Otc 19495

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    OTC 19495 Model Tests for Steel Catenary Riser in Marine Clay Thomas Langford, Norwegian Geotechnical Institute; Charles Aubeny, Texas A&M University Copyright 2008, Offshore Technology Conference This paper was prepared for presentation at the 2008 Offshore Technology Conference held in Houston, Texas, U.S.A., 5–8 May 2008. This paper was selected for presentation by an OTC program committee following review of information contained in an abstract submitted by the author(s). Contents of the paper have not been reviewed by the Offshore Technology Conference and are subject to correction by the author(s). The material does not necessarily reflect any position of the Offshore Technology Conference, its officers, or members. Electronic reproduction, distribution, or storage of any part of this paper without the written consent of the Offshore Technology Conference is prohibited. Permission to reproduce in print is restricted to an abstract of not more than 300 words; illustrations may not be copied. The abstract must contain conspicuous acknowledgment of OTC copyright.  Abstract The issue of fatigue damage caused by cyclic interaction of the steel catenary risers with the seabed has gained prominence with the widespread use and lengthening of the spans for this type of system. This paper presents the findings from a series of large-scale model tests of soil-riser interaction in re-constituted high plasticity marine clay from the Gulf of Guinea. Data are  presented on soil stiffness during virgin penetration, unload-reload stiffness as a function of displacement amplitude, the effects of soil-riser separation during robust load cycles, force-controlled versus displacement controlled load conditions, stiffness degradation under cyclic loading, and stiffness regain due to consolidation and thixotropy. Introduction Steel Catenary Risers (SCRs) are utilized to connect floating platforms with seabed systems and feature prominently in deepwater projects. Figure 1 gives the general arrangement of such a riser, where the ‘Touchdown Zone’ (TDZ) refers to the area where the riser is in ‘dynamic’ contact with the seabed. The interaction between SCR and seabed in the TDZ is of great importance when evaluating the structural fatigue life of the riser (Hatton, 2006); a stiffer seabed will result in greater localized stresses in the riser and vice versa. Recent work has investigated seabed-SCR interaction to improve interaction models that better capture the geotechnical behaviour within the structural analysis (e.g. Aubeny et al., 2006 and Clukey et al., 2005). Physical model testing has been an important tool to investigate seabed-SCR interaction in the TDZ, as reported by several authors including Dunlap et al. (1990), Bridge et al. (2004) and Giertsen et al. (2004). However, this work was based on test data for kaolin or low plasticity soils. The majority of deepwater projects are located in areas with clays of much higher  plasticity, such as the Gulf of Guinea, Gulf of Mexico and South China Sea. Andersen (2004) has demonstrated that the cyclic behaviour of clays is dependent on both plasticity index and overconsolidation ratio. The authors therefore decided to  perform seabed-riser interaction tests on a high plasticity soil taken from the Gulf of Guinea. Figure 2 illustrates several facets of soil-riser interaction behavior that were investigated. The first involves the mobilization of soil resistance during initial monotonic penetration of the riser into the seafloor. Upon completion of the initial penetration  phase, two limiting conditions of cyclic loading of the riser were investigated. The first involved force-controlled loading conditions in which the riser was unloaded and reloaded to a uniform level of compression resistance in each load cycle. The second involved displacement controlled loading in which the riser is reloaded to a uniform penetration depth. It is readily apparent from Figure 2 that fundamentally different soil-riser interaction will occur under displacement-controlled conditions, as soil resistance will decline with each successive load cycle. With regard to the relevance of loading mode to actual field conditions, either can be relevant depending on the condition being considered. During the trench formation  phase (which may occur repeatedly during the life of a riser), the riser successively embeds itself deeper with each load cycle, so a force-controlled mode would likely be a closer approximation to field conditions. In contrast, as the trench approaches a steady-state configuration, a soil-riser interaction behavior in a displacement-controlled mode is likely to be most relevant. Accordingly, soil-riser interaction behavior in both loading modes merit investigation.  