Technical Committee 209 Offshore Geotechnics. Comité technique 209 Géotechnique marine

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2 Technical Committee 29 Offshore Geotechnics Comité technique 29 Géotechnique marine

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4 General Report of TC29 Offshore Geotechnics Rapport Général général du TC29 Géotechnique Offshore marine Jewell R.A. Fugro GeoConsulting ABSTRACT: This general report introduces the discussion session organized by ISSMGE Technical Committee 29 (TC29) Offshore Geotechnics. The main topics include offshore wind projects, pipelines and seabed structures, seabed soils, coastal and nearshore work, and pile foundations. RÉSUMÉ : Ce rapport général introduit la séance de discussion organisée par le Comité Technique 29 (TC29) Géotechnique Offshore de la SIMSG. Les principaux thèmes abordés sont les projets d éoliennes offshore, les pipelines et structures sous-marines, les sols sous-marins, les travaux côtiers et nearshore et les fondations sur pieux KEYWORDS: offshore, caisson, piles, pipes, cyclic load, stability diagram, lateral load, tests, numerical analysis, bearing capacity 1 INTRODUCTION. The organizers of the 18th International Conference Challenges and Innovations in Geotechnics have implemented important changes to the conference format. One is the inclusion of Offshore Geotechnics at this main ISSMGE forum. Second is the focus given to the technical committees. This general report covers the session organized by the ISSMGE Technical Committee 29 (TC29) Offshore Geotechnics chaired by Philippe Jeanjean. Participation by TC29 at this 18 th conference includes the 2 nd ISSMGE McClelland Lecture by Mark Randolph, this discussion session and a workshop on recent research and development on piles under cyclic loading. The main difference in offshore geotechnics arises from the conditions and environment offshore. There is a stark contrast in access for site investigation, soil sampling, field testing, installation and observation. Activities offshore often require new tools. Soft soil conditions at seabed level are encountered in deepwater, frequently with high carbonate content, unusual mineralogy or biogenic activity. Combined and cyclic loading usually dominate design, whether caused by waves and currents acting on structures or by repeated expansion and contraction of pipelines on the seabed. Large displacement is a feature of the installation and operation of seabed pipelines. Many of the above features of offshore geotechnics are discussed in papers to this session. This general report has been organized into five main subject areas: Offshore Wind; Pipelines and Seabed Structures; Seabed Soils; Coastal and Near Shore work; Pile Foundations. Since papers on the cyclic loading of piles will be presented and discussed at the TC29 workshop, these are highlighted in this general report but will not be presented during the discussion session. The main focus for presentations will be Offshore Wind and large displacement as encountered with Offshore Pipelines. Only a limited selection of papers will be presented at the discussion session. All the papers are in the proceedings and many will be presented at the poster session. Participants are strongly encouraged to attend the TC29 workshop where the cyclic behavior and design of piles will be presented and debated based around papers to this session. 2 OFFSHORE WIND. 2.1 Site investigation Project development, engineering design and project construction are three main phases for offshore wind farms. A major challenge is to minimize geotechnical risk for foundation engineering. In current practice, geotechnical risk is addressed mainly during the engineering design phase P2 in Table 1. Ben- Hazzine and Griffiths (213) suggest that risk management may be improved through more extensive geophysical survey and preliminary site investigation during the project development phase P1. The authors highlight various sources of geotechnical risk such as inherent soil variability, measurement errors and transformation errors caused by simple empirical interpretation of data. Table 1. Timing of Geotechnical Work for Offshore Windfarms (Ben-Hazzine and Griffiths, 213) Muir Wood and Knight (213) use experience from 15 offshore wind farm projects in the UK to illustrate the manner in which geotechnical risk was managed and to define categories of poor, mediocre and good practice (or vice versa). Poor practice includes appointment of the foundation design team after site investigation is completed, insufficient planning and poor interpretation of geophysical and geotechnical surveys. Mediocre practice often involves incorrect scope for geophysical and geotechnical surveys causing extra cost and increased risk. On projects with good practice the foundation design team was appointed at the start of the project and specified the site investigation work. The ground models were 2295

5 Proceedings of the 18 th International Conference on Soil Mechanics and Geotechnical Engineering, Paris 213 developed and refined over multiple phases of geotechnical and geophysical investigation. The authors conclude that a formal approach to risk management with staged investigations and early appointment of the foundation design team is the best practice. 2.2 Wind turbine performance A wind energy project involves several disciplines within civil engineering. The complex interaction between the turbine and the supporting structure governs the dynamic behaviour. Analysis is required of the structural modal parameters that influence fatigue of a wind turbine. Indeed, continuous monitoring of modal frequencies and damping ratios during the operational life of a wind turbine can provide early warning of the onset of structural damage. Damgaard et al. (213) assess instrumentation data from 3 offshore wind turbines in the North Sea. The measured modal frequency and damping ratio are found to vary with time. However, the observed magnitude and pattern of change with time might result from scour erosion or backfilling around the monopile structure. This highlights that the dynamic interaction between an operating turbine and the supporting structure may be affected by changes with time of the local seabed level, something that could be reduced by scour protection. 2.3 Foundation systems The power capacity of a wind turbine, typically in the range 2MW to 5MW, but increasing toward 1MW in future, determines the required height above mean sea level and the maximum horizontal and vertical loads to be supported. Deeper water increases the moment arm for both wind and wave loading and often signifies larger design storms. There are many novel features to the required foundation engineering as discussed in papers to this session, Table 2. Table 2. Papers on Foundation Systems for Offshore Wind. SINGLE LOAD CYCLIC LOAD Shallow Arroyo et al A Foundation Monopile Roesen et al E.1g; A Caisson Kim et al E.c; A Versteele et al A Hybrid (pile/caisson) Arshi et al E.1g; A E = Experimental (E.1g lab floor; E.c centrifuge); A = Analysis In brief résumé, shallow foundations and monopiles generally provide efficient support for wind turbines. However, these foundations are less effective when the moment load to be resisted increases due to larger turbines and/or deeper water. One approach is to improve the capacity by combining a monopile and shallow foundation, Figure 1. Alternatively, a caisson can support combined vertical, horizontal and moment loading at seabed level. Three or four caissons may be combined to support a structure where greater load carrying capacity is required, Figure 2. The issues for foundation engineering discussed in the papers include: (1) the impact of the loading path and loading direction on safety factor and use of a single failure envelope; (2) improved performance by combining a shallow foundation and monopile; (3) the incremental displacement of cyclically loaded monopiles; and, (4) assessment of simultaneous pore water pressure generation and dissipation for caissons under storm loading. (Arshi et al, 213) Figure 1. Hybrid Monopile and Shallow Foundation. (Versteele et al, 213) Figure 2. Monopod and Multipod Foundations Bearing capacity analysis The assessment of bearing capacity for offshore wind turbine foundations differs from onshore practice in several respects. The issues include: (a) separate correction factors in analysis for shape, depth, load inclination and eccentricity that are cumbersome and prone to calibration error; (b) use of separate partial factors on loads and resistances as in DNV-OS-J11 when the difference between favourable and detrimental effects can be subtle; and (c) simultaneous application of two major horizontal loads from wind and wave acting in separate directions. Arroyo et al highlight these issues and question the suitability of the conventional design approach for offshore wind foundations (Arroyo et al 213). A more satisfactory framework for capacity checks would be through failure-envelopes as detailed in the paper. Arroyo et al examine a synthetic design example to illustrate their case using the geometry of a Thornton Bank GBS and a set of derived loading parameters, Figure 3. The complex interaction between horizontal and moment loading, and the impact of different directions for wind and wave loading are illustrated by the authors using Figure 4 and the results in Table 3. It is conventional to increase both detrimental loads at the same time, so that the load increment causing the limit to be reached is in the same direction as the reference combined load. The authors consider the cases where only wind or wave loading is increased. Such analysis might be used to assess the impact of any error in the assessment of those loads. Because of the greater influence on moment of wind loading, the analysis shows that an underestimate 21% (in this case) in the wind load would be sufficient to cause failure compared with an underestimate 5% for the wave load, Table 3. Arroyo et al (213) conclude that failure envelopes offer a powerful framework for analysis of shallow foundation capacity; the approach is particularly well suited for offshore wind structures that require refined design in the face of considerable uncertainty. 2296

