A comprehensive review of the literature concerning case histories of failures of cuts of natural slopes and embankments constructed on very soft to medium stiff clays was conducted .
The merits and limitations of the different field and laboratory testing techniques and construction procedures were outlined. The consideration of the paleogeological history as well as the hydrological and geological characteristics of the sites of cuts in natural slopes comprised of jointed competent rocks with interbeds of weak shales, clays and mudstones in establishing the design criteria of the cuts is emphasized. Also, the proper selection of the investigation and testing techniques of weak sensitive clays for the determination of their shear strength parameters used in the stability analysis of embankments constructed on these weak soils is emphasized.
Recommendations, based on experience and judgment concerning the site investigations and field and laboratory testing techniques and construction procedures are developed to help designers and practicing engineers in their task of constructing safe and economic structures.
The targeted factors of safety largely depend on the type of material involved, level of risk and uncertainty of gathered data used in the stability analysis.

D
Although advanced recent techniques used in g eotechnical testing have substantially contributed to the maturity of foundation engineering as an independent discipline, the diversity of challenges in applications manifested by unexpected failures of presumably safe engineering structures demonstrates that this field is still in a stage of further development and refinement. This is particularly true in the realm of risk assessment, forensic engineering , and design analysis of earth structures founded on soft clay foundations. Discrepancies between the in-situ and laboratory measured values of soil properties and also between the assumed and actual stratigraphic sequence of clay deposits often seriously affect the accuracy and thus the reliability of the stability analyses . This discrepancy between the theoretical analyses and the actual behavior of embankments founded on soft clays is further exacerbated
.by the negligence of time-scale effect on the behavior of h i gh l y sensitive clays .
Boundary conditions, sometimes, seriously affect the results of stability analysis particularly in the case of non-uniform subsurface groundwater flow conditions.
Records about the actual behavior of embankments constructed o v er thick sensitive clays constitute a valuable data source for subsequent analysis by designers.
*
Professor, Faculty of engineering and Technology, U. of Jordan .
1

This would, undoubtedly, contribute to the enhancement of the geotechnical engineering profession in the design, construction, and monitoring of the actual performance of embankments and their soft foundations. Of particular importance also are the documented field records about the improved performance of embankments through the use of vertical drains at different spacings to enhance pore water pressure dissipation, incorporation of stabilizing berms in the design to increase safety and rate of construction, use of pre-loading to decrease post-construction settlements, and the use of geotextiles to strengthen the foundations and decrease the potential cracking due to excessive total and differential settlements.
The assurance of an adequate factor of safety to natural and man-made earth slopes is the most challenging task of geotechnical engineers. Among the various methods used to assess the level of safety of earth slopes, the limit equilibrium method is the most widely used one. This is due to its simplicity whereby the soil mass is assumed to behave as a rigid plastic body meeting the Mohr-Coulomb failure criterion and moving along a continuous slip surface . This method usually leads to practically acceptable results in soils with perfect plastic behavior but not in brittle very stiff soils. Shear failure of soft to medium stiff cohesive soil slopes is often preceded by slow time-dependent creep movement i .
e . progressive failure.
Tiande et al (1999) in their model of progressive failure of landslides showed that local plastic failure initiates at the toe of the slope where shear stress is concentrated. and tension failure occurs at the crest of the slope. This observation is in conformity with the conclusions of Lo and Lee (1973) in their finite-element study of the slope stability of strain-softening soils. Local failure at the toe slice of a slope leads to the transmission
2

of stress to the neighboring slices and eventually to the initiation of local failure.
Propagation of local failure continues with continuous creep displacement and widening of tension cracks at the crest of the slope until it culminates in a catastrophic landslide unless local failure stops at a certain slice and strength doesn't drop to a post peak value. In this case the overall factor of safety of the slope is greater than unity.
Tiande et al (1999), in their numerical simulation of failure evolution, clearly showed that the time between the development of tension crack at the crest of slope and complete failure is much shorter than that between the initiation of local failure at the toe of the slope and the development of crack at the crest of the slope. This emphasizes the importance of observing the initiation of local failure at the toe by detecting any lateral displacements at the base of the slope at an early stage to take measures that would prevent further deterioration of slope stability.
Wright et al (1973) criticized the widely used lim it equilibrium method of slope stability for three reasons viz. the negligence of stress-strain characteristics of soil, the assumption that the factor of safety is the same for every slice, and the nonsatisfaction of all conditions of equilibrium. Contrary to the assumption that the factor of safety is the same for each slice in the Bishop's Modified Method, values vary from one location to another along the shear failure surface when using the linear elastic stress distribution. To prevent local elastic overstress al o ng the critical shear surface the required factor of safety varies from 1.44 to 1.49 for purely cohesive soils having slope angles ranging from 16° to 34° respectively and with
 eØ
=

