Mechanical properties and behaviour of a partially saturated lime-treated, high plasticity clay

    Mechanical properties and behaviour of a partially saturated lime-treated, high plasticity clay Xiwei Zhang, Maria Mavroulidou, M.J. Gunn PII: DOI: Reference:

S0013-7952(15)00161-1 doi: 10.1016/j.enggeo.2015.05.007 ENGEO 4044

To appear in:

Engineering Geology

Received date: Revised date: Accepted date:

3 March 2014 27 April 2015 9 May 2015

Please cite this article as: Zhang, Xiwei, Mavroulidou, Maria, Gunn, M.J., Mechanical properties and behaviour of a partially saturated lime-treated, high plasticity clay, Engineering Geology (2015), doi: 10.1016/j.enggeo.2015.05.007

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Mechanical properties and behaviour of a partially saturated lime-

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treated, high plasticity clay

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Dr Xiwei Zhanga,(BEng, MSc, PhD);

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Dr Maria Mavroulidoub*(Dipl-Ing, D.E.A, PhD, MTEE-Greece);

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Prof M.J. Gunnc ;

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a. Address during the presented research: London South Bank University, 103 Borough Road, London, SE1 0AA, UK Present address: Associated research fellow, Key Laboratory of

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Ministry of Education on Safe Mining of Deep Metal Mines, Northeastern University,

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Shenyang, Liaoning, 110819, China;

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b. Reader of Geotechnical Engineering, London South Bank University, 103 Borough Road, London, SE1 0AA, UK

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c. Emeritus Professor of Geotechnical Engineering, London South Bank University, 103 Borough Road, London, SE1 0AA, UK Manuscript submitted to the journal Engineering Geology

Date written: February 2014 Date of first revision: December 2014; Second revision: April 2015 Word count:

10,968 (including abstract, keywords, notation, references, table and

figure captions) Number of Figures: 17 figure captions Number of Tables: 8

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*Corresponding author

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Faculty of Engineering, Science & the Built Environment

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Dr Maria Mavroulidou

London South Bank University 103 Borough Road

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London SE1 0AA

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Tel (work): 02078157646; (please use preferably email)

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Email: [email protected]

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Mechanical properties and behaviour of a partially saturated lime-

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treated, high plasticity clay

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ABSTRACT: This paper investigates the effect of suction and lime treatment on the volumetric and shear behaviour of a partially saturated high plasticity clay (London Clay). A series of triaxial tests were performed on statically compacted London Clay and lime-treated London Clay specimens. These

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were prepared at the same target void ratio and tested under partially saturated conditions. The tests

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concerned isotropic compression and two different types of shearing tests (a) shearing at constant suction and (b) shearing at constant water content, using the axis translation technique. Based on the

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results, the effects of suction and lime-induced bonding are evaluated and quantified within an

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elastoplastic partially saturated soil framework.

Keywords: lime-treated clay; partially saturated soil; suction controlled triaxial testing

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Notation

A1 and A2; curve fitting parameters representing intercepts at zero suction B1 and B2 curve fitting parameters (slopes) a1 , a2 , b1 and b2 curve fitting parameters c′ cohesion (true or apparent) cv coefficient of consolidation e void ratio f(c) lumped cementation bonding shear strength component Gs specific gravity 3

ACCEPTED MANUSCRIPT h specimen height M critical state stress ratio

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Mα critical state stress ratio with respect to the mean net stress (partially saturated soil)

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Mb critical state stress ratio with respect to the matric suction (partially saturated soil) N(s) the intercept of the normal compression line

Ny intercept of the line of isotropic compression yield points

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na and nb components of shear strength associated respectively with (p-ua) and (ua –uw)

(p-uα) mean net stress, here denoted as

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p mean stress

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nc suction-dependent cementation bonding shear strength component

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pat reference pressure (atmospheric pressure)

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p´c yield effective stress (apparent preconsolidation pressure) q and qp deviator stress and peak deviator stress respectively

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qc cementation bonding (‘true’ cohesion) at saturated state RL the loading rate of the ramped consolidation s suction (kPa)

sd yield suction factor during wetting Sr degree of saturation ua air pressure uex excess pore pressure uw pore water pressure (ua-uw) matric suction 4

ACCEPTED MANUSCRIPT v soil specific volume vw specific volume of water

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vκs specific volume corresponding to the yield suction

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w water content (gravimetric) εv volumetric strain θ volumetric water content

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λ(s), the slope of the normal compression line.

