ARTICLE. Test fill on soft plastic marine clay at Onsøy, Norway. Toralv Berre

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1 30 Test fill on soft plastic marine clay at Onsøy, Norway Toralv Berre ARTICLE Introduction Abstract: The test fill at Onsøy, close to the town of Fredrikstad in Norway, was built on a very soft marine clay with in situ vane strength as low as 11 kpa and a plasticity index varying from 34 to 50. The dimensions at the bottom of the long fill were 20 m 60 m and the height 2.3 m. The fill, as placed, had a computed factor of safety against foundation failure of 1.35 based on in situ vane strength. The fill was allowed to sit for about 3 years (from 1972 to 1975) before it, in a second stage, was brought to failure by raising the height. The main purpose of the first stage was to study the stress strain time relationships for the soft clay and compare the values of geotechnical parameters determined by careful laboratory tests with those back-calculated from the observed field behaviour. The test fill foundation was heavily instrumented to measure clay deformations and pore pressures with time. Field observations showed that the Onsøy test fill turned out to be more of a drained loading case than an undrained case in spite of the fairly rapid construction of the fill. Actually, subsequent undrained triaxial and direct simple shear tests strongly indicate that the fill might have failed if less drainage had taken place, i.e., if the drainage paths had been longer and (or) if the coefficient of consolidation had been lower. Key words: test fill, Soft clay, field measurements, triaxial tests, in situ vane tests. Résumé : Le site d essai de remblayage de Onsøy, près de la ville de Fredrikstad en Norvège, a été construit sur de l argile marine très molle dont la résistance scissométrique est aussi faible que 11 kpa et dont l indice de plasticité varie de 34 à 50. Les dimensions au bas du remblai sont de 20 m 60 m et la hauteur est de 2,3 m. Le remblai, tel que placé, comporte un facteur de sécurité calculé contre les ruptures de fondation de 1,35, basé sur la résistance scissométrique in-situ. Le remblai a été laissé en place pendant environ 3 ans (de 1972 à 1975) avant qu il n atteigne la rupture lors d une deuxième étape de rehaussement. L objectif principal de la première étape était d étudier les relations contraintes-déformations-temps pour l argile molle, et de comparer les valeurs des paramètres géotechniques déterminés à partir d essais de laboratoire à des valeurs rétro-calculés à partir d observations du comportement sur le terrain. La fondation du site remblayé a été instrumentée avec de nombreux appareils pour mesurer des déformations dans l argile et les pressions interstitielles avec le temps. Les observations sur le terrain ont démontré que le site de remblayage de Onsøy s est avéré un cas de chargement drainé plutôt qu un cas non drainé, malgré la construction relativement rapide du site. En fait, des essais triaxiaux non drainés et de cisaillement simple directs subséquents ont fortement indiqué que le remblai aurait pu céder si moins de drainage avait eu lieu, c est-à-dire si les chemins de drainage avaient été plus long et/ou si le coefficient de consolidation avait été plus faible. [Traduit par la Rédaction] Mots-clés : site d essai de remblayage, argile molle, mesures de terrain, essais triaxiaux, essais scissométriques in-situ. In 1966 the Norwegian Geotechnical Institute (NGI) was engaged by Danish consulting firm Kampsax to assist with the investigation of problems with two highway projects in Thailand where test fills on the soft and very plastic Bangkok clay failed. These failures happened even though the calculated factor of safety based on in situ vane tests was about 1.5 (Eide and Holmberg 1972; Eide 1978). Bjerrum (1973) therefore suggested to use a correction factor on the in situ vane strengths and indicated that this, for the Bangkok clay, was primarily a correction for the rate effect on the undrained shear strength. Stability calculations based on standard triaxial and direct simple shear (DSS) tests also showed that these strength values needed a correction, which was assumed to be mainly a rate correction, to get reasonable factors of safety for the Bangkok clay (Holmberg 1972, 1974). The experience with the soft Bangkok clay, together with a general interest in studying the mechanical properties of plastic clays, led to NGI s decision to build the test fill at Onsøy in The main purpose for this test fill was to study the effect of time on strength and deformation characteristics for approximately undrained conditions. The fill was laid out with a factor of safety against foundation failure of about 1.35 according to the in situ vane strength. The first stage of the test fill was expected to give large deformations under approximately undrained conditions, but not failure. After 3 years, the fill was brought to failure in about 1 week by raising the height. The failure stage will be presented in a separate paper. Site conditions The test fill is located at Onsøy about 4 km north of the centre of Fredrikstad (Fig. 1). NGI has previously carried out an extensive series of vane tests with different shapes of the vane in the same area to study strength anisotrophy. These tests are reported by Aas et al. (1986). The terrain around the test fill slopes gently (1:200) towards the Seut River, which is at about sea level. During the test period, the elevation of the natural ground surface under the fill was about +0.7 m. When the tide was especially high, the water from the river rose about 0.1 m above the ground surface around the fill, but the average ground water level was about 0.2 to 0.3 m under the ground surface. The pore pressure was slightly artesian, and at a depth of 20 m the pressure was about 5% higher than the fresh Received 21 December Accepted 16 October T. Berre. Norwegian Geotechnical Institute, Sognsveien 72, P.O. Box 3930, Ullevaal Stadion, 0806 Oslo, Norway. for correspondence: toralv.berre@ngi.no. Can. Geotech. J. 51: (2014) dx.doi.org/ /cgj Published at on 23 October 2013.