𝐻 𝑡𝑎𝑛Ø 𝑐
= 0 (i.e. Ø = 0 o ) where

= unit weight of slope material and
H=height of slope. For soils with
 eØ
= 50 the required factor of safety ra n ges from
4

1.12 for a slop e angle of 16° to 4.36 for a slope angle of 34°. However, for a wide range of encountered conditions Wright et al (1973) recommended that a factor of safety of 1 .
5 is adequate to prevent local elastic overstress and thus the possibility of initiating progressive failure in the slope.
Since failure of soft foundations beneath stiff embankments is often of a rotational type the failure mode is represented by different types of shear tests due to the rotation of the principal stresses viz. undrained triaxial compression, direct simple shear, and undrained triaxial extension tests as shown in Figure 1 (Bjerrum 1972).
Therefore, it is of an utmost importance to select the undrained strength value that best represents the average strength along the ent ire failure surface. Proper consideration should also be given to the stratigraphic and structural features of the foundation materials like the types and thicknesses of intercalations, degree of anisotropy, and orientation of laminations. Usually the undrained strength from the undrained direct simple shear (DSS) represents the best average to be used in stability analyses. In cas e insitu vane shear test values are to be used Bjerrum (1972) recommended th e application of correction factors for the effect of rate of shearing and for the effect of anisotropy.
In the case of constructing embankments over so ft clay foundations there could be a significant difference in the stress-strain characteristics between the stiff compacted fill of the embankment and the soft clay foundation. I t is thus recommended that the undrained strength of both the embankment fill and the foundation material be reduced by applying the reduction factors RE and RF to the
5

undrained strength of the embankment and the foundation materials respectively as shown in Figure 2 (Duncan and Buchignani, 1975).
Ladd (1991) emphasized the importance of controlling the excess pore water pressure (p.w.p) in the soft clay foundation during all stages of construction on such soils. This could be done either by slowing the rate of construction or by staged construction. Staged construction involves suspension of construction at certain critical sections of embankment for periods that could range from few days to few weeks. Staged construction aims at enhancing gain in the shear strength of soft clay foundation through the gradual time-dependent dissipation of p.w.p. Staged construction often requires the installation of an efficient monitoring system that allows the measurement of excess p.w.p at different depths beneath different sections of the embankment as well as the vertical and horizontal displacement The objective of this system is to enable the designer to carry more reliable stability analyses at the end of each stage of construction before proceeding to the next stage. Assessments of the shear strength of clay deposits based on well established relationships with the vertical effective stresses should be correlated and ascertained with the insitu shear strengths measured by reliable field testing techniques. This is to ensure the validity of the assumed shear strength parameters used in the stability analyses and thus the reliability of the factor of safety computations . Some of the major pitfalls of designers in their stability analyses is their reliance on the piezometric measurement of p.w.p (to limited depths and at limited locations) alone or their reliance on the measurement of insitu shear strength by using some crude methods like the insitu vane shear tests
5

without any correction for the effects of some factors like anisotropy, rate of shearing, plasticity, and size of testing apparatus. Stabilizing berms are sometimes used to improve the stability of the embankment and accelerate the con s olidation process.
Pre-loading with fill heights exceeding the final design grade are also used to reduce the post-construction settlement. To meet construction schedules and reduce post­ construction settlements vertical sand or prefabricated wick drains are often used to provide horizontal drainage and accelerate both the consolidation of the clay foundations and their strength gain. Settlement calculations require the determination of soil stratification, soil properties, and its past geologic history within the significantly stressed zone beneath the embankment.
Uncertainties associated with the insitu measure d values of undrained shear strength of clay foundations as compared with the laboratory measured values should be dealt with by applying appropriate correction factors depending on the sensitivity of the soil, sampling disturbance, rate of shearing, type and size of shearing apparatus, degree of anisotropy, and the effect of progressive failure. Bjerrum (1972) listed 14 cases of embankments that failed although most of them showed a factor of safety well greater than 1.0 based on the insitu vane shear tests. He introduced a correction factor, µ, to the undrained strength as measured by the insitu vane shear test (VST) in the form:
C u
(corrected)
= µ.C
u
(VST) where,
= 1.7 -0.5 log P
1
(P
1
= plasticity index)
6