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κs elastic stiffness parameter with respect to a change in matric suction

λs plastic stiffness parameter with respect to a change in matric suction

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λy slope of the line of isotropic compression yield points

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ρd dry density

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(σ-ua)f is the net stress normal to the shear plane at failure τf shear strength of the soil

a

and

internal angle of friction, peak angle of friction, and critical state angle of friction

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  ,  p ,  c

b

friction angle terms associated respectively with net stress and matric suction effects

1 Introduction

Chemical soil improvement using additives such as lime is a technique that has been used extensively in construction, most commonly for pavement applications. With this application in mind, most international literature on lime treated soils focuses on simple tests such as California Bearing Ratio (CBR) or unconfined compressive 5

ACCEPTED MANUSCRIPT strength (UCS) tests. With an increasing use of the technique in a wider range of applications, the need has emerged for more sophisticated testing and thorough

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experimental evidence of the soil parameters involved in constitutive models able to

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describe the behaviour of lime-treated soils under saturated and partially saturated conditions. The latter aspect is particularly relevant, as lime-treated soils are typically compacted after treatment and hence, by definition, partially saturated. Therefore

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prior to constitutive modelling, quality suction-control triaxial testing data is required.

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To the Authors’ knowledge such data has been lacking in the literature.

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The aim of this paper is to present results from an extensive experimental programme

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funded by the UK Engineering and Physical Sciences Research Council (EPSRC) on

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the hydro-mechanical properties of a high plasticity lime-treated clay (London Clay), based on suction-controlled triaxial testing, using the axis translation technique. The

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results are interpreted in terms of elasto-plastic modelling parameters within an unsaturated soil mechanics framework.

2. Background 2.1 Lime treatment of clayey soils The treatment of clayey soils with lime is a widely used technique of ground improvement. The improvement consists in a more or less instantaneous reduction in plasticity (Sherwood, 1993) as well as shrinkage and swelling characteristics (provided that sulphates are not present) (e.g. Bell, 1996; Al-Rawas et al, 2005). 6

ACCEPTED MANUSCRIPT These immediate/short term effects on the soil upon lime treatment, commonly mentioned as ‘lime modification’ (NLA, 2004), are attributed to rapid ion exchange

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reactions between exchangeable clay ions and calcium ions provided by the lime.

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Cation exchange would be the first step towards more permanent changes; following a modification of the electrolyte content in the water due to the increased exchangeable calcium ion concentration, flocculation and agglomeration of the soil particles occurs

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and transforms the plastic soil to a granular and less plastic material (Bell, 1996). It

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was also argued that this flocculation could also be due to the early formation of small quantities of calcium silicate or calcium aluminate hydrates; these could create some

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bridging between particles and consequently flocculation (Diamond & Kinter, 1965).

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This could justify the increase in CBR and UCS shortly after lime addition, frequently

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reported in the literature (see e.g. Sherwood, 1993; Bell, 1996).

If enough lime beyond the Initial Consumption of Lime (ICL) content is present (i.e.

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the lime percentage which increased the pH of the soil to 12.35-12.45), UCS and CBR were shown to continue to increase potentially over long time periods, through slower pozzolanic reactions between lime, silica and alumina, which produce cementing agents (Brandtl, 1981; Bell, 1996; Sherwood, 1993). These reactions are usually reported as ‘stabilisation’ reactions (NLA, 2004, Sherwood, 1993). They are caused by the highly alkaline environment upon lime addition, which promotes dissolution of siliceous and aluminous compounds from the clay mineral lattice. The compounds dissolved from the clay mineral lattice react with calcium ions in pore water to form calcium silicate hydrates, calcium aluminate hydrates and hydrated calcium alumino7

ACCEPTED MANUSCRIPT silicates, which coat the soil particles and subsequently crystallise to bond them (Bell, 1996). However strength does not increase linearly with lime content; it may in fact

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decrease upon excessive lime addition as cementation is limited by the available

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amount of silica; when all silica in the clay is used up, further lime addition would not result in the formation of any new cementation products; thus any further lime may reduce strength as lime has no good frictional properties (Bell 1988; Brandtl 1981;

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Hausmann 1990).

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An additional reaction between soil and lime is carbonation in the presence of carbon dioxide. Namely, when the CO2 is dissolved in the soil pore water it reacts with the

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hydroxyle ions, forming carbonate ions, which subsequently react with the calcium ions. This results in the formation of caclium carbonate CaCO3, a weak cement whose

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formation is normally undesirable, as the reaction consumes lime which would have otherwise been used in pozzolanic reactions for the formation of stronger cementitious

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bonds. Another undesirable effect of carbonation is the fact that it delays penetration of ions on the surface of the clay and increases the time for these to reach the reaction sites (Barker, 2002).