2 Berre 31 Fig. 1. Map showing the Onsøy test field in the years 1972 to water and about 3% higher than the natural salt water hydrostatic pressure. The depth to bedrock in the area varies from about 22 to about 53 m. Below the middle of the fill the depth to bedrock is about 53 m. The bedrock is probably covered by a relatively thin layer of bottom moraine. Above this layer there is clay up to the ground surface. Micropaleontological investigations had been performed in connection with a slide that took place 7 m above sea level about 6.3 km north of the test fill. According to Hartmark (1962), the micropaleontological investigations showed that the clay is postglacial down to a depth of about 14 m. Below a depth of about 17 to 18 m there is Yoldia clay that was deposited to years ago. This description will probably also apply approximately for the ground below the test fill. The thickness of the weathered crust below and around the test fill varies from about 0.8 to 1.5 m, decreasing towards the river. Down to a depth of about 8 to 10 m there are black spots in the clay and pores with humus components. The black spots are about 2 10 mm wide and the distances between them are of similar magnitude. The spots probably indicate high concentrations of iron sulfur. In some of the samples horizontal and vertical cracks were discovered. Below a depth of about 10 m the colour of the clay is grey with no black spots. Figure 2 shows a boring profile. The laboratory data were determined on 54 mm samples from a borehole close to the fill (as shown in Fig. 3). It is seen that the natural water content varies between 57% and 67%. The average value of the plasticity index varies from about 34 in the upper 9mtoabout 44 between 9 and 15 m and to about 51 between 15 and 20 m. The average value of the undrained shear strength, determined by in situ vane tests, is about 11 kpa just below the weathered crust. At a depth of about 10 m the average vane strength is about 17 kpa. Instrumentation under the test fill The fill was 60 m long and 20 m wide (at the ground surface). It was believed that there would be approximately plane strain conditions under the middle third of the fill, and therefore most of the instrumentation was placed within this zone. The instrumentation at large depths was placed as near to the middle of the fill as possible, because the length of the zone with plane strain decreases with depth. The instrumentation outside the middle third was placed mainly along the longitudinal axis of the fill. A plan of the instrumentation is shown in Fig. 4. The electrical (vibrating wire) and hydraulic piezometers were placed at different sections, because the hydraulic piezometers were also used to determine the horizontal stress by hydraulic fracturing as described by Bjerrum and Andersen (1972). Such measurements create cracks that one did not want to have in the zone where the potential failure surface under the fill might pass through. The following instrumentation is in addition to what is shown in Fig. 4: 1. Ring magnets along a horizontal plastic tube in the fill, for measurements of horizontal deformations. The tube was placed 0.5 m above the original ground surface at the middle of the fill, perpendicular to the longitudinal axis. 2. Fifty wooden poles in the area around the fill. On top of each pole stick a plate (of copper) with a cross mark was mounted. The distances between the poles were measured, and geodetic surveys recorded the vertical movements of the cross marks. The poles were driven down to a depth of about 0.6 m. 3. Brittle sticks of wood in the ground for determination of any developing failure surfaces. Building up the fill Before the fill was laid out, trenches were excavated through the weathered crust to reduce the effect of the strength of the

3 32 Can. Geotech. J. Vol. 51, 2014 Fig. 2. Boring profile from 54 mm tube samples. Observed cracks are probably due to the sampling process. Positions of the boreholes are shown in Fig. 3. Depth, m Type of soil Root-holes Sample Nº Water content (%) Density (Mg/m 3 ) Undrained shear strength (kpa) Humus layer, thickness 10 mm Clay, grey with black spots Clay, grey Sandgrains and shells Shells Sandgrains 14 Cracks 15 Horizontal crack 16 Horizontal crack 17 Vertical 18 crack 19 Key: Water content Liquid and plastic limit Unconfined compression test Fallcone, undisturbed Fallcone, remoulded Field vane Average value 1.66 Average field vane crust on deformations and stresses below the crust. The trenches were made 0.7 to 0.8 m deep and about 0.15 m wide. A plan of the trenches is shown in Fig. 3 together with the dimensions of the fill. They were excavated mainly parallel to the axis of the test fill, plus one at each end. The trenches under the fill were filled with sand while those outside the fill were left open. The material in the test fill was sand laid out in five layers. The thickness of the first four layers was about 0.5 m while the thickness of the 5th layer was about 0.25 m. Measured deformations and stresses below the fill Figure 5 shows vertical settlements below the middle of the fill, as measured by the precision settlement gauge at point A, plotted versus the logarithm of time. Filling started on day No. 20 and was completed on day No. 35. Deformations were then allowed to take place until day No. 1116, i.e., for about 3 years, before the fill was loaded to failure. The curve showing settlement versus log time does not show any signs of levelling out during this period. Figure 6 shows load from the fill and measured vertical and horizontal displacements and excess pore pressure (i.e., change in pore pressure) plotted versus day number for various points below the fill. Figure 7 plots measured vertical and horizontal displacements versus depth at day No. 37 (i.e., 2 days after the end of filling) and at day No. 102.