6 Technical Committee 29 / Comité technique 29 the caisson. Centrifuge tests are currently underway to define better the benefit of caisson versus simple shallow foundation in such hybrid foundation systems. (Arshi et al 213) Figure 5. Moment resistance chart for hybrid foundations Lateral displacement due to cyclic loading Figure 3. Thornton Bank GBS (after Peire et al 29). (Arroyo et al 213) Figure 4. Incremental loading paths to failure. The focus above is the limit resistance of offshore wind tower foundations subject to a single application of combined loading. In practice, the structures are subject to several episodes of extreme loading caused by major storms and a great number of cycles of low amplitude loading from normal wind and wave conditions. The latter source of repeated loading may cause fatigue or serviceability problems (Roesen et al. 213). The authors report a series of 1g laboratory tests on monopiles in sand subject to one-way cyclic loading over more than 5, cycles. One limitation in these 1g tests is a more rigid pile compared with a typical prototype, but the trend of results should be similar. The cyclic loading is described by two non-dimensional ratios: the maximum moment compared with the static moment capacity ζ b = M max /M static in the range 1 > ζ b > ; and the ratio of the minimum to the maximum moment ζ c = M min /M max which has a value 1 for a static test, for one-way loading (the case examined by Roesen et al) and -1 for two way cyclic loading. The pile displacement is measured by the rotation θ at the soil surface. The results of a static load test and the measured displacement in one-way cyclic load tests (ζ c = ) with load intensity ζ b =.2 to.4 are shown on Figure 6. Table 3. Results of analysis on incremental load to failure Hybrid foundations When considered as a monopile design, the addition of a shallow footing at seabed level may be thought of as adding fixity to the monopile head thereby generating improved resistance and stiffness to lateral loading, Figure 1. A simplified design analysis would assess the limiting moment capacity of the shallow foundation acting alone and include the equal and opposite moment resistance to the analysis of the monopile. The shallow foundation not only increases load carrying capacity but also reduces the bending moment supported by the monopile, by about 25% in the example cited by Arshi et al (213). Experimental modelling of these hybrid systems at 1g and in the centrifuge are described together with numerical and analytical work. The significance of the geometric ratio of footing to pile diameter, and pile length to pile diameter is demonstrated and a form for design charts proposed, Figure 5. Some tests on caisson/monopile combinations are noted and indicate additional benefit due to the lateral load resistance of (Roesen et al 213) Figure 6. Static and cyclic one-way loading tests on model monopiles. The incremental rotation due to one-way cyclic loading may be quantified with respect to the rotation caused by the first, single loading Δθ N = θ N θ 1. Tests with different loading intensity may then be compared through the non-dimensional form Δθ N /θ 1 as shown on Figure 7. The test data are compared with a simple power law (Δθ N / θ 1 ) = an b, where a and b are empirical constants found from testing. The power law seems to provide a reasonable asymptotic limit for the data after about 1 or more load repetitions, Figure 7. The value of the constant a is found to vary almost linearly with the applied load magnitude ζ b. The slope of accumulating displacement with repeated loading, the constant b, appears to be a function of the combination of 2297

7 Proceedings of the 18 th International Conference on Soil Mechanics and Geotechnical Engineering, Paris 213 monopile and sand parameters, presumably including the installation process. The rate of accumulation appears not to depend markedly on the applied load magnitude. The analysis in 6.2 suggests that the data might also fit with a natural logarithm through to low numbers of cycles (Δθ N / θ 1 ) = c.ln(n). process is repeated for the next package of cyclic loading, and so on. The analysis for pore pressure generation uses two relations for the sand material. First the measured cyclic shear strength versus number of cycles to liquefaction, N l, from cyclic laboratory tests; second an empirical formula linking pore water pressure generation to number of load cycles, Figure 9. Liquefaction occurs at u/u max = 1 when N/N l = 1. (Roesen et al 213) Figure 7. Monopile rotation versus number of cycles Caisson foundations A centrifuge test of a caisson in sand is reported by Kim et al. (213). The caisson response to single combined load to failure is measured and numerical analysis applied. The test details are provided elsewhere and it is not clear whether the 1/7 th scale caisson was installed by suction during the centrifuge test or before testing. Soil material from the proposed site is used to model a planned prototype caisson foundation. The measured response of the modelled prototype 15.5m diameter 1.5m long caisson is shown on Figure 8 in terms of applied moment versus rotation. The authors report a parametric analysis using FLAC to show the significant influence of the assumed elastic modulus and cohesive strength parameters assumed for the soil. (Versteele et al 213) Figure 9. Generation of excess pore water pressure. The results of a 3D analysis for a caisson foundation are reported to illustrate the method, Figure 1. The 2m diameter by 1m length caisson is subjected to a 6 hour design storm of 216 waves. These are split for analysis into 5 individual load packages. The direction of wind and wave loading is assumed to be aligned. The results illustrate the asymmetric nature of pore water pressure generation that has potential consequences for possible differential settlement and tilting of the caisson. (Kim et al 213) Figure 8. Measured moment-rotation for prototype caisson. The offshore design standard DNV-OS-J11 (DNV, 211) requires structures to resist a 5 year design storm considering both peak loads and the entire history of cyclic loading. It is this latter requirement that is tackled by Versteele et al (213) for the case of caisson foundations in sand. Because a full analysis of cyclic loading of caissons in sand is not practically feasible with current numerical methods, the authors develop an analysis to provide insight into the competing processes of excess pore water pressure generation and dissipation during the design storm. The analysis breaks the storm into several packages of cyclic loading (magnitude, number of cycles and time). The excess pore water pressure generated at each point in the soil by the package of cyclic loading is computed analytically and input into the numerical analysis. The dissipation and redistribution of pore water pressure during the time period is computed numerically. The (Versteele et al 213) Figure 1. Example of excess pore water pressure below a caisson. Versteele et al (213) conclude that the model is useful in predicting areas beneath the caisson prone to the development of excess pore pressure. However, the analysis does not predict liquefaction behaviour or compute settlement, nor does it allow for load redistribution in the caisson due to the changing effective resistance in the soil during the design storm. There is further development work to be done. 3 PIPELINES AND LARGE DISPLACEMENT. A challenging feature for offshore pipelines is the large displacement that can occur during installation and service. Large displacement is particularly extreme for laying pipe on a soft seabed. Large displacement also results from multiple cycles of heat expansion and contraction of the operating pipeline. This requires engineering design to avoid localized 2298

8 Technical Committee 29 / Comité technique 29 pipe distortion and over-stress irrespective of the sea bed soil type. A less discussed cause of large displacement is where a pipeline crosses a seismic fault; here it is movement of the ground with respect to the pipe that causes gross distortion. Fault crossings occur both onshore and offshore. The geotechnical analysis for large displacement requires suitable tools and numerical models and both have developed significantly in recent years. A variety of large displacement numerical methods with 3D capability are commercially available for design purposes. Similarly, constitutive models for soft clay that account for competing strain rate and strain softening effects, and competing pore pressure generation and dissipation, are available for designers. Specific numerical elements to model the large displacement interaction between a pipe and the surrounding soil are also currently under development for practical application in design (SAFEBUCK JIP). However, the constitutive and numerical modeling for large displacement in dense sand is less well advanced. Pipe buckling and pipe walking is usually assumed to occur between fixed seabed structures. There may be scope for permitting the seabed structures to move horizontally to help accommodate axial pipe displacement. Several of these topics are described in four papers to the discussion session. 3.1 Dynamic embedment of offshore pipelines Dutta et al (213) examine pipe laying and dynamic embedment using Coupled Eulerian Legrangian (CEL) methods available in ABAQUS software. Progressive degradation of undrained shear strength with plastic shear strain is included using the model of Einav and Randolph (25). Similar analysis by Wang et al (21) used remeshing and small strain (RITSS analysis). The simplified problem is illustrated in Figure 11. A pipe is penetrated monotonically into a soft clay sea bed under selfweight (submerged weight of pipe). The pipe is then cycled laterally by a displacement u/d = ±.5, in the x direction of Figure 11, under constant self-weight vertical load. This causes additional pipe penetration. incorporate initial overloading of the pipe, but the forty cycles of lateral loading resulted in a dynamic embedment factor of the order 4 to 5, within the range often assumed in practice. (Dutta et al 213) Figure 12. Static and dynamic pipe penetration of a seabed. As shown previously by Wang et al (21), the analysis provides insight into the size of the zone of highly sheared and softened soil around the pipe and the shape of the berms formed by pipe penetration. The results in Figure 13 are for the heavier pipe and show dynamic pipe penetration and monotonic pipe penetration to the same depth (increased vertical load). The comparison is striking. Dynamic embedment causes more extensive plastic strain softening in the soil, coloured red, and wider and flatter berms than generated by monotonic pipe penetration. The latter could be important for the analysis of initial lateral breakout of the pipe. Dynamic embedment affects the magnitude of pipe penetration, the zone of soil remoulding and the shape of the berms formed. (Dutta et al 213) Figure 13. Dynamic and monotonic pipe penetration of a soft seabed. (Dutta et al 213) Figure 11. Pipe penetration of a seabed. The analysis uses the same dimensions, soil parameters and loading sequence as the first stage of two pairs of centrifuge tests by Cheuk and White (28) on a light and heavier pipe. The progressive pipe penetration and magnitude of horizontal resistance caused by cyclic lateral displacement is computed. The penetration of the pipe with cycles of lateral displacement is shown in Figure 12 for one pair of pipe tests. Current practice is to estimate separately the embedment due to pipe laying and due to dynamic effects. The initial embedment on pipe laying involves temporary overload at the touch down point. Typically the pipe weight is increased by a lay factor and the initial pipe penetration under monotonic loading is computed for this higher load. The effect of small amplitude cyclic lateral motion is incorporated using a dynamic embedment factor that multiplies up the initial monotonic pipe lay embedment to determine a final estimated pipe embedment. The centrifuge tests and analysis here did not 3.2 Pipeline fault crossing Damage is caused to pipelines that cross a seismic fault that subsequently displaces. Rupture of oil or gas pipelines can cause fire and environmental risk. For critical pipelines, the magnitude and direction of localized fault displacement should be assessed and appropriate engineering implemented to avoid pipe rupture due to ground movement. Moradi et al completed centrifuge tests on buried steel pipe subject to an upward thrust fault at 3⁰ to the vertical in the direction of the pipe. A fault displacement 7mm was applied across an 8mm diameter buried pipe with.4mm wall thickness tested in a centrifuge at 4g. In one test the pipe is simply buried in the compacted sand. A low density and light weight loose backfill was used in the second test. The axial and bending strain in the pipe was measured in both tests. The light backfill allowed the pipe to buckle and displace over a greater length considerably reducing the damage to the pipe. The pipe embedded in the sand suffered more localized deformation and damage, as illustrated by the photos post testing, Figure