He also introduced correction factors for the effects of rate of shearing,
R
which depends on the
P
I of the clay and for the anisotropy,
A
which depends on the inclination of the sliding surface and the
I of the clay in the form:
C u
(field)= C u
(VST)
R
A
He indicated that the effect of anisotropy,
A
is higher in the lean clays of low plasticity than in the highly plastic clays. It could be also observed from the data listed in Table II (Bjerrurn, 1972) that the factor ;
R
is slightly higher in lean clays than in highly plastic clays. Bjerrum (1972) also emphasized the effect of progressive failure, particularly in sensitive clays with strain softening characteristics in initiating failure in embankments where the computed factor of safety, based on the VST, is well greater than 1.0. Failure starts at the highly stressed zone beneath the center of the embankment and gradually extends sideways until a full shear failure plane develops.
Overstressing the sensitive clay causes a sudden drop in its strength to the post peak value with complete destruction of its structure. When failure occurs the soil will be differently strained at the different locations along the failure surface. While the soil strength beneath the highly stressed zone is close to the residual one the soil strength beneath the less stressed zones at both ends of the failure surface will be close to the peak one . On one side the soil strength is best represented by the undrained triaxial compression test while on the other side it is best represented by the undrained triaxial extension test.
Tavenas and Leroueil (1980) recommended, in th e light of the many case histories of failure where the computed factor of safety far exceeded 1.0 (Figure 3)
7

that Bjerrum's approach in the stability analysis of embankments founded on soft clays be considered only as a crude empirical one.
In the case of non-uniform soil conditions like the presence of intercalations of silt layers, lenses of sand, or interbeds of gypsum, calcite, or aragonite within the clay deposit it has been demonstrated by many case studies that the cone penetration test (CPT) is far more superior than either the VST or the SPT in determining the variation of strength along the soil profile.
Trial embankments are sometimes constructed and instrumented and monitored to asses s and verify subsurface ground conditions and soil strength and consolidation characteristics for use in embankment stability and settlement analyses.
They are also used to check for the appearance of any indications of impending failure that could be employed during the construction of embankment like the development of longitudinal or transverse cracks. For stability analysis during staged construction of embankments over soft clays Ladd (1991) recommended the use of the undrained strength analysis (USA) rather than the effective strength analysis (ESA) because failures during staged construction often occur under undrained conditions. This requires the determination of the effective overburden stress (
 o
) and the preconsolidation stress (
 c
) along the soil profile. This requires running consolidation tests on undisturbed samples representing the whole soil profile under consideration.
Increments in effective stress during construction are then computed by proper consideration of stress distribution and pore pressure readings of the installed piezometers. Using SHANSEP (stress history and normalized soil engineering properties) approach suggested by Ladd (1974) increments in cu of soil can then be computed during construction by considering the determined normalized strength
8

parameters of the soil and its over consolidation ratio (OCR). This approach is applied on soils which prove by testing to have a normalized soil engineering behavior.
The C u for a normally consolidated (NC) clay could be estimated from the equation suggested by Skempton (1957): where
 ’ o
= effective overburden stress
Jamiolkowski et al (1985) suggested the following equation for lightly over consolidated (OC) clays: c u
/  ’ c
= 0.23
0.04 (  ’ c
=effective pre- consolidation stress)
Mesri (1989) suggested: c u
/  ’ c
=
0.22
Ladd et al (1977) suggested the following relationship between the strengths of OC clays and NC clays:
Mayne and Mitchell (1988) suggested the following equation for estimating
 ’ c for a natural clay deposit
They also suggested that OCR can be estimated from the C u (fie l d) of the natural clay deposit in the form:
9