4 Berre 33 Fig. 3. Plan showing the dimensions of the fill, trenches through the weathered crust (depth 0.75 m, width 0.15 m), and positions of the 54 and 95 mm boreholes where piston samples were taken prior to the filling. Figure 8 shows vertical displacements for points along the longitudinal axis of the fill. These displacements were measured just below the weathered crust (at depths varying from 0.5 to 1.5 m) at day Nos. 555, 791, and 1065, i.e., about 1.4, 2.1, and 2.8 years after the fill had been placed. Figure 9 shows vertical displacements along an axis normal to the longitudinal axis at the middle of the fill just below the weathered crust. Figure 10 plots vertical strain along the vertical centreline under the centre of the fill at various times. The strain values have been obtained from measurements on ring magnets that are floating in the clay outside a plastic tubing. The positions of the magnets are accurately measured by a detector that is lowered inside the tubing. For example, in Fig. 10 there are results from six ring magnets with distances H1, H2, and so on between them. The average strain for each layer is computed as the compression between the magnets, in percent, of the original thickness of the layer. It is plotted as a horizontal column at the original position of the centre of the layer although the centres are moving downwards as the clay under the fill is compressed. The strain values are plotted versus initial (not true) depths below the original ground surface. A continuous curve is then drawn so that the area below the curve for each column is approximately the same as below the column. Figure 10 shows the tops of the columns in addition to the continuous curves. Figure 11 shows the same type of plot for a set of ring magnets 5 m east of the centreline of the fill. Figure 12 shows plots of horizontal displacements 5 m west and 5 m east of the vertical centreline versus true (i.e., current) depth below original ground surface at various times. The displacement value, in millimetres, divided by 50 is equal to the average horizontal strain, in percent, for the clay mass between the point where the horizontal displacement is measured and the middle of the fill. At day No a 0.5 m thick layer of sand was placed on top of the fill as a start of the incremental loading for the failure stage. The measurements at day No should have been taken before placing this layer; however, this layer does not seem to have had any significant influence on the horizontal displacements at that time. Figure 13 plots the ratio between horizontal and vertical strain ( H / V ) versus depth (i.e., initial depth below original ground surface) at day No. 36 and at day No The H -values are the average strain values over a 10 m wide distance below the middle of the fill. For an undrained plane strain condition one would expect this ratio to be close to 1 (no volume change) and then gradually decreasing with time as drainage takes

5 34 Can. Geotech. J. Vol. 51, 2014 Fig. 4. Instrumentation for measuring displacements and pore pressures. All dimensions in metres. place. The curves in Fig. 13 indicate that consolidation has taken place due to high initial permeability, especially in the weathered crust, down to a depth of about 3 m followed by some decrease in permeability in this depth interval as consolidation takes place. Figure 14 plots excess pore pressures below the centre of the fill versus depth (i.e., initial depth below original ground surface) at various times. The excess pore pressures (i.e., change in pore pressure since start of loading) have been corrected for changes due to the settlements of the piezometer tips (change in the excess pore pressure, corrected for the settlement of the piezometer, is equal to the change in elevation of the top of the water level in the piezometer). Figure 15 shows the same type of plot along a vertical line, 5 m west of the centre of the fill. Figure 16 plots excess pore pressures at various times below the middle of the fill at various distances from the central axis at an initial depth of 8 m below the original ground surface. It should be noted that the excess pore pressures increased for several days after end of loading before dissipation started. Values of in situ minor principal stress may be obtained by induced hydraulic fracturing around hydraulic piezometers using the method described by Bjerrum and Andersen (1972). Hydraulic fracturing tests performed prior to building up the test fill gave K o -values (i.e., ratio between horizontal and vertical effective stress) equal to 0.60 ± 0.05 in the depth interval 2 to 20 m. Using