9 Proceedings of the 18 th International Conference on Soil Mechanics and Geotechnical Engineering, Paris 213 (Moradi et al 213) Figure 14. Pipe response to shear fault displacement in a centrifuge. 3.3 Seabed structures that displace horizontally The sea bed in deep water is generally soft and often requires large shallow foundations to support seabed facilities. If some movement could be tolerated the size could be reduced. Further, if the structure connects with a pipeline subject to walking or other axial force, there may be merit in allowing the structure to slide horizontally to help relieve concentrated load. Bretelle and Wallerand (213) examine the design for a shallow foundation that displaces horizontally in a cyclic fashion, as might be caused by repeated pipe expansion and contraction. The influence of soil softening, foundation settlement and potential change in stiffness with time is examined through relatively straightforward analysis. The authors conclude that shallow foundations designed to displace horizontally could be useful for subsea pipeline networks. 3.4 Large displacement in dense sand While numerical analysis for large deformation is increasingly amenable for engineering design, a relatively simple constitutive model for dense sand that provides stable large deformation analysis is still subject to study. Li et al (213) propose a Critical State Mohr Coulomb (CSMC) model: deformation up to peak strength is elastic and thereafter dense sand dilates (including non-associated flow) and reduces in strength to the critical state angle of friction. The concept of the state parameter defined by Been and Jeffries (1985) is used. A key objective is analysis for punch through of a spudcan footing in dense sand overlying soft clay. Li et al (213) have not reached that target. However, development of the model starkly highlights non-uniform deformation and preferential shear band formation in dense sand post peak that makes data acquisition (lab tests) and model calibration such a challenge. In analysis for bearing capacity of a circular plate on uniform sand, the authors found that the elastic stiffness of the sand influences bearing capacity by as much as 5% over the realistic range, reminiscent of rigidity index in penetration problems. Stiffness was found to have greater impact than dilation angle. The analysis for bearing capacity is described in terms of a combined bearing capacity factor N qγ that applies across the range from N q alone to N γ. The proposed formula for N qγ includes soil stiffness and dilation angle along with peak friction angle, foundation size, soil unit weight and surcharge. 4 SEABED SOILS. The three papers on soil properties cover diverse topics. Ho et al (213) describe undrained cyclic triaxial compression tests on isotropically consolidated Singapore Marine Clay. The focus is the behavior of the clay when it is sheared monotonically to failure after cycling. The tests show that when the current mean effective stress in the sample reduces below half the original preconsolidation pressure, p /p c.5, due to cyclic loading, some increase in mean effective stress commences at higher stress ratio in each cycle. At mean effective stress p /p c.6 (first few cycles) the mean effective stress of the clay always reduces. This behavior is similar to normally versus over consolidated clay. The final effective stress path for monotonic triaxial compression to failure after cycling similarly depends on the mean effective stress p /p c at the end of cycling. However, the shear strength is found to be largely independent of the previous number of load cycles and the strain amplitude. Kim and Safdar (213) report cyclic direct simple shear tests on compacted silty sand to define the limiting cyclic stress ratio versus number of cycles for two initial void ratios. Tyldesley et al (213) describe site investigation to define parameters for wind farm foundation design on a deep deposit of carbonate clayey silt till in Ontario Canada. This onshore site investigation demonstrates the use of insitu tests and shear wave velocity measurement, interpreted together with laboratory tests, to assemble knowledge on soil strength and stiffness properties. 5 COASTAL AND NEARSHORE WORK. There are three papers on diverse topics. Madrid et al (213) describe site investigation, cyclic laboratory tests and numerical analysis for the stability of a caisson breakwater in about 2m to 25m depth of water. The caissons are founded on a rubble mound infilling a large zone where the deep underlying soft clay soil was removed, Figure RIP RAP 3kg CAISSON RUBBLE MOUND CONCRETE BLOCK RIP RAP 3kg RIP RAP 4 ton SEAWARD SIDE. (Madrid et al 213) Figure 15. Caisson breakwater and stability analysis for wave impact. There is much detail in the paper on soil testing and soil properties, loading cases for various phases of project construction and hydrodynamic testing to determine dynamic uplift. A good description is provided on the way cyclic loading and shear strength reduction were treated for design. Relic footprints from earlier jack-up activity can occur next to the location for new shallow foundations. Ballard and Charue (213) describe a study on a circular zone of remoulded soft clay (S r = 2) with a diameter equal to the size of the square mudmat and with soft clay thickness of half that size. The limiting envelope for combined moment and horizontal resistance is computed for a range of applied vertical load (V/V ult ), and a range of distance between the mudmat and the remoulded zone/footprint that causes the moment and horizontal resistance to be reduced, as well as V ult. 2D and 3D analyses show very substantial benefit from the 3D geometry in this case. A detailed design and project record for installation of large diameter, buried HDPE pipes in a nearshore environment prone to seismic loading is described by Bellezza et al (213). Details for the case history and the various code requirements considered in design are documented. Initial measurements are provided on the vertical deflection of the installed pipes. 6 PILE FOUNDATIONS. A lack of code guidance on capacity, stiffness and displacement for cyclically loaded piles is being addressed by collaborative research including the original GOPAL study and the current SOLCYP project, supplemented by individual research work. Several papers to this session report on SOLCYP results from instrumented field tests, calibration chamber and centrifuge

10 Technical Committee 29 / Comité technique 29 tests. SOLCYP will be presented and discussed at the TC29 workshop and recorded for publication. Therefore only some key aspects are described below to avoid duplication. There are four axial load magnitudes: the mean Q mean and the half-amplitude of the cyclic load Q cyclic define the maximum Q max = Q mean + Q cyclic and the minimum Q max = Q mean - Q cyclic pile loads. These loads are typically referenced to the ultimate pile capacity in tension Q T or compression Q UC. The ultimate capacity and the capacity under cyclic load is determined at a limiting displacement (.1D or less) or due to an increasing rate of displacement; either continuing displacement after a static load increment or the cyclical displacement rate (mm/cycle). 6.1 Stability diagram: cyclic axial loading The stability diagram is a non-dimensional map for cyclic pile behavior. The diagram in Figure 16 is for axial tension tests on model driven piles in dense sand in a calibration chamber (Silva et al, 213). A similar diagram is found for the equivalent field test data (Rimoy et al, 213). The chart defines the region of stable cyclic load combinations for a number of load cycles to be resisted. One way loading (Q mean >Q cyclic ) is more stable. The dashed line shows the limit of valid load combinations. Puech et al (213) provide the equivalent stability diagrams for cyclic compression loading of bored piles in dense sand from field and centrifuge tests. The field tests by Benzaria et al (213a) are shown on Figure 17. The results from centrifuge model tests are very similar. Note that only one-way loading was tested and the data should not be extrapolated beyond. While bored piles in stiff clay have lower ultimate resistance compared with driven piles, cyclic compression tests on bored piles in over-consolidated clay indicate a much larger range of stable load combinations compared with piles in dense sand (Benzaria et al, 213b). Information on deformation can be included on the stability diagram as shown by Rimoy et al (213) for field tests under cyclic axial tension on driven piles in sand, Figure 18. The data on the accumulation of displacement with cycling shows mostly stable behavior that suddenly degrades near the limiting number of cycles, a rather brittle behavior under tensile load. A consistent observation for driven piles in dense sand is that stable cycling increases the ultimate capacity when tested subsequently. This is attributed to densification with some relaxation in lateral effective stress around the pile, as measured in the exceptional data of Silva et al. (213). Conversely, strong cyclic loading (as would lead to instability) acts to reduce the ultimate axial pile capacity that can be mobilized subsequently. Data on large axial pile tests in silt on a 4.2m long test pile are also interpreted in a stability diagram by Chen et al (213). 6.2 Cyclic lateral loading Rosquoët et al (213) report data on lateral displacement for model driven piles in sand tested under one-way loading at 4g in a centrifuge. As in 2.3.3, but for displacement rather than rotation, the lateral displacement y N compared with the first load displacement y 1 [Δy N = y N y 1 ] is linked with the number of cycles. A logarithmic form Δy N / y 1 = c.ln(n) fits the data well where c varies with the amplitude of cyclic load and the maximum lateral load (equivalent to 2Q cyclic / Q max in the axial terminology above, written as DF/F by Rosquoët et al). Based on their test data the authors suggest c =.1(DF/F).35 as a general fit, but sand and pile properties might also be important. (Tsuha et al (212) calibration chamber tests) Figure 16: Stability diagram: tension tests on model driven piles in sand. (Puech et al, 213) Figure 17: Stability diagram: field compression on bored piles in sand. (Rimoy et al, 213 field tension tests on driven piles in sand) Figure 18: Stability diagram: accumulated displacement. The natural logarithm form has the advantage of fitting the data through to low numbers of cycles (recall 2.3.3). Rosquoët et al (213) also note that the maximum moment in the pile does not increase significantly with lateral cyclic loading. Finally, the work to extend p-y analysis for laterally loaded piles failed to capture the measured behavior beyond the first few cycles. 231