2.2 The mechanical behaviour of partially saturated soils The mechanical behaviour of partially saturated uncemented soils in terms of volume change and shear strength characteristics was extensively investigated in the last four decades (e.g. Fredlund & Morgenstern, 1976; Ho & Fredlund, 1982; Alonso et al., 8

ACCEPTED MANUSCRIPT 1990; Toll, 1990; Wheeler & Sivakumar, 1995; Cui & Delage, 1996; Sharma, 1998; Romero, 1999; Rampino et al., 2000; Cabarkapa, 2001, Jotisankasa, 2005). It is

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common knowledge that suction increases the shear strength of the soil. The efforts of

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experimental work have however focused on the mathematical description of the shear strength of a partially saturated soil as a function of the selected stress state variables. An early expression for the shear strength behaviour of partially saturated

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soils was proposed by Bishop (1959), using an effective stress approach. Later

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Fredlund et al (1978) suggested an expression of the shear strength for partially saturated soils in terms of an extended Mohr-Coulomb failure criterion, using two

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independent stress state variables (net stress and matric suction), written as:

  c'  u  tan    u  u  tan  a

f

a

w

b

(Eqn. 1)

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f

where  f refers to the shear strength of the soil, c ' is the cohesion (true or apparent), the net stress normal to the shear plane at failure,   is the angle of

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  ua  f is

internal friction related to the applied net normal stress and considered equal to that of saturated soils, and b a friction angle term associated with matric suction ua  uw  effects. Based on data fitting from other researchers’ work, within suction ranges between 0 and 200 kPa, early work by Fredlund et al. (1978) suggested that the value of b would be a constant. Later, new experimental evidence from tests including higher suctions, demonstrated that the increase of shear strength due to suction becomes nonlinear when the range of suction is extended to large values (Escario & Saez, 1986; Fredlund et al., 1987; Cabarkapa, 2001). Modifications of the above 9

ACCEPTED MANUSCRIPT criterion were therefore suggested to introduce an appropriate expression for the non-

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linear variation of the parameter b (Fredlund et al, 1996, Vanapalli et al, 1996).

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Concerning the main aspects of the volumetric behaviour of a partially saturated uncemented soil upon isotropic/one dimensional loading and unloading under a certain constant suction it was shown that the volume change response of a partially

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saturated soil depends on the initial and final stress state with respect to the mean net

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stress and matric suction, as well as on the particular path followed from the initial to final state (Alonso et al.1990). The yield stress increases with increasing suction; the

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slope of the loading and unloading-reloading curve of a partially saturated soil is not

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the same as that of the same soil in a saturated state (i.e. suction tends to stiffen the

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partially saturated soils against the loading). This behaviour is similar to that manifested in natural fully saturated cemented soils, as a result of cementation. As for

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saturated soils, upon isotropic or 1-D unloading, some rebound/swelling of the partially saturated soil occurs. It was reported that the gradient of the partially saturated soil swelling lines is almost independent of suction (e.g. Sivakumar, 1993).

On the other hand, concerning the effect of the suction under a certain constant mean net stress (p-uα), it was shown that a change in suction may induce irrecoverable volumetric strain for most clay soils. A soil subjected to a wetting or drying path (i.e. decreasing or increasing suction respectively), may swell or shrink, under the combined effect of suction and mean net stress (i.e. it may either swell at low mean 10

ACCEPTED MANUSCRIPT net stress values or collapse upon wetting at high mean net stress values, and contract upon the drying path, as described e.g. in Matyas & Radhakrishna, 1968 or Alonso et

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al. 1990). Upon drying, shrinkage is generally observed for most soil specimens that

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are initially saturated, often with an apparent yield point beyond which an irreversible volumetric contraction can be observed. Experimental evidence shows that this threshold value (yield suction) depends both on the suction history and the initial void

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ratio (Chen et al, 1999). For compacted expansive clays the volume change due to

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variations in suction upon drying/wetting was found to be much larger than for nonexpansive clays. Substantial irreversible shrinkage or swelling strains may occur.

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Whether irreversible swelling or shrinkage will occur depends on the history of

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compaction pressure and subsequent stress paths (as a function of suction or net

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stress) and not solely on the soil type (e.g. Cui et al, 2002).

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Whereas the above findings have now been established for compacted soils that have not been subjected to chemical treatment, there is paucity in experimental evidence concerning the behaviour of partially saturated chemically stabilised soils. For such soils, both stiffness and shear strength behaviour are likely to be affected by the relative magnitude and interaction between chemically-induced bonding and suction. It is therefore of interest to investigate how the trends followed by these soils relate to the findings concerning untreated partially saturated soils.

3 Materials and methods 11

ACCEPTED MANUSCRIPT 3.1 Materials The soil used in this study was London Clay, a high plasticity stiff overconsolidated

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marine clay (in its natural state) extensively encountered in construction in the

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London area and the South Eastern England (a very densely populated area with intensive industrial activity) including pavement construction, airports (e.g. Heathrow Terminal 5), underground railway (an example of recent engineering works being the

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Crossrail project), embankment and building foundation construction.