6 Berre 35 Fig. 5. Settlement of ground surface below the middle of the fill as measured by a field precision settlement gauge. Fig. 6. Measured displacements and excess pore pressures versus time. this method, Fig. 17 shows plots of changes in total and effective minor principal stresses, below the centre of the fill at day Nos. 50, 476, and Figure 18 shows the same type of plot at 5 m west of the centre of the fill, but here 3 is not horizontal. Description of laboratory tests The routine laboratory tests on Onsøy clay referred to in Fig. 2 were performed in 1972 on 54 mm tube samples, while subsequent triaxial, DSS, and oedometer tests in 1973 were performed on 95 mm tube samples. The positions of the borings are shown in Fig. 3. The field vane curve shown in Fig. 2 is the average of 11 borings at and around the test fill location. Later (in 1985) advanced laboratory tests were performed on special block samples from a boring at the same field in Onsøy (Lacasse et al. 1985). Those samples proved to be of much better quality than the 1973 samples. In 2009 it was therefore decided to take a new series of block samples from boring Block 1 about 110 m northwest of the original test fill site (which is now covered by fill masses and buildings). A vane boring close to boring Block 1 gave shear strength values very close to the average values from the earlier vane borings around the fill, and the level of the ground surface at these new borings is very close to that of the fill prior to the filling. Figure 19 shows shear strength values from standard undrained triaxial and DSS tests performed on block samples from boring Block 1. The test specimens were consolidated anisotropically to the same effective stresses as in the field before the test fill was placed. The K o -value was taken to be 0.60, i.e., the value determined by hydraulic fracturing tests in the field. For the DSS tests the specimens were loaded up to 0.8 times the apparent preconsolidation stress p c, then unloaded to the present effective vertical stress in the field, v0, to obtain a realistic horizontal effective stress before start of shearing. The specimens were then sheared undrained with a rate of shear strain of 1.0% and 0.8% per hour for triaxial and DSS tests, respectively. Some key data for the drained laboratory tests are given in Table 1. Note that in addition to the C c -values given in Table 1, values of constrained modulus, M, are also plotted for a values up to where a is kept constant overnight. That is done in Figs. 20 to 23. Figures 20 to 23 show the results of constant rate of strain (CRSC) oedometer tests performed as described by Sandbækken et al. (1986) on block samples from Block 1 from depths:of 1, 3.9, 7.5, and 10.8 m. The oedometer tests were, except for the periods where the stress was kept constant, performed with a CRSC of 0.5%/h to 0.6%/h. The stress was, except for the test on the deepest specimen, increased until an axial strain of about 10%, then kept constant overnight, reduced to an OCR of about p c / v0, kept constant there for about 4 h, and finally increased until the strain was at least 23%. The test on the deepest specimen (Bamberg 2009) was performed with about the same rate of loading as for three other tests, but it was loaded to 16% axial strain before being unloaded to an OCR of The coefficients of permeability, k, have been back-calculated from the measured excess pore pressure at the undrained bottom of the specimen by the equations given by Wissa et al. (1971). The k-values are plotted, in a logarithmic scale, versus strain. The plots may be approximated by straight lines. Figs. 20 to 23 also show results of K o -triaxial tests together with the results of the oedometer tests at depths 1.2, 3.9, and 7.5 m, following mainly the same procedures as described by Berre (1982). The specimens for the K o -triaxial tests were first consolidated anisotropically to the in situ effective stresses as described above for the undrained tests. Then the axial and radial stresses were increased once a day so that the specimen area was kept as constant as possible. For the tests at depths of 1.2 and 3.9 m the daily increase in axial stress was 7%, while for the test at 7.5 m the daily increase was 3.5%. The K o -tests were performed with a back pressure equal to 196 kpa. For the triaxial tests constant head permeability tests were performed at stresses close to v0, p c, and at the end of the test. In addition to the tests described above, drained extension triaxial tests with constant axial effective

7 36 Can. Geotech. J. Vol. 51, 2014 Fig. 7. Measured distribution of vertical and horizontal displacements below the fill. Fig. 8. Settlements along the longitudinal axis of the fill just below the weathered crust.

8 Berre 37 Fig. 9. Settlement just below the weathered crust at the middle of the fill. Fig. 10. Vertical strain contours below the centreline of the fill (ring magnet tube No. 72). stress and increasing radial effective stress have also been performed. The results of those tests will be used in future numerical analyses. Preliminary analysis of field results Finite element analyses of the clay foundation behaviour are being performed, considering partial consolidation and creep. Procedures and results of those will be described in a separate paper. In this paper preliminary analyses are presented considering only the completely undrained and fully drained conditions. The analyses give a basis for comparing the values of geotechnical parameters obtained on samples in the laboratory with those derived from observed field behaviour. Undrained stability analysis Table 2 presents effective shear strength parameters from undrained triaxial tests on specimens consolidated anisotropically to the present effective stresses in the field. Figure 24 shows the most critical slip circle, obtained by the limit equilibrium method, when using the vane strength measurements shown in Fig. 2. The strengths used down to 1 m depth are shown with a dashed line in Fig. 2 because the strengths were only measured below 1 m depth. Outside the fill the circle is seen to end up in a pre-excavated trench where the strength is zero. The factor of safety when using the vane strengths for this circle was found to be If the vane strength is corrected by the combined correction for anisotropy and rate effects as proposed by Bjerrum (1973),

9 38 Can. Geotech. J. Vol. 51, 2014 Fig. 11. Vertical strain contours 5 m east of the centreline of the fill (ring magnet tube No. 53). Fig. 12. Horizontal displacements 5 m west and 5 m east of the centreline.