11 Proceedings of the 18 th International Conference on Soil Mechanics and Geotechnical Engineering, Paris Extension of t-z analysis to cyclic loading Burlon et al (213) extend static t-z analysis for piles in sand to the case of tensile and compressive cyclic load and compare the results with centrifuge test data. The analysis turns out not yet to be practical as four new parameters are introduced that require measurement in cyclic pile tests. Even then, the fit with the test data is good only for relatively few load cycles. 6.4 Plugging of open-ended displacement piles Laboratory tests on sand to measure pile plugging, using PIV observations, are described by Lueking and Kempfert (213). The fully plugged limit IFR= was not achieved in the tests. The results of 2D Plaxis analysis are reported to investigate the mechanism of plugging. Based on this study, the authors propose two largely empirical calculation methods for the analysis of end-bearing for open-ended, partially plugged piles. 6.5 Cyclic pressuremeter tests Reiffsteck et al (213) report new work on the application of Ménard and self-boring pressuremeter tests to measure the change in soil properties with cyclic loading and, potentially, liquefaction resistance. Data are reported for several sites where cyclic pile tests were completed (SOLCYP). The authors emphasise the importance of a high quality borehole and the need for at least 5 cycles of repeated load to characterize the shear modulus. Soil horizons susceptible to liquefaction could be identified by large volume expansion. The acquisition of pore water pressure data during the test would greatly improve the test control (data on drainage/rate effects) and interpretation. 6.6 Osterberg cell testing The general reporter is not aware of recent cyclic pile tests using O-Cell technology (Osterberg, 1989). A two level O-Cell test arrangement, for example, permits end bearing to be eliminated from cyclic axial compression tests on piles. 7 CONCLUSIONS. As demonstrated by papers to this session, the practical challenges of offshore geotechnics actively drive forward the development of soil mechanics and geotechnical engineering. This is partly due to more extreme loading and deformation than usually encountered onshore. The fruits of this research and development are of great value for the overall understanding and pactice of geotechnical engineering. That is why offshore geotechnics should remain part of this key ISSMGE conference. 8 ACKNOWLEDGEMENTS The author would thank TC29 Chairman Philippe Jeanjean for the invitation to prepare this General Report, and colleagues J.C. Ballard, P. Peralta and V. Whenham for valuable support. 9 REFERENCES Arroyo M., Abadias D., Alcoverro J. and Gens, A Shallow foundations for offshore wind towers. Proc. 18th ICSMGE, Paris. Arshi H.S., Stone K.J.L., Vaziri M., Newson T.A., El-Marassi, M., Taylor R.N. and Goodey R.J Modelling of monopile-footing foundation system for offshore structures in cohesionless soils. Proc. 18th ICSMGE, Paris. Ballard J.C. and Charue N Influence of jack-up footprints on mudmat stability - How beneficial are 3D effects? Proc. 18th ICSMGE, Paris. Been, K. and Jefferies, M.G. (1985). A state parameter for sands. Géotechnique, Vol. 35(2), pp Bellezza I., Mazzieri F., Pasqualini E., D Alberto D. and Caccavo C Design and installation of buried large diameter HDPE pipelines in a coastal area. Proc. 18th ICSMGE, Paris. Ben-Hassine J. and Griffiths D.V Geotechnical Exploration for Wind Energy Projects. Proc. 18th ICSMGE, Paris. Benzaria O., Puech A. and Le Kouby A. 213a. Essais cycliques axiaux sur des pieux forés dans des sables denses. Proc. 18th ICSMGE. Benzaria O., Puech A. and Le Kouby A. 213b. Essais cycliques axiaux sur des pieux forés dans l argile surconsolidée des Flandres. Proc.18th ICSMGE, Paris. Bretelle S., Wallerand R. Fondations Superficielles Glissantes pour l Offshore Profond Méthodologie de Dimensionnement. Proc. 18th ICSMGE, Paris. Burlon S., Thorel L. and Mroueh H Proposition d une loi t-z cyclique au moyen d expérimentations en centrifugeuse. Proc. 18th ICSMGE, Paris. Chen R.P., Ren Y., Zhu B. and Chen Y.M Deformation behavior of single pile in silt under long-term cyclic axial loading. Proc. 18th ICSMGE, Paris. Cheuk, Y.C. and White, J.D. (28). Centrifuge modelling of pipe penetration due to dynamic lay effects. Proc. Int. Conf. on Offshore Mechanics and Arctic Engineering, Portugal. OMAE Damgaard M., Andersen J.K.F., Ibsen L.B. and Andersen L.V Time-Varying Dynamic Properties of Offshore Wind Turbines Evaluated by Modal Testing. Proc. 18th ICSMGE, Paris. DNV-OS-J11 (211). Design of Offshore Wind Turbine Structures. Det Norske Veritas (DNV) Offshore Standard, September 211. Dutta S., Hawlader B. and Phillips R Numerical investigation of dynamic embedment of offhore pipelines. Proc. 18th ICSMGE. Einav, I. and Randolph, F.M. (25). Combining upper bound and strain path methods for evaluating penetration resistance. Int. J. Numer.Meth. Engng., Vol. 63, pp Ho J., Goh S.H. and Lee F.H Post Cyclic Behaviour of Singapore Marine Clay. Proc.18th ICSMGE, Paris. Kim D.J., Youn J.U., Yee S.H., Choi J., Choo Y.W., Kim S., Kim J.H., Kim D.S. and Lee J.S Centrifuge test and numerical modelling for a suction bucket monopod foundation. Proc. 18th ICSMGE, Paris. Kim J.M. and Safdar M Behaviour of marine silty sand subjected to long term cyclic loading. Proc. 18th ICSMGE, Paris. Li X., Hu Y. and White D A large deformation finite element analysis solution for modelling dense sand. Proc. 18th ICSMGE. Lueking J. and Kempfert H.-G Plugging Effect of Open-Ended Displacement Piles. Proc. 18th ICSMGE, Paris. Moradi M., Galandarzadeh A. and Rojhani M The new remediation technique for buried pipelines under permanent ground deformation. Proc. 18th ICSMGE, Paris. Muir Wood A. and Knight P Site investigation and geotechnical design strategy for offshore wind development. Proc.18th ICSMGE. Osterberg, J.O New Device for Load Testing Driven Piles and Drilled Shafts Separates Friction and End Bearing. Proc. Int. Conf. on Piling and Deep Foundations, London, A.A. Balkema, p Peire, K., Nonneman, H. & Bosschem E. (29) Gravity Base Foundations for the Thornton Bank Offshore Wind Farm. Terra et Aqua, No. 115, pp Puech A., Benzaria O., Thorel L., Garnier J., Foray P., Silva M. and Jardine R Diagrammes de stabilité cyclique de pieux dans les sables. Proceedings 18th ICSMGE, Paris. Reiffsteck P., Fanelli S., Tacita J.L., Dupla J.C. and Desanneaux G Utilisation des essais d'expansion cyclique pour définir des modules élastiques en petites déformations. Proc. 18th ICSMGE. Rimoy S., Jardine R. and Standing J Displacement response to axial cyclic loading of driven piles in sand. Proc. 18th ICSMGE. Roesen H.R., Ibsen L.B. and Andersen L.V Experimental Testing of Monopiles in Sand Subjected to One-Way Long-Term Cyclic Lateral Loading. Proc. 18th ICSMGE, Paris. Rosquoët F., Thorel L., Garnier J. and Chenaf N Pieu sous charge latérale : Développement de lois de dégradation pour prendre en compte l effet des cycles. Proc. 18th ICSMGE, Paris. Silva M., Foray P., Rimoy S., Jardine R., Tsuha C. and Yang Z Influence des chargements cycliques axiaux dans le comportement et la réponse de pieux battus dans le sable. Proc. 18th ICSMGE. Tyldesley M., Newson T., Boone S. and Carriveau R Characterization of the geotechnical properties of a carbonate clayey silt till for a shallow wind turbine foundation. Proc. 18th ICSMGE. Versteele H., Stuyts B., Cathie D. and Charlier, R Cyclic loading of caisson supported offshore wind structures in sand. Proc.18th ICSMGE, Paris. Wang, D., White, D. J. and Randolph, M. F. (21). Large deformation finite element analysis of pipe penetration and large-amplitude lateral displacement. Canadian Geotech. Jrnl, Vol. 47, pp