The London Clay samples used in this study came from an excavation at Westminster

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Bridge in London and depths between 30-31 m below ground level. The soil was air-

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dried at an average temperature of 220 C and a relative humidity of 60% for a month

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and was subsequently pulverised. Figure 1 shows the particle size distribution of the portion of two samples of the pulverised soil passing the BS 425 μm sieve (BSI

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1990a), (represented by the two different curves in Fig. 1). The composition of the soil is shown in Table 1. Note the presence of smectite, which causes the London Clay soil to manifest a swelling/shrinking soil behaviour.

Fig. 1 Particle size distribution of the London Clay soil

Table 1 Composition of London Clay soil used in this study

Commercially available hydrated lime was used in this research. Its suitability for soil 12

ACCEPTED MANUSCRIPT stabilisation was confirmed (BSI 1990b). The relative proportion of calcium hydroxide to calcium oxide was found to be 4.88:1.00 based on chemical analysis on

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the lime sample carried out in duplicate. Plasticity tests performed on London Clay

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mixed in dry condition with lime at percentages of 0%-8% lime by dry mass of soil showed no change in the plasticity characteristics of the lime-treated soil beyond approximately 4% of lime addition. Hence 4% was considered to be the minimum

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necessary lime percentage for treating this clay. The percentage was confirmed by

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initial consumption of lime (ICL) test results (BSI 1990b) which showed that the ICL of this soil was between 3 and 4% (see Mavroulidou et al, 2013a and b). Therefore an

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amount of 4% lime by dry soil mass was used to treat this soil. The main physical

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characteristics of the two soils (treated vs untreated) are shown in Table 2.

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Table 2 Physical characteristics of untreated and 4% lime treated London Clay soils

3.2 Specimen preparation Specimen dimensions were 76 mm height and 38 mm diameter for all suction controlled triaxial tests, and 100 mm height and 50 mm diameter for the two suction unloading (i.e. suction decrease over a wetting path) tests. Due to available equipment limitations some of the independent isotropic compression tests, as well as some of the saturated soil triaxial tests performed at zero suction, (presented in Mavroulidou et al, 2011 and included here for the sake of comparison) were also conducted using specimens of 100 mm height and 50. Note that although for the latter set of results 13

ACCEPTED MANUSCRIPT (saturated soil testing at zero suction) two different specimen sizes were used (see Table 3), it was subsequently found that all strength results plotted as unique strength

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lines and hence the difference in the specimen size did not have an effect on the

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results. Similarly, the results of the isotropic compression tests conducted on larger specimens (100 mm height and 50mm diameter) were confirmed to be consistent with the isotropic stage results of the smaller triaxial testing specimens (of 76 mm height

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and 38 mm diameter).

The soil was placed in the split moulds in six or eight equal layers (for the 76 mm and

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100 mm height specimens respectively) and compressed at a monotonic displacement

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rate of 1 mm/min until the required height was reached. The loading ram was then

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held in contact with the soil for another five minutes to reduce the rebound upon unloading (Jotisankasa 2005). This static compaction procedure was selected as the

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best way of exerting sufficient control over the compaction process of a clayey soil (Sivakumar, 1993; Sharma, 1998; Jotisankasa, 2005), so that almost identical specimens were prepared. It was observed that the two soils (treated vs. untreated) had different dynamic (standard Proctor compaction) characteristics due to the effect of lime treatment (see Table 2). Thus for consistent comparisons and due to the length of the tests limiting the amount of feasible investigations within the research time scales, a decision had to be made to keep either the same target compaction dry density or the same compactive effort for the two soils during static compaction, while acknowledging that the variation of either parameter would have some effect on the 14

ACCEPTED MANUSCRIPT resulting properties and behaviour of the compacted soils (see e.g. Chen et al. 1999; Sivakumar and Wheeler, 2000; Alonso and Pinyol, 2008). It was decided to compact

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statically all specimens at the same target dry density of 1.43 g/cm3 (corresponding to

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the maximum standard Proctor dry density of the untreated soil) and not at each soil’s (i.e. untreated vs. treated soil) Proctor optimum characteristics (although this resulted in a higher compactive effort for the lime-treated soil, i.e. 550 kPa for the untreated

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vs. 1000 kPa for the treated soil). The water content for both types of specimen was

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kept to the dry side of the respective Proctor optimum to ensure that the resulting structure after compaction was qualitatively similar for the two soils. Thus for the

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untreated London Clay the water content was about 25% (which is slightly drier of the

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Proctor optimum); for the lime treated specimens an additional 2% of water was used

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(also dry of the Proctor optimum of the lime treated soil), to ensure that enough water was available for chemical reactions, considering that the optimum water content of

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the treated soil was higher than that of the untreated soil (i.e. 30% for the treated vs 26% for the untreated soil). Using the filter paper method the suctions of the ascompacted specimens were determined and found to be fairly consistent i.e. approximately 650 kPa for untreated London Clay and 600 kPa for 4% lime treated London Clay specimens, with average as-compacted degrees of saturation of about 74 % and 78% respectively for the two soils.