10 Berre 39 Fig. 13. Ratio between horizontal and vertical strain versus depth below the centreline of the fill. The strains are calculated from field measurements. the factor of safety is reduced from 1.35 to The shear strength of the sand fill was found from drained compression triaxial tests assuming that the friction angle that will be mobilized in the fill is equal to the minimum angle found from the tests at large strains, i.e., equal to If the strength values from the triaxial and the DSS tests presented in Fig. 19 are used for the slip circle in Fig. 24, a factor of safety of 1.14 is obtained. If it is required that all shear resistance values along the slip circle be defined at the same magnitude of shear strain, the factor of safety for the slip circle in Fig. 24 is reduced to The mobilized shear strain is then equal to 9%. For the middle third of the fill, plane strain triaxial tests will be more relevant than axially symmetric tests. According to tests reported by Ladd et al. (1971) on re-sedimented Boston Blue Clay, the ratio between axisymmetric and plane strengths were 0.97 and for compression and extension triaxial tests, respectively. If such ratios are taken into account for the slip circle in Fig. 24, the computed factor of safety will increase from 1.06 to According to Berre and Bjerrum (1973), the undrained shear strength from compression triaxial tests on plastic Drammen clay when time to failure is about 10 weeks (i.e., about 10 5 min) is about 15% lower than found by standard tests where time to failure is about 2 h. The plasticity index is about the same for the plastic clay from Drammen and the Onsøy clay down to a depth of about 8 m. If the long-term undrained strength for the Onsøy clay is assumed to be 15% lower than for tests performed with conventional laboratory rates, the calculated factor of safety will be reduced to 1.10(0.85) = According to a summary given by Lunne and Andersen (2007), the rate effect on the undrained shear strength is likely to be about the same for triaxial (compression and extension) and DSS tests. Another factor that has to be considered is the temperature effect on the undrained shear strength. Preliminary tests on the Onsøy clay indicate that the strength at ground temperature, i.e., at about 7.5 C, may be about 10% higher than the strength measured at room temperature. This will increase the calculated factor of safety to 0.94(1.10) = This means that according to the laboratory tests, the ground below the Onsøy fill would have been very close to failure if no drainage had taken place. It might have failed if the test fill build-up had been more rapid (it took 15 days), the maximum drainage path down to the critical slip surface had been longer (more than about 4.5 m), and (or) the minimum values of the coefficients of consolidation had been lower (lower than about m 2 /s). This view is supported by Figs. 14 and 15, which show that the excess pore pressure for depths smaller than about 10 m continues to increase up to about 30 days after end filling. This is considered to be due to creep strains, which is a type of strain that will take place with time at constant effective stresses. This type of strain will also take place when the effective stresses change, and in this case leads to a temporary increase in excess pore pressure after end of filling. In this connection it should be noted that the coefficients of consolidation from oedometer tests just above the apparent preconcolidation stress, p c, is about 10 times lower for the Bangkok than for the Onsøy clay. Therefore the increase in computed factor of safety at failure with increasing plasticity, as reported by Bjerrum (1973), is not only due to the increasing rate effect on the undrained strength, but also due to the decreasing amount of drainage with increasing plasticity. Settlement calculations The strain contours in Figs. 10, 11, and 12 and the curves in Fig. 13 show that the horizontal strains in the field are relatively small compared to the vertical ones below the longitudinal axis of the fill. For an undrained plane strain condition the two strains should be approximately equal. Below the longitudinal axis the horizontal strains probably are even smaller compared to the vertical ones shown in Fig. 13. It was therefore decided to base the drained interpretation on oedometer and drained triaxial tests, and limit the analysis to what is going on under the longitudinal axis of the fill. Figure 25 shows a plot of vertical strain versus depth, calculated from oedometer tests, as the difference between the strain at v0 + ( V ) final according to the 24 h virgin compression line minus the strain at v0 (this implies that in Fig. 25, in the three figures introduced in the section titled Comparison of stress strain-time relationships obtained from measurements in the field and from laboratory tests, and in Table 2 all strain values are zero at v = v0 ). The increase in total vertical stress with depth when the fill is completed, ( V ) final, has been calculated from charts from the theory of elasticity for a strip load, as presented by Lambe and Whitman (1979). Figure 25 also shows strain values obtained from the three K o -triaxial tests, which are seen to agree quite well with the values obtained from the oedometer tests. The strain values from the K o -triaxial tests are all after 24 h and therefore directly comparable with the values from the 24 h virgin compression line for the oedometer tests. The deepest oedometer test is for a sample from a depth of 10.8 m. Below this depth the plot is extrapolated down to about 18 m. The strain versus depth curve from the oedometer tests in Fig. 25 should represent the final strain, final, when all excess pore pressures have dissipated. The distribution of final corresponds to a total settlement at the ground surface of 1.2 m while the observed settlement at day No is about 0.68 m. The difference between 1.2 and 0.68 m is interpreted to be mainly due to incomplete consolidation. Below, an attempt has been made to estimate the degree of consolidation, U, at day No. 1116, assuming one-dimensional consolidation. All retardation in the compression is assumed to be

11 40 Can. Geotech. J. Vol. 51, 2014 Fig. 14. Excess pore pressure below the centreline of the fill, according to the hydraulic piezometers, corrected for settlement of the piezometers. due to expulsion of pore water, i.e., all secondary compression is included as primary compression. The coefficient of consolidation, c V = Mk/ w, where M =( V ) final / final. The values used in the computations are given in Table 3. The average values of the coefficient of permeability, k oed, and the c v -values, calculated as explained above, are m/s and m 2 /s, respectively. For one-dimensional consolidation, the relationship between time factor T, time t, and length of maximum drainage path H, is as follows: (1) T tc v H 2 When c v is set equal to m 2 /s, t to s (corresponding to day No. 1116), and H to 18 m, the value of T becomes When converting this into degree of consolidation, U, a solution given by Janbu (1970) is used where the distribution of strain with depth is triangular with maximum strain and free drainage at the top of the consolidating layer. With this solution, a T-value of gives a U-value of 0.31, i.e., a total compression of 1.20(0.31) = 0.37 m, which is much smaller than the observed value of about 0.68 m. It is seen from Fig. 20, and especially from Fig. 21, that k-values back-calculated from excess pore pressures for CRSC oedometer tests and represented by the so called k-lines are much smaller, especially at low stresses, than directly measured k-values by constant-head permeability tests (large dots) on triaxial test specimens. A possible reason for this may be that the clay contains remnants of vertical root holes. Such channels may not significantly speed up rate of consolidation for thin oedometer specimens, but may be very important for the much longer drainage paths in the field. Attempts have therefore been made to calculate the coefficient of permeability from pore pressure and deformation measurements in the field. A block of clay under the central part of the fill was used when doing this calculation. The depths to the top and the bottom of the block were 3.5 and 8.5 m, respectively. The widths perpendicular to and parallel with the long axis of the fill were 10 and 1 m, respectively. Deformations and pore pressure differences parallel with the long axis of the fill were assumed to be zero (i.e., plane strain conditions). The time interval considered was from day No. 66 to day No The deformation measurements tell how much the volume of the block changes per unit of time. The slopes of the excess pore pressure curves (in vertical and horizontal sections) give information about the hydraulic gradients at the boundaries of the block. The coefficient of permeability, k, was assumed to be the same in the vertical and horizontal directions. This computation gave a k-value of