2 OTC 19495    S  o   i   l   R  e  s   i  s   t  a  n  c  e ,       Q Penetration,  z Reversal at Constant ForceReversal at ConstantDisplacementInitialPenetration   Figure 1. SCR and Touchdown Zone   Figure 2. Typical Soil-Riser Interaction Behavior    Test Program The test program consisted of 4 test footprints, as presented in Table 1. Table 1 Summary of tests performed Test Dates Load/Displacement control Penetration velocity Description Cyclic stages 1 10 August Displacement 0.05 mm/s Cyclic penetration/extraction 1 2 10 to 13 August Load 0.05 mm/s Small-cycle load-controlled tests 6 3 14 to 15 August Displacement 0.05 mm/s Small-cycle displacement-controlled tests 13 4 17 August Displacement 0.50 mm/s Cyclic penetration/extraction 1 Tests 1 and 4 were single-stage cyclic tests used to investigate the bounding curves for penetration and extraction given in Figure 2. Tests 2 and 3 were multi-stage cyclic tests where each stage featured a different specified load or displacement history and/or time delay between the stages. These multi-stage small cycle tests were used to investigate the unloading and reloading stiffness within the bounding curves. Test Equipment The tests were performed in clay prepared within a steel bin of plan dimensions 3.6 x 1.7 m, as shown in Figure 3. A hydraulically-powered biaxial test system was used for the project, as shown in Figure 4. In this case only the vertical actuator was used for the testing for which the maximum test stroke is 1000 mm. The rig has been previously described in Langford et al. (2007). The pipe element used for the testing was a ‘rough’ coated element of length 1300 mm and diameter 174 mm. Figure 3. Test Arrangement in Plan View   Figure 4. NGI Test Bin and Instrumentation   Clay Preparation and Strength Evaluation The clay was recovered from the offshore field using a box corer and transported to Norway in bags at the natural water content of around 150%. Water was added to the clay in order to create a workable slurry with a water content around 340%. The clay was consolidated in the bin, first using dead weights and then a vacuum applied within a rubber membrane. The  OTC 19495 3 final consolidation stress was 9.5 kPa and the resulting settlement curve is presented in Figure 5. After consolidation and a 1 month swelling period, the final height of the clay for testing was just under 240 mm. 110100 Time, days 200220240260280300320340360    C   l  a  y   h  e   i  g   h   t ,  m  m Settlement data   0123456 Undrained shear strength, kPa    P  e  n  e   t  r  a   t   i  o  n ,  m Model 1T-bar 1T-bar 2T-bar 3T-bar 4   Figure 5. Settlement Curve for Clay   Figure 6. Undrained Shear Strength from T-bar Test  The resulting shear strength profile was investigated by means of 4 mini T-bars penetration tests (diameter  D  = 25 mm by length 120 mm). The test probe was penetrated and extracted at a constant rate of v  = 20 mm/sec and the undrained shear strength was interpreted from the T-bar resistance using a factor  N  T-bar   = 10.5 , giving the profiles in Figure 6. It can be seen that the shear strength increases linearly from ‘seabed’ to 0.2 m and model profile was defined increasing from 2 kPa at the seabed with a strength gradient of 13 kPa/m to be used in the subsequent interpretation of results (i.e. s u  = 2+13·z kPa). The interpretation of T-bar results should be treated with some caution since the T-bar factor increases during the first few diameters of penetration before reaching a steady value, as discussed in White and Randolph (2007). However, the change of T-bar factor with depth is a topic of research with some uncertainity and is therefore not taken into account here; this means that the surface shear strength may be underestimated. Due to dependence of undrained shear strength on strain rate, it is useful to express shear strength at a given velocity in terms of the shear strength at some reference velocity. Throughout this paper, the reference velocity was selected as the T-bar  penetration rate normalized by diameter, V  ref   = v/D  = 0.8 sec -1 . When interpreting measurements from subsequent riser tests conducted at various velocities, the s u  appropriate to a given riser velocity V   was adjusted for rate effects using the following relationship: s u  = s u,T-bar   [1 + λ  ref   log (V/V  ref  )] (1) where s u,T-bar   is the strength measured in the T-bar test, V  ref   is reference velocity, and λ ref   is a strain rate multiplier corresponding to V  ref  . Initial Penetration For each of the four tests, the pipe was initially penetrated to a depth of 52 mm (0.3 pipe diameters) at a constant rate of displacement. The first 3 tests were penetrated at 0.05 mm/s, whereas Test 4 was penetrated faster at 0.5 mm/s. Figure 7 shows a significant rate effect between these two penetration rates, where the faster test (Test 4) exhibits 15 to 20% higher resistance than the slower tests. This increase in penetration resistance over a log cycle of penetration rate is consistent with rate effect studies reported by Lunne and Andersen (2007) for laboratory shear strength measurements. Some scatter in soil resistance occurred even among the tests conducted at the same penetration rate, which may have been due in part to soil disturbance from the neighboring footprint.   