12 Shallow foundations for offshore wind towers Fondations superficielles pour des installations éoliennes maritimes Arroyo M., Abadías D., Alcoverrro J., Gens A. Dep. of Geotechnical Engineering and Geosciences, Technical University of Catalonia, Barcelona, Spain ABSTRACT: Direct foundations are present in about 25% of the installed offshore wind power towers. The peculiarities of this type of structure are well known: high dynamic sensitivity, complex couplings between environmental actions, machine operation and structural response, complex installation and maintenance, difficult site investigation. There is a clear need for optimized foundation design tools that would enable cost reduction and a more detailed assessment of the risk of every installation. One such tool is likely to be the systematic use of failure envelopes for capacity checks. The paper explores the benefits of such an approach with various realistic design examples. RÉSUMÉ : Les fondations superficielles interviennent dans la réalisation de 25% des structures éoliennes maritimes. Les particularités de ce type de structures sont bien connues: haute sensibilité dynamique, couplages complexes entre les actions environnementales, le fonctionnement de la machine et la réponse structurelle, installation et maintenance difficiles, investigation géotechniques onéreuses. Un besoin évident d'optimisation des outils de conception est nécessaire pour permettre la réduction des coûts et une évaluation plus détaillée du risque de chaque installation. Le recours systématique à des enveloppes de rupture pour les justifications de la capacité portante des fondations peut bien être un tel outil. Ce papier explore les avantages d'une telle approche avec divers exemples de conception réalistes. KEYWORDS: direct foundation, capacity, offshore, energy, wind farms 1 INTRODUCTION Offshore wind is an increasingly large contributor to the energy production mix of several European countries, particularly those bordering the North and Baltic seas. An exponential increase in installations is currently anticipated in this region. It is reasonable to expect that other regions of the world will follow suit. Offshore wind turbines (OWT) are generally larger than those installed on land, with 3 to 5 MW of nominal capacity being now the norm, but with turbines of up to 1 Mw coming soon to the market. Rotor diameters of more than 1 m and nacelle locations 8 m above mean sea level are common. The result is a relatively lightweight and slender structure, supporting a rotating machine finely tuned to maximize power production while minimizing structural loading. While initial OWT installations took place near shore (< 1 km) at locations with relatively shallow water depths (< 2 m), current developments are clearly located offshore (1-1 km from the nearer coast) with water depths of 2-5 m being typical. Several floating support concepts are now being developed; however, commercial installations are still always supported by some kind of fixed structure. For these, the foundation of choice would depend in any case on the particular site conditions, construction equipment availability and, to a certain extent, local traditions. To this date pile foundations have been largely dominant, mostly as single large (4-6 m diameter) monopile installations, and lately also as smaller (1-2 m) piles for jackets and tripods. However, examining the industry databases (e.g. Burton et al 211) it appears that at the end of 211 about 25% of the installed power was supported by direct foundations or gravity base substructures (GBS). Most of these GBS installations took place in relatively shallow waters, but there are some examples already at larger distances from the coastline and in deeper waters. Perhaps the most significant is the Thornton Bank I project, 27 km offshore Zeebrugge in Belgium, where 6 OWT of 5 Mw were installed in water depths of 2-3 m. The foundation design for this installation was described by Peire et al (29) and its outline is reproduced here in Figure 1. These are large (44 m height; 23.5 m base diameter) concrete shells, floated into place and later ballasted with a mixture of sand and olivine with the base at 4 m below the original seafloor level. The geotechnical profile at the site comprises medium and high density sands and stiff tertiary clays. 2 DESIGN ISSUES FOR DIRECT OWT FOUNDATIONS There are several specific standards dealing with OWT. Perhaps the highest ranked is IEC (29) which, from the point of view of structural design, establishes design cases and site ambient load specification procedures, introduces a safety format and gives broad indications about structural design procedures. However, detailed specification of structural and foundation design procedures is deliberately referred to other documents, like the ISO 199X offshore standard series or DNV-OS-J11 (21). As might be expected, the indications given by such standards are, on the one hand, firmly based in conventional design practice when being specific, and somewhat elusive with problems that lack a clear conventional solution. An example of the later is the consideration of fatigue or foundation failure under cyclic loading. An example of the former is the consideration of foundation bearing capacity which, for shallow foundations, follows a conventional superposition and correction procedure not very different from those outlined by Brinch-Hansen (197) or Vesic (1975). When designing foundations for OWT, there will be of course issues of geotechnical capacity under extreme loads. However the design drivers might be sometimes related to other considerations, such as dynamic characteristics of the whole 233

13 Proceedings of the 18 th International Conference on Soil Mechanics and Geotechnical Engineering, Paris 213 structure (Van der Temple and Molenaar, 22) or displacement limits imposed by operating constraints (e.g. foundation tilting limits of are sometimes quoted). However, even if we narrow our focus to bearing capacity considerations there are reasonable grounds to question the suitability of the conventional design approach. 3 FAILURE ENVELOPES 3.1 Concept Failure envelopes were introduced (Butterfield & Ticof, 1979) as an alternative to classical bearing capacity analyses. They were based on the concept of interaction diagram, which was applied to the system of loads acting on the foundation. Most developments to date but not all-, refer to the case in which that system can be reduced to loads acting within a plane (V, H, M) where M represents the moment acting within the plane, M normalised by a characteristic foundation dimension, M/B. Failure envelopes are implicit in the traditional approach to bearing capacity. However, it was clearly appreciated from the beginning that an explicit failure envelope was useful to link previously separate checks on different foundation failure modes (e.g. sliding and bearing capacity) into a coherent view. Failure envelopes offered advantages also from the experimental viewpoint, because they provide a clearer framework for experimentation, even suggesting new, more efficient, procedures (like swipe tests). Failure envelopes are also attractive because they can fit well with generalized force-displacement foundation models ( macroelements ; Nova and Montrasio, 1991) that are used to compute foundation displacements and represent an economical solution to non-linear soil-structure interaction studies. Finally, failure envelopes are interesting because they enable a more coherent approach to foundation safety. 3.2 Safety considerations Already Georgiadis (1985) clearly identified as one major advantage of failure envelopes that they allow a very natural consideration of the influence of different loading paths. To do that, it is important to distinguish between the reference design load state and incremental loading paths (Figure 2). M Figure 1. Thornton Bank GBS (Peire et al, 29). Indeed there are several aspects of the traditional approach to bearing capacity that are poorly suited to deal with OWT. Firstly, using separate corrections for shape, depth, load inclination, load eccentricity is cumbersome and prone to calibration error if the effects that are being corrected for are not truly independent. This is perhaps the reason behind the large scatter between inclination factor formulations (Siefert & Bay- Gress; 2); that uncertainty is particularly undesirable for structures, like OWT, that are mostly designed to sustain horizontal loads. Secondly, the traditional approach to bearing capacity quickly leads to conundrums when the security format (as is the case for most modern codes, like DNV-OS-J11) is based on separate partial factors for loads and resistances. As discussed in detail by Lesny (27) the same action might have a detrimental or favourable effect depending on which other actions are being simultaneously considered. Also it is fairly evident that a traditional bearing capacity check is far from eliminating the most likely path towards failure. Finally, it is very difficult to generalize the traditional approach to cases when two major horizontal loads (wind, wave) are acting in separate planes. All these problems are best dealt with if the traditional approach to capacity checks is replaced by a failure-envelope based one. (H, M) (H, M) (H r, M r ) Figure 2. Schematic load envelope illustrating a reference design load and one incremental load path Any load system (V, H, M) shall remain within the failure envelope. It is however convenient to establish a nondimensional safety measure. To do so a simple approach is, for any incremental loading direction, to obtain the crossing point with the failure envelope (V r, H r, M r ) and then define a generalized safety factor, SF, as SF V,H, M 1 (1) V,H, r M ( V V V, H H H, M M M) (2) r r r r r r It is thus made explicit the fact that safety is not only dependent on the initial design situation but also on the incremental loading path. This definition includes, as a particular case, the traditional safety factors against bearing capacity (the incremental load direction and the reference H 234

14 Technical Committee 29 / Comité technique 29 design load are collinear) or sliding (incremental loading direction collinear with the Horizontal component of the reference design load). Another particular case included is that of plastic overturn, a prescribed check for breakwater design in Spanish regulations (Puertos del Estado, 25) in which the lever arm of the horizontal loading is maintained (i.e. the incremental load is aligned with the the Horizontal and Moment components of the reference load). 3.3 Example formulations 4 EXAMPLE APPLICATION To illustrate the argument we propose an example, synthetic but realistic. The case is developed using the characteristics of the gravity base substructure built at Thornton Bank (Peire et al. 29) and the design loading specified for a Baltic windfarm development site, Kriegers Flak (Bulow et al, 29). This reference gives some basic characteristics for the OWT superstructure (Table 1). Table 1 Super-structure characteristics Rated power Rotor diameter Nacelle height above msl Nacelle-rotor weight Tower weight 5 MW 126 m 9 m 4.1 MN 3 MN Figure 3 Failure envelope by Gottardi et al (1999) There are many failure envelopes in the literature. For foundations failing without drainage at the soil-foundation interface Gourvenec & Randolph (211) offer an excellent review. For the example below a sand profile is assumed and drained conditions are reasonable. In these circumstances a convenient expression for a failure envelope is that proposed by Gottardi et al. (1999) (Figure 3) FV (, H, M) 2 2 h m hm 2 2 a (4 v(1 v)) h m hm (3) Where (a, h, m ) are shape factors, empirically determined as (-.22,.12,.9) for quartzitic sand, and we use a nondimensional notation in which v = V/V, h = H/V, m = M/(DV ) and D is the foundation diameter. The normalizing factor V is the maximum load (i.e. centered vertical) that the foundation can sustain. Here that maximum load is computed assuming no embedment and introducing the bearing capacity factor N from Bolton & Lau (1993) into 1 D V DN 2 4 It is worth noting that (a) it is relatively straightforward to generalize expression (3) to more complex loading situations e.g. Lesny 21- although the experimental base for adjusting the parameters in those circumstances is somewhat scarce, (b) that the shape of (3) above has proven rather resilient and very similar expressions have been found to fit well other foundation test results in materials like carbonate sand or even clay (Martin & Houlsby, 21), as long as the contact surface remains drained. Of course the choice of V would change according to the material and foundation shape. 2 (4) The same reference also includes resultants from ambient loads for a range of depths and load hypothesis (e.g. extreme, fatigue). Using these data, Table 2 has been computed for a 3 m depth case and extreme load scenario. It appears that, in this particular case, 8% of the total horizontal thrust is due to sea action, but this load is the source of less than 2% of the overturning moment at foundation level. This might partly reflect the fact that at that particular site sea current is relatively strong, lowering the action line of sea forces. These ambient loads should be combined with the OWT selfweight. Using the Thornton Bank design like a template for substructure shape, the relevant characteristics of that part of the OWT are those listed in Table 3. As usual with gravity base OWT, the dead weight of the substructure is significantly larger than that of the superstructure. Combining all environmental actions and structure selfweight the resultant load combination acting at the foundation level is (H, V, M) = (1.1; 44.5; 284.3) in MN and MNm. This will be the reference design load state in this example. Tabla 2 Ambient loading parameters Parameter / load Unit Value Total thrust, H MN 1.1 Total overturning moment, M MNm Wind thrust, Hw MN 2.3 Wind arm lever, bw m 12 Sea thrust, Hs MN 8.7 Sea arm lever, bs m 5 Tabla 3 Thornton Bank type substructure characteristics Parameter / load Unit Value Base diameter m 23,5 Concrete weight MN 3 Fill weight MN 38 Buoyant volume m From that reference state we probe the failure surface alongside three different incremental loading paths. One will correspond to a simultaneous and proportional increase of all ambient actions (the plastic overturn case). The other two hypothesis would correspond to increases of just one of the 235