12 Berre 41 Fig. 15. Excess pore pressures 5 m west of the centreline, according to the hydraulic piezometers, corrected for settlement of the piezometers. Fig. 16. Excess pore pressures 8 m below the original ground surface, according to the hydraulic piezometers, corrected for settlement of the piezometers.

13 42 Can. Geotech. J. Vol. 51, 2014 Fig. 17. Change in total ( 3 ) and effective ( 3 ) minor principal stress below the centreline of the fill determined by hydraulic fracturing tests on day Nos. 50, 476, and Fig. 18. Change in total ( 3 ) and effective ( 3 ) minor principal stress, 5 m west of the centreline of the fill, determined by hydraulic fracturing tests. (The minor principal stress is not horizontal outside the centreline of the fill.) k field = m/sec, which is about 3 times higher than the laboratory oedometer k-line values in this depth interval. With an increase of 3 in the k-values, the T-value increases from to and U increases from 0.31 to 0.47, i.e., to a settlement of about 1.20(0.47) = 0.56 m. Figure 26 shows time settlement curves calculated as explained above together with the observed curve. The principal sketch in Fig. 27 summarizes the main assumptions made in the drained analysis. As mentioned previously, all retardation of the settlement is assumed to be due to the delay associated with the expulsion of pore water, and the relation between effective stress and strain is assumed to be linear. The slope of this line is used to calculate the c v -values. The true stress strain curve in the field will be curved, due partly to the nonlinearity of the stress strain relationship and partly to secondary compression taking place during and after the expulsion of pore water. However, as seen from Figs. 20 to 23, the final effective stress, i.e., v0 +( V ) final, for most of the oedometer tests is considerably higher than p c. The secondary compression taking place overnight (before start of unloading) is seen to be relatively small compared to the total amount of strain, final, due to ( V ) final. Thus, assuming that the retardation of the settlement is due mainly to expulsion of pore water may be a fairly reasonable assumption below the centreline of the test fill. Although consolidation in this case appears to be quite dominating already from the start of the loading, there will always be an undrained shear component in addition to the volumetric consolidation, which also implies horizontal displacements. An estimate of this initial undrained settlement component based on the secant modulus halfway to failure for CAUC triaxial tests, gives an additional settlement of at least 0.05 m. The total computed settlement at day No then becomes equal to = 0.61 m, which is only slightly smaller than the observed value of 0.68 m. Comparison of stress strain time relationships obtained from measurements in the field and from laboratory tests In Fig. 28 vertical strain from the oedometer tests is calculated as the difference between the strain at v0 +( v ) 1116 (according to the laboratory 24 h virgin compression line) and the vertical strain at v0. The values of ( v ) 1116 are calculated by the theory of elasticity, using the measured excess pore pressures at day No The measured vertical field strains are also shown in Fig. 28. The strains calculated from oedometer tests are seen to be considerably smaller than the measured ones down to a depth of about 6 m. The main reason for this is believed to be that horizontal strains in the field, although small compared to the vertical ones, are still large enough to increase the vertical strains significantly. It may also be that the real horizontal deformations in the field are somewhat larger than the ones observed by the inclinometer channels because of the very abrupt change in stiffness and strength of the clay just below the weathered crust. Below a depth of about 7.5 m the calculated strain is larger than the values measured in the field, probably due to sample disturbance, which increases with increasing depth. It should also be remembered

14 Berre 43 Fig. 19. Undrained shear strength from triaxial and DSS tests versus depth. Figures written beside the plotted points, in parentheses, are volumetric strains (in %) at end of consolidation. Values below 1% are usually considered as very good, indicating only little sample disturbance. According to the classification system introduced by Lunne et al. (2006), based on e/e i, the triaxial test specimens come in class 1 (i.e., very good to excellent) while the DSS specimens come in class 2 (i.e., good to fair). Table 1. Some key data for drained laboratory tests performed on material from boring Block 1. Test No. Type of test Sample No. Depth (m) v0 w i (kpa) (%) s (g/cm 3 ) e i p c (kpa) OCR C c C s C Min. value of c v above Temp. p c (m 2 / S ) ( o C) ac (kpa) Confined by membrane Oed 1x CRSC 3-A1-Ø Oed 2x CRSC 10-A2-Ø Oed 3x CRSC 19-A1-Ø Oed 4x CRSC 26-B T1x CAK o UC 3-B1-T Yes T2x CAK o UC 10-A3-T Yes T3x CAK o UC 19-A2-T Yes T4x Creep test 13-A1-T None Old T5x Creep test 13-A2-T Yes Old T6x Creep test 13-A4-T Yes New FE program used to calculate stresses Note: CRSC, constant rate of strain oedometer test; CAK o UC, triaxial test with anistropic consolidation to v0 followed by K o consolidation and then undrained shearing in compression; FE, finite element.

