laboratory of Hydraulic Engineering Simulation and SafetyTianjin UniversityTianjinChina

Section editors and affiliations

  • Yinghui Tian
    • 1
    • 2
  1. 1.Department of Infrastructure EngineeringThe University of MelbourneParkvilleAustralia
  2. 2.State Key Laboratory of Hydraulic Engineering Simulation and SafetyTianjin UniversityTianjinChina