After compaction, the lime treated specimens were left to cure for the required time (one week) in several layers of cling film and stored at controlled environmental 15

ACCEPTED MANUSCRIPT conditions (constant temperature and humidity). The required curing time of one week

was determined based on prior Unconsolidated Undrained (UU) triaxial testing. This

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showed that for this amount of lime (4%) and curing method, and for six different

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curing periods between 1 day and 166 days (the latter period covering the typical duration of the tests presented here), curing beyond seven days did not result in any

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of the cementation reactions (see Fig 2).

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further improvement of the shear strength of the soil, suggesting no further evolution

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Fig. 2 Deviator stress vs axial strain UU testing plots for different curing periods

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The characteristics of triaxial testing specimens after compaction as well as the list of

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tests presented are summarised in Table 3. Note that in places the tests will be discussed in comparison with results based on a set of triaxial tests conducted at s=

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0kPa. These were published elsewhere (Mavroulidou et al, 2011); hence they are not shown in detail in this paper.

Table 3 List of tests and specimen characteristics after compaction

Figure 3 shows indicative Soil Water Retention Curves derived from filter paper testing of compacted specimens (subject to subsequent saturation, followed by a drying and then a wetting path). These curves are included here to illustrate the partially saturated behaviour of the two soils (untreated and treated) and place these in 16

ACCEPTED MANUSCRIPT the context of the partially saturated soils analysis carried out in the following sections of the paper. From the figures it can be seen that according to the filter paper tests the

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compacted soils subjected subsequently to a wetting path are likely to be partially

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saturated in the range of suctions considered. It can also be seen that the treated soil appears to present a point of maximum curvature / ‘air entry value’ at lower suctions compared to the untreated soil (which is expected due to the flocculation of the

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particles induced by lime). Conversely the slope of the curves of the treated and

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untreated soils beyond these points of maximum curvature (which is linked to the

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microporosity of the soils) is similar.

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Fig. 3 Indicative Soil Water Retention Curves based on filter paper tests

3.3 Testing procedure

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As the suctions of the specimens at compaction were higher than the target suctions for triaxial testing, the specimens had to be wetted (following a suction unloading/decrease path) to achieve the target suction levels for testing, namely 0 kPa (saturation), 50, 100, 200 and 300 kPa respectively. A small mean net stress of 20 kPa was applied throughout the suction equalisation stage, to minimise specimen disturbance, avoid damaging of the membranes and ensure a good contact of the soil with the ceramic disk interface. Based on Sivakumar (1993) the suction equalisation stage was considered to have been completed when the rate of the pore water volume change was less than 0.044 cm3/d and 0.1 cm3/d for specimens of 38 mm diameter and 17

ACCEPTED MANUSCRIPT 76 mm height, and specimens of 50 mm diameter and 100 mm height respectively. When the above flow criteria were met, the particular suction level was maintained

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for a further 48 hours as an additional precaution, ensuring suction uniformity

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throughout the specimen. This also gave the specimen enough time to deform under the applied suction. The suction equalisation period typically varied from 10 to 30

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days depending on the suction level.

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After suction equalisation, the specimens were isotropically consolidated at the required mean net stress. The required loading was applied in one single step (ramped

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consolidation). During loading, the mean net stress (p-ua) was incremented at a

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sufficiently slow rate to minimise the development of excess pore water pressure and

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maintain constant suction during testing (Zhan, 2003). The maximum excess pore pressure uex induced in the partially saturated soil by the loading, was calculated as

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(Sivakumar, 1993):

uex 

RL h 2 2cv

(Eqn. 2)

where uex is the excess pore water pressure; RL is the loading rate of the ramped consolidation; h is the height of the specimen; cv is the coefficient of consolidation.

Taking into account the above equation and based on average cv values derived from preliminary tests, a rate of 0.6 kPa/h on the mean net stress was used for the isotropic 18

ACCEPTED MANUSCRIPT loading/unloading tests; its adequacy was confirmed by the monitoring of the pore water pressures during testing, which showed that this rate allowed for excess pore

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pressure dissipation during testing.

Following isotropic compression, a number of specimens were sheared. The majority of the specimens were sheared under constant suction (CS) (at constant suctions

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varying between 100-300 kPa); during these tests the specimens were sheared at a

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very slow rate of 1.14 μm/min, following a q/(p-ua)=3 path, while maintaining a constant cell pressure and a constant suction during the shearing (confirmed by the

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monitoring of the pore water pressures during shearing). The rate of shearing was

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consistent with rates used by other researchers (summarised in Delage, 2004). In

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addition, three constant water content (CW) tests were performed under undrained conditions at a constant rate of axial strain (2.42 µm/min). The faster rate adopted for

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the latter tests was justified as these were undrained tests. During shearing the pore water pressure was measured, and hence the suction changes were determined.