15 44 Can. Geotech. J. Vol. 51, 2014 Fig. 20. Results of oedometer test: Oed 1x, depth = 1.01 m, v0 = 10.1 kpa. ( v ) final = 47.3 kpa. Also shown in the first figure is the K o -part of triaxial test T1x, depth = 1.23 m. Fig. 21. Results of oedometer tests: Oed 2x, depth = 3.87 m, v0 = 28.6 kpa. ( v ) final = 45.5 kpa. Also shown in the first figure is the K o -part for triaxial test T2x, depth = 3.91 m. Fig. 22. Results of oedometer test: Oed 3x, depth = 7.45 m, v0 = 50.6 kpa. ( v ) final = 43.0 kpa. Also shown in the first figure is the K o -part of triaxial test T3x, depth = 7.48 m. that the vertical strains associated with the calculated initial settlement come in addition to the oedometer strains given in Fig. 28. The curves in Fig. 28 indicate that the reason why the estimated time settlement curve in Fig. 26 agrees fairly well with the values observed in the field is, to some degree, due to compensating errors. The calculated oedometer strains at small depths are too small (making the c v -values too high). This is counteracted by the fact that horizontal drainage and possible very high permeability of the weathered crust are both neglected in the computations, and that the calculated strains from oedometer tests at large depths are too large due to sample disturbance. Figure 29 shows the results of three drained triaxial tests. These specimens were loaded, with respect to effective stresses, in the

16 Berre 45 Fig. 23. Results of oedometer test Oed 4x, depth = m, v0 = 69.9 kpa. ( v ) final = 37.6 kpa. Table 2. Effective shear strength parameters from undrained triaxial tests. At a r Type of test At a a r max r max CAUC = = 24.2, c= = 5.3 kpa = = 29.8, c= = 3.3 kpa CAUE = = 23.9, c= = 1.9 kpa == 39.6, c= = 0 kpa Note: Value of parameter within denotes numerical or absolute value of parameter. CAUC and CAUE denote triaxial tests with anistropic consolidation to v0 followed by undrained shearing in, respectively, compression and extension. same way and with the same rate as an element in the field below the centreline of the fill at a depth of 5.2 m. As mentioned previously, the total minor principal stress along the vertical axis of the fill was determined by hydraulic fracturing (Bjerrum and Andersen 1972). However, such measurements were only done at day Nos. 2, 50, 475, and To get an estimate of how the horizontal total stress changed between day No. 2 and day No. 50, which includes the filling period from day No. 20 to day No. 35, the total stresses in the field were computed with a finite element program assuming no drainage and a bilinear stress strain curve based on undrained triaxial tests on 95 mm tube samples. The total horizontal stresses under the centreline of the fill, measured at day No. 2 and day No. 50, were assumed to be the correct ones, and the horizontal stresses at day No. 36 (i.e., one day after the end of filling) were assumed to be equal to those at day No. 50. The finite element analysis was used to estimate the values in-between day No. 20 and day No. 36. The total vertical stress values were calculated by the finite element program. Effective stresses were obtained by subtracting the measured pore pressures from the calculated total stresses. These effective stress paths were followed for tests T4x and T5x, as shown in Fig. 29. Other testing conditions for tests T4x and T5x were as follows: Test T4x: Cylindrical specimen, the radial effective stress was equal to the minor effective principal stress in the field. The test was done without rubber membrane using the paraffin method described by Iversen and Moum (1974). The testing temperature was about 20 C. Test T5x: As test T4x except that the specimen was enclosed by a rubber membrane. Test T6x was as T5x, but performed at a temperature of 7.8 C, i.e., close to the in situ ground temperature. For this test the total stresses in the field were computed by Gustav Grimstad, NGI, by a more advanced finite element program that includes partial drainage and nonlinear stress strain curves from tests performed in 2010 on block samples from boring Block 1. The measured total horizontal stresses and the measured pore pressures were used in combination with the finite element computations in the same way as for the finite element computation described above, except that the finite element program was used to calculate the stresses also between day No. 36 and day No. 50. Figure 29 also shows vertical strains under the centreline of the fill and average horizontal strains over a 10 m wide zone under the central part of the fill, both obtained from the field measurements. These strain values have been obtained from Figs. 10 and 12. Figure 30 shows vertical strains versus logarithm of time. The following is seen from the plots in Figs. 29 and 30: 1. For the two tests T4x and T5x performed at room temperature, the strains develop much faster than the field strains up to about day No. 80, i.e., about 45 days after end of filling. From then on the laboratory strains develop more slowly than in the field. However, the extrapolated curves in Fig. 30 (shown with dashed line) indicate that the laboratory strains after several years seem to end up roughly at the same values as in the field. Tests set up at constant effective stresses at room temperature to determine rate and magnitude of secondary compression will most probably show the same trend as tests T4x and T5x, thus initially the rate of secondary compression will, due to sample disturbance, be much too large, and if the tests are terminated too early, i.e., in this case earlier than about 3 months, the estimated final strain may agree less well with the field value. 2. For test T6x performed at +7.8 C (i.e., at approximately the same temperature as in the field), the strains are significantly lower than for the two other tests, although the effective stresses are somewhat higher. But also for this test the vertical strain, according to Fig. 30, is seen to increase much faster than in the field. The stress path followed for test T6x is considered to be the most correct one and should therefore also have been followed for tests T4x and T5x. The strains for these latter tests would then have been somewhat higher. The stress path followed for T6x is seen to give rather small radial strain. Testing under plane strain conditions should then give about the same vertical strains as testing under axisymetric conditions. When dismounting the specimen from the triaxial cell, the confining membrane was found to be significantly stiffer than at room temperature prior to the testing. However, the effect of this stiffening seems to be insignificant up to about day No Around day No. 30, in the laboratory tests, there is an extra increase in rate of axial strain that coincides with the assumption that the apparent pre-consolidation stress determined from the oedometer and K o -triaxial tests probably is reached at this time. 4. According to Fig. 13 the ratio H / V in the field at a depth of 5.2 m is about 0.46 at day No. 36. (For totally undrained and