Based on data from initial penetration, the strain rate multiplier was estimated as λ ref   = 0.11 for the marine clay tested. It is noted that the magnitude of λ ref   is linked to a specific reference velocity, in this case V  ref   = 0.8 sec -1 , and meaningful comparisons are possible only when the same reference velocity is considered. Figure 8 shows bearing factor  N   p  ( = P / s u ) as a function of depth for the four tests. The undrained shear strength s u  used in this calculation was obtained from Figure 6, with the correction made for rate effects using Eq. 1. Measured bearing factors were compared to two theoretical estimates. The first is a simplified expression from Murff et al. (1989) that approximates  plasticity based solutions for a rough pipe:  4 OTC 19495 2 2(2)(/)(/)  p  NzDzD π  = + +   (2)   The second is an empical expression matched to finite element solutions by Aubeny et al. (2005):  N   p  = a·(z/D) b   (3) where the coefficients a  and b  vary according to pipe surface roughness. Figure 8 shows the theoretical bearing factor for a rough pipe, for which a  = 6.73 and b  = 0.29. Measured apparent bearing factors somewhat exceed both theoretical estimates, with the most significant diferences occurring at 0.1-0.2  D . Of course, the discrepancy can be attributed to inaccuracy in the theoretical estimates. However, as noted earlier, the estimated soil strength near the mudline is a matter of some uncertainty, since the T-bar factor is influenced  by the proximity to a free surface. Furthermore, preparation of the clay in the test basin is believed to generate a crust of stiffer soil near the free-surface, which is not necessarily detected in the T-bar measurement due to the free surface effect. Accordingly, the actual soil strength at shallow depths is likely to be higher than indicated by Figure 6. If this is in fact the case, the apparent bearing factors in Figure 8 should be considered as upper bound estimates. 02468101200. Test 1, V   = 0.00029 sec -1 Test 2, V   = 0.00031 sec -1 Test 3, V   = 0.00031 sec -1 TEst 4, V   = 0.0021 sec -1    E  q  u   i  v  a   l  e  n   t   P  r  e  s  s  u  r  e ,      P    (   k   P  a   ) Penetration,  z/D   012345600. 1Test 2Test 3Test 4    B  e  a  r   i  n  g   F  a  c   t  o  r ,      N    p Penetration Depth,  z/D Aubeny et el. (2005)Murff et al. (1989)   Figure 7. Initial Penetration Resistance   Figure 8. Penetration Bearing Factor    Displacement Controlled Cyclic Tests In the displacement controlled test (Test 3), following initial pentration the riser was subjected to a series of 100-cycle load  parcels. The cyclic displacement during each load cycle was approximately 4 mm (0.02  D ). At the end of each parcel, the loading was interrupted to allow rest periods of 1 to 4 hours. Figure 9 shows the envelope values of measured soil resistance in compression and uplift. Figure 10 shows the corresponding normalized unloading secant stiffness versus cyclic displacement magnitude for selected load cycles, namely: (A) the first load cycle of the first parcel, (B) the 99 th  load cycle of the first parcel, (C) the first load cycle of the second parcel, and (D) the 99 th  load cycle of the second parcel. Secant stiffness in Figure 10 is normalized by maximum compression resistance as follows: K   = k  sec  /  N   p  s u  = ( Δ Q  / Δ  z ) /  N   p  s u  (4) where Q  is soil resistance force per unit length of pipe, and  z  is vertical displacement. Referring to Figures 9 and 10, the following comments apply: ã   As cyclic loading progresses the compression resistance declines by about half after 100 cycles (paths A to B in Figure 9). Figure 10 shows the secant stiffness degrading in a commensurate fashion, with the stiffness for cycle B being about 60% of that for cycle A. ã   After a 1-hr rest period some regain in soil resistance occurs, with about a 20% increase in compression resistance occurring between B and C in Figure 9. The secant stiffness again increases in a commensurate fashion over the rest  period, although the stiffness does not appear to increase in exact direct proportion to the increase in maximum compression resistance.
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