In addition to the triaxial tests, two suction unloading (wetting path) tests were performed in the triaxial cell under two different constant mean net stresses of 100 kPa and 200 kPa. They aimed at studying the effects of the mean net stress on the deformation of the lime-treated soil during wetting. After suction equalisation at the target suction, the specimens were compressed isotropically up to the target mean net stress of 100 kPa or 200 kPa (see column 5 of Table 3). During the suction unloading 19

ACCEPTED MANUSCRIPT /decrease tests (following a wetting path) performed in the triaxial cell, the above mean net stresses were then kept constant (and so was the air pressure) while the pore

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water pressure changed at a specified rate 1 kPa/hour to the respective target value

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resulting in a decrease in suction. This was followed by a 2-4 day period during which suction was maintained for excess pore water pressure equilibrium within the specimen. Water volume changes and overall volume changes were recorded during

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the testing. To find the suction decrease line, the results of these tests were

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complemented by filter paper results, not discussed here in detail as many of these were published elsewhere (Zhang et al, 2010; Mavroulidou et al., 2011; Mavroulidou

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et al, 2012; Mavroulidou et al, 2013a and b).

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4. Presentation of results

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4.1 Wetting induced swelling

The presented results refer to wetting induced-swelling from different tests and hence conditions, namely: zero mean net stress (based on filter paper results not detailed here), 20 kPa net stress (obtained from the suction equalisation stage preceding triaxial testing) and mean net stress of 100 kPa and 200 kPa (wetting soil water retention curves obtained using the triaxial cell). For the purposes of comparison the results of the wetting tests at all different net stresses in terms of specific volume versus matric suction (v-s), are plotted together in Fig. 4 (a) and (b). The latter plots the same results of the in a semi-logarithmic scale (v: ln[(s+pat)/pat)], normalised for a 20

ACCEPTED MANUSCRIPT reference pressure pat (atmospheric pressure). This representation will be used for the analyses of the results shown in section 5.1. To indicate qualitatively the differences

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during swelling between the untreated and the lime-treated soil, specific volume

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versus suction results from filter paper testing and some available results of wetting under 20 kPa net stress of untreated London Clay soil are also added. From Figures 4(a)-(b) it can be seen that although the lime-treated soil still shows some swelling

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under low confinement stresses, this is considerably reduced across the suction ranges

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considered, if compared to the untreated London Clay soil.

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Fig. 4. Specific volume vs. suction: (a) arithmetic scale; (b) semi-logarithmic scale

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Based on the curves of the wetting path results from the filter paper specimens (i.e. at a zero mean net stress), it can be seen that upon wetting, the specific volume initially

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appears to increase rather linearly, but after a suction of approximately 200 kPa was reached, some considerable swelling was observed, indicating yielding due to wetting. The specific volume after wetting versus suction under the 20 kPa mean net stress curves (i.e. from the suction equalisation stage in the triaxial test), shows deformation patterns similar to those of the tests under zero mean net stress tests but the suction yield stress point can now be identified at about 180 kPa. The maximum specific volume is about 2.06 (i.e. much smaller than that for tests under zero net stress) as swelling was partly suppressed due to the higher mean net stress. Specimens LTLCW1 and LTLC-W2 subject to wetting under a mean net stress of 100 and 200 kPa 21

ACCEPTED MANUSCRIPT respectively, showed swelling volumetric strains of 0.7% and 0.3% for a mean net stress of 100 kPa and 200 kPa, respectively, i.e., with an increase in the mean net

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stress, the swelling deformation decreased. For the specimen LTLC-W1 tested at 100

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kPa mean net stress, a likely yield point during wetting was identified at about 90-100 kPa suction corresponding to an increase in the rate of change of the specific volume during wetting (swelling). For the specimen tested at a mean net stress of 200 kPa, the

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yield suction was not possible to identify as the swelling deformation curve was

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linear; it will therefore be assumed to be zero in the analysis of the results (see section

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4.2 Isotropic compression

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5.1).

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Figure 5 presents isotropic compression results in terms of specific volume against mean net stress for both untreated and lime-treated specimens. The exact positions of

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the starting points in the plot were controlled by the specific volumes of the specimens at the end of the suction equalisation stage (i.e., after swelling) shown in Fig 4(a) and (b)).