15 Proceedings of the 18 th International Conference on Soil Mechanics and Geotechnical Engineering, Paris 213 ambient horizontal actions, (sea, wind) while the other remains constant. These hypotheses would, for instance, naturally follow from any circumstance in which the estimates of wind and wave carry different uncertainties. Figure 4 illustrates graphically the meaning of these load directions in an idealised section of the failure envelope at constant V. For this check we use the failure envelope of Gottardi et al described above. The soil profile below the foundation is characterised by a friction angle of 33 and submerged weight of 1 kn/m 3. These values might correspond well to the characteristic values of a medium-dense sand profile, frequently encountered in North Sea locations. It is assumed that the foundation base is perfectly rough. 5 CONCLUSION Failure envelopes offer a powerful framework to analyze shallow foundation capacity problems. They seem particularly suitable for offshore wind towers, where refined design in the face of large load uncertainties is likely to be a frequent situation. 6 ACKNOWLEDGEMENTS The research on direct foundations for offshore wind towers described in this paper was partly founded by the company ACCIONA ENERGY within the framework of the CENIT- AZIMUT project supported by the Spanish Ministry of Science. 7 REFERENCES (TNR 8) Figure 4 Incremental load paths in the example Table 4 Example: results Hypothesis Hr (MN) Hr / Hi ΔH (%) Sea Both Wind Some relevant results from the computation are presented in table 4. For each incremental loading path a failure point is identified in the envelope, with values (H r, M r ). In the table the value H r corresponding to each loading path is reported in the first column. In the second column this value is normalized by the reference state horizontal load. This corresponds to the generalized safety factor defined above, which, only for the hypothesis in which both loads are simultaneously increasing, coincides with the plastic overturn safety factor of ROM.5-5. As a reference the value required for that safety factor in breakwaters is commonly above 1.3 (Puertos del Estado; 25). For the other two load hypothesis in which only one environmental action is increased no similar reference exists to judge on the computed safety factor. For these cases it is perhaps more meaningful the number in the third column of Table 4, where the difference between failure and reference thrust is expressed as a percent of the reference ambient load that is increasing. In the case computed, a 21% error in the reference estimate of wind thrust would result in foundation failure, whereas it would be necessary a 5% underestimate of the hydrodynamic thrust to fail the foundation. The previous computations have always been made under the hypothesis of increased thrust and constant lever arm. This can be interpreted as action magnitude uncertainty. Alternative hypothesis dealing with lever arm uncertainty can be equally set up with relative ease. Note, finally, that most geotechnical uncertainty can be lumped in the V estimate to achieve a relatively straightforward approach to reliability evaluation. Brinch-Hansen J. (197). A revised and extended formula for bearing capacity. Danish Geotechnical Institute Bulletin, n 28, 5-11 Bülow, L., Jorgensen, L. and Gravessen, H. (29) Kriegers Flak Offshore Wind Farm. Basic data for conceptual design of foundations. March 29, Vattenfall Vindkraft AB Burton, T., Sharpe, D., Jenkins, N. and Bossanyi, E. (211) Wind Energy Handbook, 2nd Edition, John Wiley & Sons, Chichester, UK Butterfield R., Ticof J. (1979). The use of physical models in design (discussion). 7th European Conference on Soil Mechanics and Foundation Engineering, Brighton, UK, Vol.4, Georgiadis, M (1985) Load-path dependent stability of shallow footings, Soils & Foundations, 25,1, Gottardi, S., Houlsby, G.T. y Butterfield, R. (1999) Plastic response of circular footings on sand under general planar loading, Géotechnique 49, No. 4, Gourvenec, S. y Randolph, M. (211) Offshore geotechnical engineering, Spon Press, New York DNV (21) DNV-OS-J11 Design of offshore wind turbine structures IEC (29) International standard Wind turbines Part 3: Design requirements for offshore wind turbines. International Electrotechnical Comission Lesny, K. (21) Foundations for offshore wind turbines : tools for planning and design, VGE, Essen Lesny, K. (27) Design approaches of Eurocode 7 and their effect on the safety of shallow foundations, ICASP1, Applications of statistics and probability in Civil Engineering, Taylor & Francis Martin, C., M., Houlsby, G. T. (21) - Combined loading of spudcan foundations on clay: numerical modeling. Géotechnique, 51, No. 8, pp Nova R. y Montrasio L. (1991). Settlements of shallow foundations on sand. Géotechnique, vol.41(2), Peire, K., Nonneman, H. & Bosschem E. (29) Gravity Base Foundations for the Thornton Bank Offshore Wind Farm. Terra et Aqua, N. 115, pp Puertos del Estado (25) ROM.5-5 Recomendaciones Geotécnicas para Obras Marítimas y Portuarias, Sieffert, J.G., y Bay-Gress, Ch. (2). Comparison of the European bearing capacity calculation methods for shallow foundations; Geotechnical Engineering, Institution of Civil Engineers, Vol. 143, pp Van der Tempel, J. y Molenaar, D.P. (22) Wind turbine structural dynamics a review of the principles for modern power generation, onshore and offshore, Wind Engineering, 26,4, Vesic, A. S. (1975) Bearing capacity of shallow foundations, Ch. 3 in Winterkorn H.F. & Fang H.Y., Foundation Engineering Handbook, Van Nostrand Reinhold 236

16 Modelling of monopile-footing foundation system for offshore structures in cohesionless soils Modélisation d un système de fondation superficielle isolé pour sur les structures maritimes dans les sols pulvérulents Arshi H.S., Stone K.J.L. University of Brighton, UK Vaziri M. Ramboll UK Limited, UK Newson T.A., El-Marassi M. University of Western Ontario, Canada Taylor R.N., Goodey R.J. City University London, UK ABSTRACT: While monopiles have proven to be an economically sound foundation solution for wind turbines, especially in relatively shallow water, their installation in deeper water and in hard ground may require a more complex foundation design in order to satisfy the loading conditions. One approach is that foundation systems are developed which combine several foundation elements to create a hybrid system. In this way it is possible to develop a foundation system which is more efficient for the combination of vertical and lateral loads associated with wind turbines while maintaining the efficiency and simplicity of the design. Previous studies have reported the results of single gravity tests of the hybrid system where the benefits of adding the footing to the pile are illustrated. This paper presents experimental results on the performance of skirted and unskirted monopile-footings. A simplified design approach based on conventional lateral pile analysis is presented. RÉSUMÉ : Alors que les fondations de type monopile se sont révélées être une solution économiquement viable pour les fondations d éoliennes, en particulier dans les eaux relativement peu profondes, leur installation dans des eaux plus profondes et dans un sol dur peut exiger une conception plus complexe afin de satisfaire les conditions de chargement. Une approche possible est que les systèmes de fondations développés combinent plusieurs éléments de fondation pour créer un système hybride. De cette manière, il est possible de développer un système de fondation plus efficace vis à vis des charges verticales et latérales associées aux éoliennes, tout en maintenant une conception efficace et simple. Des études antérieures sous gravité simple ont montré l efficacité d un système hybride en combinant une semelle et un pieu. Cet article présente des résultats expérimentaux sur la performance de systèmes avec et sans pieu pour des semelles. Une approche de conception simplifiée basée sur l'analyse classique d un pieu sous charge latérale est présentée. KEYWORDS: Hybrid monopile footing, offshore piles, laterally loaded piles, wind turbine foundations 1 INTRODUCTION Due to the needs of on-going developments in the oil and energy sector, the design of offshore foundations is constantly evolving. In the hydrocarbon extraction sector, exploration and development is moving in to ever deeper water resulting in ever more challenging geotechnical conditions. Similarly the expansion of the offshore wind sector involves the development of deepwater sites, together with requirements for heavier high capacity turbines. Conventional offshore foundations are not always economical or practical for this new generation of turbines, and there remains a requirement to develop foundation solutions which can better satisfy future developments in the offshore wind sector. The foundations of a typical offshore wind turbine are subjected to combined loading conditions consisting of the selfweight of the structure (V), relatively high horizontal loads (H) and large bending moments (M). The preferred foundation system to date has been the monopile, which has the advantage that it can be employed in a variety of different soil conditions. However, a disadvantage in the use of monopiles in deep water sites is that the system can be overly compliant. For sites with intermediate water depths, it may be possible to stiffen the lateral response of the monopile at the mudline. Figure 1. Schematic illustration of the prototype hybrid system. One such approach to increase the lateral resistance of a monopile is the hybrid monopile-footing system. As schematically represented in Figure 1, this foundation system 237