17 46 Can. Geotech. J. Vol. 51, 2014 Fig. 24. Slip circle used in stability calculations with in situ vane and laboratory test results. When using laboratory tests, triaxial extension, DSS, and triaxial compression tests were used for parts a b, b c, and c d of the slip circle, respectively. All dimensions in metres. Fig. 25. Vertical strain versus depth under the centreline of the fill, according to oedometer tests, when all excess pore pressure has dissipated. In this figure v = 0 when v = v0.

18 Berre 47 Table 3. Parameter values used when estimating the degree of primary consolidation (U) in the field from oedometer tests. Depth (m) ( V ) final (kpa) final (%) k oed (m/s) c v (m 2 /s) Note: The value of final = 0 when v = v0. Fig. 26. Settlement of ground surface below the middle of the fill as measured by a field precision settlement gauge. Comparison with estimated values. Fig. 27. Principal sketch showing main assumptions made in the drained analysis. Fig. 28. Vertical strain versus depth under the centreline of the fill at day No In this figure v = 0 when v = v0. plane strain conditions it should be equal to 1.0) The laboratory tests in Fig. 29 show that at day No. 36 the ratio r / a is about 0.13, on average, for tests T4x and T5x and about 0.06 for test T6x. As a first approximation the laboratory values from axisymetric laboratory tests should be multiplied by 2 to be comparable with plane strain values, i.e., 0.26 and These values are still much lower than the field value of 0.46, so the relatively low field values are in a way supported by the laboratory tests. The very low laboratory values may be due to the fact that sample disturbance, under the prevailing stress conditions, causes more increase in a than in r. Conclusions The Onsøy test fill foundation was heavily instrumented to measure deformations and pore pressures with time for the soft

19 48 Can. Geotech. J. Vol. 51, 2014 Fig. 29. Comparison of strains measured in the field at depth 5.2 m under the centreline of the fill with strains from various triaxial tests where the effective stresses have been increased by the same magnitude and by the same rate as in the field. In this figure all strains are equal to zero when v = v0. plastic marine clay, and a major part of the recorded data are presented. The following preliminary conclusions may be drawn from the investigations made regarding stability and settlements of the fill: 1. Bjerrum (1973) proposed a correction factor to be applied to the in situ vane strength before using the strength values to compute the factor of safety against undrained failure. The factor depends on the plasticity index of the clay. Bjerrum considered this correction to be largely due to the fact that the undrained strength of soft clays increases with increasing rate of loading, and more so the higher the plasticity. However, the results of undrained triaxial and DSS strength tests with rate corrections on the Onsøy clay indicate that failure might have occurred below the test fill even with a factor of safety based on in situ vane tests of up to 1.35, i.e., a correction factor of 1/1.35 = 0.74, compared to 0.88 from Bjerrum s diagram, if absolutely no drainage had taken place in the field. This indicates that the coefficients of consolidation and the length of the drainage paths may be equally important as the rate effect and should therefore be considered when estimating the factor of safety against undrained ground failure. 2. A fairly conventional settlement analysis for the central part of the test fill with calculation of initial settlement based on undrained triaxial tests and calculation of consolidation settlement based on oedometer tests, performed at room temperature, and realistic field values of the coefficient of permeability combined with triangular strain distribution with depth gave an estimated time settlement curve that agrees fairly well with the values observed in the field. However, this appears, to some degree, to be due to compensating errors. The calculated oedometer strains at small depths are too small about 3 years after start of filling. This is counteracted by the fact that horizontal drainage and possible very high permeability in the weathered crust are both neglected in the computations, and that the strains calculated from oedometer tests at large depths are somewhat too large due to sample disturbance. 3. When the effective stress time history observed in the field is applied to an Onsøy block sample under drained conditions, the early vertical strains appear to develop much faster in the laboratory than in the field. This is believed to be due to sample disturbance. However, after long periods of time (i.e., after several years) the measured strains in the field and in the laboratory tend to approach each other. 4. Tests set-up at constant effective stresses at room temperature to determine the rate and magnitude of secondary compression will, most probably, show a much too high initial rate of