Fig. 5. Comparative isotropic compression results

Despite the limited pressure ranges, yield stress points (noted

c)

were identified using

Casagrande’s method (see Table 4). These yield stress points were subsequently used to establish the LC yield curve in the s: space, as discussed later (see section 5.2). 22

ACCEPTED MANUSCRIPT From Table 4, it can be seen that the yield stresses of the lime-treated soil were higher than that of the untreated soil at the same suction level, which is consistent with the

IP

T

behaviour of bonded geomaterials. It can also be seen that, as for the untreated

SC R

partially saturated soil, yield stress increased with suction. However, although suction increased the yield stress for both soils, for the lime-treated soil the additional effect of chemically-induced bonding was proven to be considerable, comparing the yield

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stress values of the two soils at the same suction level. Table 4 also shows the values

MA

of N(s) and λ(s) (i.e. respectively the intercept at the reference pressure and the slope of the normal compression line) for different suctions. For uncemented partially

D

saturated soils, these values are commonly used to describe the post-yield

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relationship, i.e.:

TE

compression curves, for which Wheeler and Sivakumar (1995) suggested a linear

v  N ( s)   ( s) ln

p pat

(Eqn. 3)

AC

where the atmospheric pressure, pat =100 kPa, was used as a reference pressure to make the expression dimensionally consistent;

is the mean net stress (p-uα); p is the

mean total stress and uα the air pressure.

Note that the slope λ(s) value for the lime-treated soil was not clearly identifiable as the curves were slightly non-linear. This is realistic because, if extrapolated to higher mean net stresses that the available equipment for this study could not match, the gradient λ(s) during the de-bonding processing would be expected to continually increase with the mean net stress until the cemented soil curve coincides with that of 23

ACCEPTED MANUSCRIPT the uncemented soil when the breakage of all cementation bonds is complete. In this study, the compressibility line of the cemented lime-treated soil did not eventually

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converge to that of the untreated soil. This implies that the range of pressures applied,

SC R

only captured perhaps some initial yield linked to the beginning of a gradual, partial breakage of the cementation bonds but not that corresponding to full destructuration. According to Rao & Shivananda (2005), who tested saturated lime-treated black

NU

cotton soil specimens, a more dramatic yield point corresponding to complete

MA

destructuration was reached at pressures of 3-13 MPa, far beyond the ranges

D

achievable by the available equipment in the current study.

TE

An interesting observation made based on Figure 5, is that for the lime-treated soil

CE P

there is no clear indication that suction allows the Normal Compression Line (NCL) to cross the zero suction NCL of the lime-treated soil (assuming extrapolated NCL

AC

curves due to the limited extent of the actual experimental curve). This is unlike the behaviour of uncemented partially saturated soils reported in the literature (e.g. Alonso et al., 1990; Wheeler & Sivakumar, 1995), which was also noted for the untreated London Clay soil tested here (see Fig. 5). The isotropic compression stage results of all subsequent shearing triaxial tests (not shown here for the sake of brevity) based on additional specimens, were very consistent with the independent isotropic compression testing results shown in Figure 5, confirming the findings. Moreover the findings are consistent with those obtained from suction-controlled oedometer apparatus (K0 compression) testing for the same soils (which reached higher 24

ACCEPTED MANUSCRIPT compression levels, up to 2000 kPa) published previously (Mavroulidou et al 2013a). These also showed that the partially saturated lime-treated soil NCL did not cross the

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zero suction NCL. This would indicate that the behaviour of the lime-treated soil in

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isotropic compression for the ranges of suction studied was largely controlled by the lime-induced bonding rather than suction. For this statement to be conclusive the data

MA

the triaxial equipment used in this study.

NU

needs to be extrapolated to yet higher pressures, which were beyond the capacity of

Overall, based on the results in Table 4 and Figure 5, it can be concluded that lime

D

treatment generally had a favourable effect on the compressibility behaviour of the

CE P

TE

material compared to the untreated soil.

AC

Table 4 Soil parameters from isotropic compression data

4.3 Shearing

The results are presented in terms of peak and ultimate (end of test) strengths, also considering the critical state line of the untreated London Clay soil. Due to the number of results, these are plotted in several figures for the sake of clarity. A summary of the results is also given in Table 5.

Table 5. Results of suction-controlled shearing tests

25

ACCEPTED MANUSCRIPT Figures 6(a)-(d) show comparative stress-strain plots, namely: 6 (a) shows limetreated specimens sheared under constant suction (CS) at suctions of 100 and 200 kPa

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and varying mean net stresses; 6(b) plots constant suction (CS) testing results for

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specimens sheared at a suction of 300 kPa and varying mean net stresses, compared with the respective constant water content (CW) results for specimens subjected to an initial suction of 300 kPa; 6(c) and (d) compare the lime treated soil results with

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available results from untreated London Clay specimens to assess the differences in

MA

strength and behaviour.