17 Proceedings of the 18 th International Conference on Soil Mechanics and Geotechnical Engineering, Paris 213 comprises of a circular footing attached to the monopile at the mudline. A 2-D analogy of this system is that of a retaining wall with a stabilising base (Powrie and Daly, 27). The role of the footing is to provide a degree of rotational restraint at the pile head, leading to an improvement in the lateral resistance of the pile. It has also been shown that the use of a relatively thick pile cap leads to an increase in the lateral resistance through the development of passive soil wedges (Mokwa, 1999), in a similar way to the behaviour of skirted foundations (Bransby and Randolph, 1998). Analysis of the hybrid system would involve both lateral pile analysis and bearing capacity analysis. The lateral response of piles is well reported in the literature and various methods of analysis have been proposed by numerous researchers, such as Matlock and Reese (196), Broms (1954), Poulos (1971), Reese et al. (1974), Randolph (1981), Duncan et al (1994) and Zhang et al. (25). Where the plate diameter is relatively small, the system is similar to a single capped pile, for which methods have been developed for analysing the influence of the pile and pile cap under axial loading (Poulos and Randolph, 1983), and the effect of the pile cap on the lateral performance of single piles has also been investigated by others (Kim et al., 1979), (Mokwa and Duncan, 21: 23), (Maharaj, 23). The bearing capacity problem has also been investigated under different loading conditions relevant to offshore foundations, see for example references Houlsby and Puzrin (1999), and Gourvenec and Randolph (23). 2. EXPERIMENTAL INVESTIGATIONS The potential performance of the hybrid system was investigated in single gravity studies (Stone et al. (27)) and is illustrated in Figure 2. These studies suggested that the additional rotation restraint provided by the footing can result in a stiffer lateral response of the pile and greater ultimate lateral load. The degree of restraint at the pile head was dependent on the size of the footing, the initial contact between the soil and the footing and the stiffness of the soil beneath the footing. Observations of heaved and displaced soil in front of the edge of the footing also suggested that a degree of passive soil resistance is likely to be generated under the lateral movement and rotation of the footing. footing and the pile was also investigated where it was suggested that the hybrid foundation system tends to be more effective if vertical movements are allowed at the pile-footing connection. This movement allows the footing to act independently from the pile where the positive contact between the footing and the soil underneath is solely controlled by the vertical load acting on the footing. Table 1. Notations for skirted hybrid foundations system ID Footing size (mm) Skirt length (mm) Dead load (N) Footing to pile connection P.W P.F8.W1.FR 8 1 Slipping P.F8.S1.W1.FR 8 1 Slipping P.F8.S2.W1.FR 8 1 Slipping P.F8.S3.W1.FR 8 1 Slipping More recent single gravity tests are presented in Figure 3 where skirts with different lengths have been added to the footing. The tests were conducted in sand and the results indicate that the presence of the skirts has a relatively significant contribution on the lateral load capacity of the system. The results show that adding the skirts to the footing and increasing the skirt length tends to increase the lateral load bearing capacity of the foundation system by about 5% in comparison to a non-skirted hybrid system. It is also apparent that footings with very short skirts do not tend to show any apparent additional advantage to that without the skirt. This could be due to the fact that the stresses around the skirt induced by the soil are very small at 1g. Further studies in the centrifuge are in the taking place to investigate the effect of the skirts and the results will be reported soon. Figure 3. Load vs. deflection plot for the hybrid system with skirts. Figure 2. Lateral load response of the hybrid system (after Stone et al. 27). Arshi (211), and Arshi and Stone (212) reported the results of a comprehensive series of single gravity testes carried out on the foundation system where the elements affecting the overall performance of the foundation system was investigated in depth. It was reported that the size for the footing has a direct effect on the overall lateral load bearing capacity of the foundation system. Furthermore it was reported that the ratio between the vertical and horizontal load has a significant effect on the lateral performance of the foundation system where larger vertical loads tend to improve the lateral load bearing capacity of the hybrid system. The connection between the Stone et al (211) reported the results of a series of centrifuge tests in sand. The results of the combined vertical and lateral loading tests are best represented through plots of lateral load versus lateral displacement. Figure 4 shows a plot of the lateral load versus lateral displacement for the monopilefooting (HL 1) and single pile (PL 1) with a vertical load of 6N at 5 g. It is apparent from this plot that the initial lateral stiffness of the monopile-footing and pile are similar for the first 1 1.5mm of lateral displacement. However the monopilefooting continues to exhibit a stiffer response than the single pile as the lateral displacement increases. Further analyses of these data provided information on the redistribution of bending moment in the pile due to the plate. 238

18 Technical Committee 29 / Comité technique 29 Figure 4. Load deflection graph for centrifuge tests carried out on the hybrid system (after Stone et al. 211). In Figure 5, the bold lines represent the bending moments at 5% and 2% of the maximum deflection for the pile only case and the dashed lines show the behaviour of the hybrid system. The results show that adding the footing to the pile reduces the bending moment at any given deflection, and as a result increases the moment capacity of the system at any given applied lateral load. The results indicate about 25% improvement in the bending moment for at both deflections. developed by the footing to be generated as a function of the footing rotation. The results generated by this approach are illsutetared in Figure 6 where it is shown how different pile to footing diamater increses the moment capacityof the piles, where this variation lies between a fully free and a fully fixed pile. The dashed lines in Figure 6 show the ultimate moment capacities of the hybrid system. Although this method successfully leads to obtaining the ultimate load bearing capacity of the hybdegree of rigidity (D.O.R 75%, 5% and 25% showing the the ultimate capacity of the system when %75, 5% and 25% of the ultimate moment at pile head is applied to the free headed pile) of the system are shown as a benchmark for comparing how differenet pile to footing diameters relate to the fully fixed moment. As apparent in Figure 6, increasng the sie of the footing tends to increse the lateral load bearing capacity. As the footing size increases, it gets close to the fully fixed head condrion. This also indicates that there for a given pile diamater and length, there ought to be a footing size afterwhcih increseing the footing size further will not enhance the lateral load bearing capacity of the foundation system. Figure 6. Moment vs. rotation plot for the hybrid system with different pile to footing ratios. Figure 5. Bending moment distribution along the pile length for the hybrid system. 3. ANALYSIS Whilst some advanced numerical modelling of monopiled footings has been undertaken (El-Marassi et al. 28; Stone et al. 21; Arshi et al. 211; Arshi and Stone 212), the method presented here utilises conventional lateral pile analysis methodology where the hybrid system is idealised to a lateral pile with a resisting moment applied at the mudline. The resisting moment capacity provided by the footings were estimated analytically using conventional bearing capacity theory and applied at the mudline acting in the oposite dirtection to the loading. This approach only considers the ultimate condition of the system and does not allow the moment In addition to this, design charts have been developed which relate the pile embedment length to pile and footing diameters. Numerous design charts have been developed covering a wide range of pile diameters, pile lengths, footing diameters and normalized moment capacities an example of which is shown in Figure 8 where the L/D ratios vary from 1 to 1 and the footing to pile diameter ratios varies from to 1. The moment capacity of the hybrid system has been normalised and is shown against footing to pile diameter ratio. The lines in between represent different pile embedment depth where for a given moment capacity the designer could utilise this graph to choose the appropriate pile length as well as pile and footing diameters. It is also notable that for any given value of normalized moment capacity the designer has the option of choosing a short pile relatively large footing diameter, or long pile with relatively small footing diameter. The flexibility in this design approach is beneficial in particular designing the hybrid system in difficult soil conditions. 239