OCR = B
𝐶 𝑢
(𝑓𝑖𝑒𝑙𝑑) 𝜎′ 𝑜 where B=22 (P
I
) -0.48
Tavenas and Lerouiel (1980) suggested a procedure for using the
Cu
(VST) in stability analysis. This method involves the reduction of the measured undrained shear strength values in the zone of the weathered top clay crust and the use of the measured values beneath this zone (Fig. 4) The computed factor of safety, assuming full mobilization of strength in the embankment (La Rochelle et al. 1974), is then divided by the factor F f which is a function of the liquid limit and sensitivity of soil to get the corrected factor of safety (Figure 5) . The top curve in Figure 5 is for highly sensitive clays and the bottom one for clays of low sensitivity
It is worth noting that soil disturbance during testing in the laboratory often leads to the underestimation of the maximum pre-consolidation stress, soil compressibility, and soil coefficient of permeability (Rixner, 2001). However, all the above empirical equations are approximate and the strength-stress relationships should, for important or sensitive structures, be verified by actual testing. This is particularly true in the case of construction on thick highly compressible and sensitive clays with high liquidity index, whereby careful monitoring and interpretation of instrumentation d a t a by highly qualified and experienced geotechnical engineers is considered crucial to the safety and successful completion of construction works .
1
0

The philosophy of staged construction and proper monitoring and interpretation of geotechnical data could also be applied to projects that involve cuts in natural slopes with possible daylighting of sensitive plastic clays. The following case studies are presented to demonstrate the importance of proper selection and interpretation of site investigations and soil testing and the scheduling of works accordingly.
LANDSLIDES AT WADI ES-SIR SEWAGE
TREATMENT PLANT
The Wadi Es-Sir Sewage Treatment Plant (WESTP) is located about 10 kms to the south west of Wadi Es-Sir town and about 20kms to the east of the Jordan-Dead
Sea Rift (Fig. 6). The site is characterized by its relatively steep topography with an average inclination of about 15%. It is also characterized by its dry hot summer and its moderately cold winter with an average annual rainf all ranging between 250mm and350mm.
During th e construction works for the lagoons th e site was affected by four landslides at different dates, namely 8 September, 1993; 28 November, 1993; IO July
1995; and 23 February 1997 (Fig. 7).
The investigation works that were carried out in the site at different stages indicated that the geologic cross-section generally consists of (from top to bottom): i.
ii.
Loose heterogeneous man-made fill.
Collu v ium consisting of sandy silty clay intermixed with variable percentages of graved to boulder- size fragments of limestone and marlstone.
11

iii.
Bedrock consisting of poor quality and disturbed intercalations of reworked clayey marl, marlstone, and limestone affected by tectonic movements and old landslides . It belongs to the Naur Formation (Al-2) of the lower Ajloun Group which is known in Jordan as the formation most susceptible to landslides due to the presence of the wet plastic weak clayey marl.
iv.
Yellowish brown to yellowish green clayey marl sometimes intermixed with variable percentages of gravels and cobbles of limestone and marlstone .
Figure 8 shows features of the July 10, 1995 landslide which took place during excavation inspite of the flattening of the cut slopes to IV:4H after the occurrence of the November 28, 1993 landslide. The clayey marl forming the sliding surface has a LL of 77.I and a PI of 48.1
. The geomorphological features adjacent to the slide attest to the fact that the region is plagued with multiple old landslides forming slip surfaces where the shear strength is close to the residual one .
Figure 9 shows the effects of the February 23, 1997 landslide on its adjacent
July 10, 1995 slide area after it was further flattened to IV:5H and provided with gabion walls and surface drainage ditches . The slide caused extreme disturbance to the area with destruction and dislocation of the gabion walls and drainage ditches.
The post-failure investigations of the above landslides indicate that the main causes of slides were:
(a) The presence of pre-existing slip surfaces manifested by the geomorphological features of the region which indicate that the region had
12

experienced intensive tectonic disturbance and many old landslides in its past geologic history .
(b) The presence of highly plastic beds of clayey mar l underlying the jointed and weathered strata of limestone and marlstone topped with loose colluvium and man-made fill. The marly beds are dipping unfavorably out of the slope at angles between 8 and 12 degrees which are considered critical to the stability of these slopes dominated by highly plastic marls. The strength along these dipping beds is close to the residual one with their residual shear strength parameters experimentally estimated at a cohesion of 3 to 8 kN/m
2 and angle of friction of 6 to 9 degrees.
(c) The high rate of water infiltration during intense rainstorms through the highly penneable beds of loose colluviurn and jointed rocks leading to sudden rise in pore pressures and softening of underlying clayey marl beds .
(d) The steepness of the ground and the low factors of safety adopted for the slope cuts leading to creep and progressive failure of the marls that are susceptible to strain-softening.
(e) The poor drainage conditions allowing the saturation of the fill, colluvium and underlying poor bedrock materials in the absence of the vegetative cover causing a substantial increase in the driving forces and a decrease in the resisting forces.
DIKE 19-ARAB POTASH PROJECT
DEAD SEA-JORDAN
Dike 19 is an 8.3 km long embankment and form s with dike 20 one dike with a total length of 1 1.6 km enclosing a salt pan with a storage capacity of 71.3 Mm
3
13

