Stability of Natural Slopes and Embankments Underlain

By Weak Clays

Prof. Dr. Yousef M. Masannat*

Abstract

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.

1

Stability of Natural Slopes and Embankments

Underlain by Weak Plastic Clays

By

Yousef M. Masannat*

INTRODUCTION

Although advanced recent techniques used in geotechnical 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 highly 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 over thick

sensitive clays constitute a valuable data source for subsequent analysis by designers.

* Professor, Faculty of engineering and Technology, U. of Jordan.

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.

STABILITY OF SLOPES

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 limit 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 along 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 tan Ø

c = 0 (i.e. Ø = 0o) where = unit weight of slope material and

H=height of slope. For soils with eØ = 50 the required factor of safety ranges from

5

1.12 for a slope 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 entire 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 case insitu vane shear test values

are to be used Bjerrum (1972) recommended the application of correction factors

for the effect of rate of shearing and for the effect of anisotropy.

In the case of constructing embankments over soft 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).

STAGED CONSTRUCTION

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

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 consolidation 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.

TESTING UNCERTAINTIES

Uncertainties associated with the insitu measured 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:

Cu (corrected) = µ.Cu (VST)

where, µ = 1.7 -0.5 log P1 (P1 = plasticity index)

6

7

He also introduced correction factors for the effects of rate of shearing, µR, which

depends on the PI of the clay and for the anisotropy, µA, which depends on the

inclination of the sliding surface and the PI of the clay in the form:

Cu(field)= Cu(VST) µR. µA

He indicated that the effect of anisotropy, µA , 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; µR, 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 the light of the many case

histories of failure where the computed factor of safety far exceeded 1.0 (Figure 3)

8

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 assess 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

9

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 Cu 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:

cu / ’c = 0.23 ± 0.04 (’c =effective pre- consolidation stress)

Mesri (1989) suggested: cu / ’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 Cu (field) of the natural clay

deposit in the form:

10

OCR = B Cu( field )

σ ' o

where B=22 (PI)-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 Ff 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 .

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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 rainfall ranging between 250mm

and350mm.

During the construction works for the lagoons the 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. Loose heterogeneous man-made fill.

ii. Colluvium consisting of sandy silty clay intermixed with variable

percentages of graved -to boulder- size fragments of limestone and

marlstone.

12

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

13

experienced intensive tectonic disturbance and many old landslides in its past

geologic history.

(b) The presence of highly plastic beds of clayey marl 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/m2 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 forms with dike 20 one dike with

a total length of 11.6 km enclosing a salt pan with a storage capacity of 71.3 Mm3

14

(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 was 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 about 56

Mm3 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. However, 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 generally 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 was 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 of the dike. The

instrumentation system installed at 1 km intervals along the dike comprised pneumatic

piezometers to measure pore pressures in the foundation soils, standpipe piezometers

to measure long-term water levels in the built dike and its foundation after

15

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. B = ∆u

∆ σ v. It was originally

specified that the maximum values of B 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 soils 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 B 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 B 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

16

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 .

Stability analyses

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 (cu) 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

to 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 test (CPT),

17

direct simple shear test (DSS), undrained triaxial compression (TC) or extension (TE)

tests:

i) The test results should be corrected for the effect of the size of the vane

apparatus.

ii) The test tends to overestimate the Cu value in anisotropic laminated plastic

clays where the Cu along the laminae is smaller than across them.

iii) The presence of gravels or salt crystals or fibrous inclusions increases the

measured Cu values.

iv) 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 back stability analyses of failed

embankments placed on soft clays.

v) The test tends to overtimate Cu 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/m3 instead of about 18.6 kN/m3 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).

18

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. These 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

19

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 thus impeding the drainage

and the fast dissipation of pore pressures.

Post-construction investigations

The post construction investigation didn't disclose 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

20

readings and stability analyses, could have saved the dike as shown in the illustrative

sketch (Fig. 13).

CONCLUSIONS AND RECOMMENDATIONS

1- The 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 thus progressive

failure of slopes.

2- For 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.

21

3- Rapid 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.

4- Proper 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.

5- In 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.

6- In 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.

7- Undrained 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

22

of clay foundations experienced, in their past geologic history, strong

disturbance either by liquefaction or slippage.

AKNOWLEDGMENTS

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.

23

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