19 Proceedings of the 18 th International Conference on Soil Mechanics and Geotechnical Engineering, Paris 213 Figure 7. Example of a design chart for the hybrid system developed using analytical and numerical methods. 4. DISCUSSION & CONCLUSION It is apparent that the ultimate lateral response of a single monopile foundation can be enhanced by the presence of a footing resulting in a greater ultimate lateral capacity. This improvement was observed at both load versus deflection as well as the bending moment versus depth plots. Whilst the effect on the initial lateral stiffness may not be significant, the lateral stiffness beyond this initial movement was significantly enhanced through the presence of the footing. The effect of adding skirts to the hybrid system has been shown to further increase the lateral performance of the hybrid system, and centrifuge tests are planned to investigate the skirted system in more detail. A simple analytical approach using conventional lateral pile analysis methods is presented from which preliminary design charts can be generated. This approach can be developed to generate realistic design charts where the lateral capacity of the hybrid system is related to the development of bearing capacity coupled to the lateral resistance of the pile shaft. 5. REFERENCES Arshi HS. (211). Structural behavior and performance of skirted hybrid monopile-footing foundations for offshore oil and gas facilities. Proceedings of the Institution of Structural Engineers: Young Researchers Conference 11. London: IStructE Publications, 8. Arshi HS, Stone KJL and Newson TA. (211). Numerical modelling on the degree of rigidity at pile head for offshore monopile-footing foundation systems. 9 th British Geotechnical Association Annual Conference, London. Arshi HS and Stone KJL. (211). An investigation of a rock socketed pile with an integral bearing plate founded over weak rock. Proceedings of the 15 th European Conference of Soil Mechanics and Geotechnical Engineering. Amsterdam: Ios Pr Inc, Arshi HS. (212). A new design solution for increasing the lateral resistance of offshore pile foundations for wind turbines located in deep-water. Proceedings of the Institution of Structural Engineers: Young Researchers Conference 12. London: IStructE Publications, 1. Arshi HS and Stone KJL. (212). Lateral resistance of hybrid monopilefooting foundations in cohesionless soils for offshore wind turbines. Proceedings of the 7 th International Conference on Offshore Site Investigation and Geotechnics. London: Society for Underwater Technology, Bransby MF and Randolph MF. (1998). Combined loading of skirted foundations. Géotechnique. 48(5), Broms BB. (1964). Lateral resistance of piles in cohesionless soils. ASCE Journal of the Soil Mechanics and Foundation Division. 9(SM3), Duncan JM, Evans LT and Ooi PS. (1994). Lateral load analysis of single piles and drilled shafts. ASCE Journal of Geotechnical Engineering. 12(6), El-Marassi M, Newson T, El-Naggar H and Stone KJL. (28). Numerical modelling of the performance of a hybrid monopiledfooting foundation. Proceedings of the 61 st Canadian Geotechnical Conference, GeoEdmonton 28. Edmonton, (Paper No. 48), Gourvenec S and Randolph M. (23). Effect of strength nonhomogeneity on the shape of failure envelopes for combined loading of strip and circular foundations on clay. Géotechnique. 53(6), Houlsby GT and Puzrin AM. (1999). The bearing capacity of a strip footing on clay under combined loading. Proc. R. Soc. London Ser. A. 455, Kim JB, Singh LP and Brungraber RJ. (1979). Pile cap soil interaction from full scale lateral load tests. ASCE Journal of Geotechnical Engineering. 15(5), Maharaj DK. (23). Load-deflection response of laterally loaded single pile by nonlinear finite element analysis. EJEG. Matlock H and Reese LC. (196). Generalized solutions for laterally loaded piles. ASCE Journal of Soil Mechanics and Foundations Division. 86(SM5), Mokwa RL. (1999). Investigation of the Resistance of Pile Caps to Lateral Loading. Ph.D Thesis. Virginia Polytechnic Institute, Blacksburg, Virginia. Mokwa RL and Duncan JM. (21). Experimental evaluation of lateralload resistance of pile caps. ASCE Journal of Geotechnical and Geoenvironmental Engineering. 127(2), Mokwa RL and Duncan JM. (23). Rotational restraint of pile caps during lateral loading. ASCE Journal of Geotechnical and Geoenvironmental Engineering. 129(9), Poulos HG. (1971). Behaviour of laterally loaded piles: Part I-single piles. ASCE Journal of the Soil Mechanics and Foundations Division. 97(SM5), Poulos HG and Randolph MF. (1983). Pile group analysis: a study of two methods. ASCE Journal of Geotechnical Engineering. 19(3), Powrie W, and Daly MP. (27). Centrifuge modelling of embedded retaining wall with stabilising bases. Geotechnique. 57(6), Randolph MF. (1981). The response of flexible piles to lateral loading. Géotechnique. 31(2), Reese LC, Cox WR and Koop FD. (1974). Analysis of laterally loaded piles in sand. Offshore Technology Conference. Vol. II (Paper No. 28), Stone KJL, Newson TA and Sandon J. (27). An investigation of the performance of a hybrid monopole-footing foundation for offshore structures. Proceedings of 6 th International on Offshore Site Investigation and Geotechnics. London: SUT, Stone KJL, Newson TA and El Marassi, M. (21). An investigation of a monopiled-footing foundation. International Conference on Physical Modelling in Geotechnics, ICPMG21. Rotterdam: Balkema, Stone KJL, Newson TA, El Marassi M, El Naggar H, Taylor RN, and Goodey RA (211). An investigation of the use of bearing plate to enhance the bearing capacity of monopile foundations. International Conference on Frontiers in Offshore Geotechnics II - ISFOG. London: Taylor and Francis Group, Zhang L, Silva F and Grismala R. (25) Ultimate lateral resistance to piles in cohesionless soils. Journal of Geotechnical and Geoenvironmental Engineering. Vol. 131(1),

20 Influence of jack-up footprints on mudmat stability How beneficial are 3D effects? Influence des dépressions laissées par les jack-ups sur la capacité portante des mudmats quels sont les effets bénéfiques d une analyse en 3D? Ballard J.-C., Charue N. Fugro GeoConsulting Belgium ABSTRACT: Jacket platforms are piled into the seabed but need to be supported temporarily by mudmats during installation. They sometimes need to be located next to seabed features such as pug marks formed by previous deployments of jack-up rigs. These features may influence the bearing capacity of the mudmats. This is a 3D problem for which simplified approaches are unsatisfactory, simplified 2D plane strain simulations can lead to over-conservative results. This paper presents a project example in very soft clay for which the software package Plaxis 3D has been successfully used. The presence of a pug mark was found to degrade significantly the yield surface in the VHM load space. A comparison between 2D and 3D analyses shows that the beneficial 3D effects are substantial, especially when the pug mark is located at the corner of the mudmat. The zone of influence of the pug mark is also much more limited when the problem is modelled in 3D. RÉSUMÉ : Les platesformes de type «Jacket» sont fondées sur pieux mais nécessitent d être supportées temporairement pendant l installation par des mudmats (fondations de type superficiel). Ces jackets sont parfois situées à proximité de dépressions laissées par l installation antérieure de jack-ups. Ces dépressions peuvent influencer la capacité portante des mudmats. Il s agit d un problème 3D typique pour lequel aucune solution simplifiée n existe. Une approche 2D (en état plan de déformation) peut même mener à des résultats trop conservatifs. Cet article présente un exemple dans de l argile molle pour lequel la suite de logiciels Plaxis a été utilisée avec succès. Les conclusions sont les suivantes : la présence des dépressions modifie singulièrement la surface de rupture dans l espace VHM. Une comparaison entre les approches 2D et 3D montre que les avantages à faire appel au 3D sont substantiels, spécialement quand la dépression est située à proximité du coin du mudmat. La zone d influence de la dépression est aussi bien plus limitée lorsque le problème est modélisé en 3D. KEYWORDS: Pug mark, mudmat, stability, VHM, 2D, 3D, Finite Element Analysis, soft clay, remoulded, jack-up, mesh 1 INTRODUCTION Jacket platforms are the most common type of offshore structure in the offshore hydrocarbons industry (Dean, 21). They consist of open-framed steel structures made of tubular leg chords, horizontal bracing, and diagonal bracing. These structures are piled into the seabed but need to be supported temporarily by mudmats during installation. Mudmats are essentially flat stiffened metal plates attached to legs or the lower braces. In soft soils, mudmats can cover the entire surface between the legs to maximise the bearing area. They are generally subjected to combined Vertical, Horizontal and Moment (VHM) loads induced by the jacket weight, wind, waves and currents. Jacket platforms are not always installed on a virgin seabed and are sometimes located next to features such as pug marks formed by previous deployments of jack-up rigs. A jack-up is a mobile, self-elevating offshore platform consisting of a hull and three or more retractable legs passing through the hull (McClelland et al, 1982). A unit moves onto location, sets its legs onto the seabed, and raises its hull out of the water. The legs are supported on independent foundations called spudcans. Penetration and extraction of spudcans in soft grounds create zones of remoulded soil and seabed depressions (Hossain et al, 212). These seabed features potentially influence the bearing capacity of the mudmats and need to be accounted for in the stability verification. Mudmats subjected to combined VHM loads and located next to a jack-up footprint is a 3D problem for which simplified approaches for analysis do not exist. Simplified 2D plane strain simulations are generally performed but they can lead to overconservative results. This type of problem is better analysed by means of 3D Finite Element (FE) analyses. This paper presents a project example in very soft clay for which the software package Plaxis 3D (Plaxis, 211) has been used successfully. The analysis allowed confidence to be established for the selected location of the mudmat with respect to a pug mark. In contrast, a simplified 2D analysis suggested that the proximity of the mudmat to the pug mark was unacceptable. It is shown for this particular example how the presence of a pug mark degrades the yield surface in the VHM load space. 3D analyses are compared with 2D analyses to quantify the beneficial 3D effects for different pug mark locations. 2 PROBLEM GEOMETRY AND SOIL CONDITIONS A 3 m by 3 m square mudmat is considered. The mudmat is located next to a circular pug mark of 3 m in diameter. Analyses were performed for 2 positions of the mudmat. The first position considers a pug mark located along the width of the mudmat while the second position considers a pug mark located at the corner of the mudmat, as illustrated on Figure 1. The distance d between the edges of the mudmat and the pug mark is varied in the analysis. Soil conditions in this example consist of very soft clay. The soft deposit is considered to be 15 m thick and underlain by stiffer soils, which are not modelled. The intact undrained shear strength increases linearly with depth according to s u = 4 +.8z (in kpa), where z is the depth below ground level in meter. This gives a strength heterogeneity = 6 where = kb/s u, k is the 2311