(Fig. 10). It has a crest width of 8 m and varies in height from 8 to 14 m . The upstream and downstream embankments slope at 2.5 H:IV with berms whose width increases with the increase in the dike's height according to certain formula specified by the designer. The compacted reworked lisan marl constitutes the major portion of the dike's body (Fig. 11). Construction of the dike commenced in March, 1998 and ended in November, 1999. On September 22, 1999, Variation Order No I wa s issued including the reduction of the height of the dike by 2m and increasing the thickness of the berms by about lm to satisfy requirements dictated by the stability analysis of the dike and the decision of impounding the salt pan in January, 2000. On March 22 ,
2000 a sudden partial failure of the dike occurred causing a rapid release of abou t 56
Mm
3 in about 30 minutes. Investigations indicated that failure started near Chainage
6+000 and caused a 2.3 km wide gap in the body of the dike between Ch 4+600 and
Ch 6+900. The design criteria for dike 19 relied heavily on the experience gained from the construction of a trial dike and dike 18 which were constructed near the site of dike 19 . How ever, it was soon discovered at the early stages of construction that the foundations of dike 19 were more compressible, more sensitive, and less permeable than those of dike 18. The foundation materials genera ll y consist of very soft to medium stiff thick to very thick bluisk grey thinly laminated silty clay with stronger inerbeds of gypsum and aragonite and occasionally with organic debris.
Staged-construction design using the observational approach wa s adopted in the construction of the dike to ensure the safety of the structure through the control of the rate of construction and modification of design features o f the dike. The instrumentation system installed at 1 km intervals along the dike comprised pneumat ic piezometers to measure pore pressures in the foundation soils, standpipe piezomete r s to measure long-term water levels in the built dike and its foundation after
14

impoundment, survey monuments to measure displacements, and horizontal magnetic extensometers to measure the vertical settlements of the foundation materials.
Control of construction rate
The main criterion that was used for controlling the rate of construction in dike
19 was the ratio of the excess pore water pressure, u. to the corresponding increment in the total vertical stresse,

, namely, B-bar, i.e. =
∆ 𝑢
∆𝜎 𝑣
. It was originally specified that th e maximum values of
𝐵̅ are 0.7 beneath the central portion of the dike and 0.5 elsewere. However, it was noticed that due to the very low penneability of the foundation s oils the dissipation of pore pressures was very slow. Therefore, in order to avoid the anticipated delay in the completion of the works and the consequent delay in impounding the pan the maximum values of the
𝐵̅ were gradually relaxed
(in increments) to 0.95 below the central portion of the dike and to 0.7 elsewhere. The designer considered the end of construction state as the most critical. To meet the contemplated date of completion of construction the required factors of safety at end of construction was also relaxed from 1.3 to 1.25. With the reduction of the height of the dike by 2m (Variation Order No 1) it became possible to complete the construction of the dike on the contemplated date as was originally planned.
Construction of the dike proceeded by placing the fill material composed of reworked marl in 0.15 m thick lifts compacted to a minimum degree of compaction of 95 percent of Standard Proctor at an optimum moisture content ranging mostly between
18 and 20 percent. However, due to the high
𝐵̅ value the placement of fill was on many occasions suspended for periods ranging from 3 to 7 days. During construction and impoundment many longitudinal and transverse cracks developed in the body of
15

the dike . The development of these cracks was most probably due to the excessive total and differential settlements along and across the body of the dike associated with substantial horizontal displacements at the base of the dike .
The settlement reached about 4m beneath the central partion of the dike before failure near Ch 6+000 where failure was most probably initiated as deduced from the post-failure investigations .
s
Many stability analyses were frequently carried out at different stages of construction using data from the insitu vane shear tests (VST) carried down to a depth of 1Om below the base of the dike particulary at the locations where cracks developed in the body of the dike. The results of these tests were employed, without correction, as the basis for stability analyses. The vane apparatus that was used measured
50mmX100mm and 63.5 mm X 127mm vs. the 75mm X 150mm which was used during the pre-construction investigations.
It was reported (Gibb 1995) that the smaller vane 50mm X 100 mm gave undrained strength (c u
) values 80 percent higher than those obtained using the 75mm
X 150mm vane. The reliance of the stability analysis on the results of the VST resulted in an overestimation of the factors of safety. Also, the VST showed inconsistent results regarding both the increase in strength with depth or with time due
consolidation. It seems that the intermittent presence of salt layers in between the laminated silty clay layers resulted in the inconsistent strength measurements. The
VST results generally suffer from the following limitations which render them unacceptable as a reliable source of data for stability analyses unless properly corrected on the basis of correlation with other tests like cone penetration te s t (CPT),
16

direct simple shear test (DSS), undrained triaxial compression (TC) or extension (TE) tests: i) ii) iii) The presence of gravels or salt crystals or fibrous inclusions increases the measured
Cu values.
iv)
The test results should be corrected for the effect of the size of the v ane apparatus.
The test tends to overestimate the
Cu value in anisotropic laminated plastic clays where the C u along the laminae is smaller than across them.
Although a correction factor has been proposed by Bjerrum (1972) to be applied to the test results, the scatter of the data as noticed by Ladd and
Foott (1974) demonstrated the uncertainty of the results which had little correlation with the values obtained from bac k stability analyses of failed embankments placed on soft clays.
v) The test tends to overtimate C u value due to the higher strain rate during the test than that in the actual case.
The
Cu value assigned to the compacted fill of the dike body in the undrained stability analyses was 100 kPa instead of the measured value of 190kPa to account for the potential cracking of the dike's body. Cracking was expected due to the high strain incompatibility between the stiff embankment fill (peak strength at about 1.5 to 2 % strain) and the soft clay foundation (peak strength at about 5% to 12% strain).
The unit weight of the compacted embankment fill was wrongly assurmd to be 15.7 kN/m
3 instead of about 18.6 kN/m
3 which resulted in an overestimation of the factor of safety.
All the above factors resulted in an actual factor of safety at the end of construction considerably less than the presumed one (1.25).
17

Impounding the salt pan commenced on January 4,2000 i.e. a short rest period was allowed between the end of construction and the commencement of impoundment This short period was not adequate to cause any considerable increase in the factor of safety. This resulted in the overstressing of some portions of the soft foundations along the potential failure surface (Wright 1973). The early impoundment of the salt pan had exacerbated the critical stability of the dike through the development of significant zones of contained plastic flow leading to progressive failure (Ladd 1991).
The rate of settlement during impoundment didn't show any noticeable decrease as compared with that during construction and was as well combined with a high rise in pore pressure particularly near Ch 6+000. This caused a continuous decrease in the factor of safety during impoundment until it culminated in a catastrophic shear failure on March 22,2000 when the F.S dropped to 1.0.
The degree of consolidation for the 10 m to 15 m zone of the foundation material beneath the dike was less than 20% to 30% and was smaller at the deeper zones where the soil gained little or no strength during construction. No use was made either of the cone penetration test, as was the case in the pre-construction investigation, or of the pore pressure readings in estimating the strength gain of the foundation material during construction. Reliance was solely based on the uncorrected insitu VST readings down to a depth of only !Om below the base of the dike to estimate gain in the underained strength of the foundation materials during construction. The se tests that were carried out to a shallow depth missed the deeper soft layers of the laminated silty clay through and along which the shear failure surface most probably have passed. No stability analyses were carried out during the impoundment stage inspite of the high pore pressure and settlement readings. These
18

high readings of pore pressure and settlement, particularly near Ch 6+000, are due to the fact that the foundation materials consist of more than 60m thick soft to v. soft sensitive, highly compressible, and relatively impervious laminated silty clays. The laminations are sometimes disturbed and convoluted, and th us impeding the drainage and the fast dissipation of pore pressures.
Post-construction investigations
The post construction investigation didn't disclos e any geological features or imperfections that could have caused the failure of the dike. Failure couldn't also be attributed to piping erosion in the foundation materials due to their cohesiveness and low permability and the low hydraulic gradient, or due to piping through the body of the dike which was well compacted and provided with affective drainage control measures. The partial collapse of the dike was found to be due to the inadequate bearing capacity of the foundation material which gained little strength during construction and due to the destabilizing effect of the impounded water behind the dike and within the upstream longitudinal cracks.
The three boreholes which were drilled in the failed section after failure have defined the location of the failure surface as being that which separates the disturbed zone above it from the undisturbed zone beneath it (Dar/Harza 2001). The failure surface near Ch 6+000 is most probably a rotational one that starts at the junction point of the upstream berm with the dike and exits at about 27m from the downstream toe of the dike with a maximum depth of about 20m beneath the dike (Fig. 12).
Longer rest period between the end of construction and commencement of impoundment and control of the rate of impoundment, based on pore pressure
19

readings and stability analyses, could have saved the dike as shown in the illustrative sketch (Fig. 13).
CONCLUSIONS AND RECOMMENDATIONS
1The factor of safety that should be adopted for cut slopes and embankments underlain by soft clay layers should be commensurate with the degree of uniformity of ground conditions, anticipated changes in the environmental and stress conditions, and the severity of the adverse economic, social, and environmental consequences of any potential failure.
A factor of safety ranging between 1.4 and 1.6, depending on the sensitivity of soils, is generally adequate to avoid overstressing and th us progressive failure of slopes.
2For construction of embankments on soft clays it is recommended to adopt staged construction based on the observational approach by monitoring the pore pressures and the vertical and horizontal displacements during and for a reasonable period after construction . The undrained stability analysis is recommended for construction on soft clays with low permeability.
SHANSEP approach recommended by Ladd could be used to establish the soil strength profile if the soil proved, by testing, to have a normalized behavior. The CPT supported by DSS tests on undisturbed samples is far better than the VST in defining soil stratigraphy and evaluating strength gain during construction.
20

3Rapid relief of stresses by deep excavation in slopes dominated or underlain by overconsolidated plastic clays could lead to shear failures under undrained conditions. Staged excavation with proper monitoring of vertical and horizontal displacements would allow good evaluation of the stability and control of construction activities .
4Proper consideration of the past geologic history supported by careful examination of the geomorphological features of slopes would allow early detection of the presence of pre-existing slip planes that could form potential failure surfaces due to the low shear strength on such planes.
5In the design of cut slopes underlain by layers of plastic clayey marls inclined unfavorably towards the excavation utmost care should be excercised not to daylight such. strata or even be close to them. Adequate confinement is needed in order to avoid overstressing of such soils that are often susceptible to strain-softening.
6In the design of stiff embankments over soft clay foundations it is strongly recommended to introduce correction factors to the undrained strength of both the embankment fill and the clay foundation as suggested by Duncan and Buchignani (1975) to account for the high stress-strain incompatibility between the stiff embankment fill and the soft clay foundation and thus to avoid the initiation of progressive failure.
7Undrained direct simple shear test better represents the average undrained strength of the clay foundations underlying stiff embankments than either the undrained triaxial compression or extension tests. Correction to the results of this test, however, should be introduced in case evidence exists that the layers
21

of clay foundations experienced, in their pas t geologic history, strong disturbance either by liquefaction or slippage .
The author expresses his deep appreciation for the University of Jordan in general and for the Deanship of Scientific Research in particular for their moral and financial support which enabled him finish this research on time during his sabbatical leave from the faculty of Engineering and Technology in the year 2006-2007.
22

1. Bjerrum L. (1972), "Embankments on Soft Ground, State of the Art Report,"
Proceedings ASCE Specialty Conference on Performance of Earth and Earth
Supported Structures, Lafayette, Vol. 2, pp. 10-54.
2.
Dar-Harza, (2001), "Dike 19 of APC-Stage 111-Expansion Works," Engineering
Assessment Report, Jordan.
3.
Duncan, J.M
. and Buchignani, A.L. (1975), An Engineering Manual for Slope
Stability Studies, Department of Civil Engineering, U. of California, Berkely,
USA.
4.
Gibb, Sir Alexander & Partners Ltd. (1995), Arab Potash Project, Stage II
Expansion Works, Onshore Geotechnical Interpretative Report, April 1995 ,
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