D

Fig. 6. Deviator stress- axial strain plots (a) CS tests of lime-treated soil for suctions

TE

of 100 and 200 kPa; (b) comparative CS and CW results for suctions of 300 kPa; (c)

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comparative results for untreated and lime treated soil for p' or

of 200 kPa

AC

comparative results for untreated and lime treated soil for p' or

of 100 kPa (d)

From Figures 6(a)-(d), it can be seen that whereas the untreated compacted London Clay specimens show a strain hardening behaviour without any apparent peak in the stress, all lime-treated specimens show a very pronounced peak in the strength within relatively low strain levels; this was followed by strain softening presumably due to the breakage of cementation bonds, with the stress decreasing dramatically after only 2–5.6 % of axial strain. The lime treatment thus causes a similar behaviour to that of overconsolidated clays, which is consistent with the observations by Leroueil and Vaughan (1990) regarding similarities in the behaviour between naturally cemented 26

ACCEPTED MANUSCRIPT geomaterials (soft rocks) and overconsolidated clays. The elastic behaviour range is

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easily identifiable in the loading stage within 1 % strain.

SC R

Increasing suction resulted in an increase in the peak strength and stiffness of the lime-treated London Clay. Whereas this is the expected behaviour of partially saturated uncemented soils due to the effect of suction, in this instance, the behaviour

NU

is likely to be due to the combined effect of suction and the cementation bonds. Both

MA

types of soil approach constant values of deviator stress as the tests proceed (which is also consistent with the behaviour noted from the volumetric strain curves and the

D

pore pressures for the constant suction and constant water content tests respectively,

TE

shown in Fig. 7a and b and discussed later). The lime-treated specimens tested under

CE P

the same mean net stress reached essentially the same ultimate strength irrespective of the suction level considering the usual scatter in the experimental results (with some

AC

anomaly in one point of the 200 kPa suction results from specimen LTLC-CS5). This is also clearly depicted in Fig. 8, plotting the deviator stress versus matric suction results as well as the q-p´ results plotted in Fig 14 in section 5.3 below. This is consistent with the observed behaviour of uncemented partially saturated soils (e.g. Cui and Delage, 1996 and Hamid, 2008). Conversely the peak deviator stress is shown to clearly increase with mean net stress.

The stress-strain curves obtained from constant water content (CW) tests are similar to those obtained from constant suction (CS) shearing; the peak strengths are slightly 27

ACCEPTED MANUSCRIPT lower compared to the CS tests (see Fig. 6(b)). This could be attributed to the reduction in the matric suction, consistent with the increase in pore water pressure

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T

when reaching the peak (plotted in Figure 7 (b)). In the post peak region the pore

SC R

pressures gradually reduced until ultimate stresses were achieved. These ultimate stresses were essentially the same as those of the respective CS specimens, with a discrepancy noted between the LTLC-CS8 and LTLC-CW2 only (at 200 kPa suction),

MA

NU

which is difficult to explain.

Figure 7(a) presents selected volumetric strain vs. axial strain results. The reason why

D

the latter plots are shown for some of the specimens only is because not all triaxial

TE

systems used in this study could give reliable volume change measurements for the

CE P

brittle specimens. Namely two of the triaxial systems were equipped with local LVDT transducers whereas one system used the newly developed and validated laser sensor

AC

volume change measurement apparatus presented in Zhang et al (2012) and Zhang et al (2014). The former devices were not adequate in measuring post peak volume changes of brittle specimens as they soon lost alignment and their measurements became meaningless (see Zhang et al, 2012 and Zhang et al, 2014). Hence only the volumetric strain data based on the latter system (able to cover the whole range of strains involved in the testing) are shown in this paper.

Fig. 7 (a) Volumetric strain and (b) pore water pressure variation with axial strain

28

ACCEPTED MANUSCRIPT The volumetric strain vs. axial strain results (Fig. 7(a)) show that as opposed to the untreated London Clay specimens, which were continuously contracting (consistently

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T

with the stress–strain curves), the lime-treated specimens showed an initial

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contraction followed by some dilation. The amount of contraction decreased with increasing suction. On the other hand, for the same level of mean net stress, dilatancy increased with increasing suction, whereas for the same suction, higher net stresses

NU

appear to increasingly suppress dilatancy of the specimens, which is consistent with

MA

the reported behaviour of partially saturated uncemented soils (e.g. Hamid, 2008). With continuous shearing, the majority of the specimens appear to tend towards

D

constant total volumetric strains, which is consistent with the stress–strain behaviour

TE

of the soils presented in Fig. 6(a)-(c). Note that the lime-treated specimens showed the

CE P

dilating tendency after, rather than before the peak stress, i.e., the peak strength was mobilised well before the maximum rate of dilation. This conforms with published

AC

results on lithified materials (Leroueil and Vaughan, 1990; Vaughan, 1993) and implies that the extra component of strength manifested as peak strength is not due to dilatancy, as it would be the case for an uncemented particulate material but instead due to the cementation bonding created by lime treatment. It can thus be argued that as dilation only happens upon softening, this is consistent with the breakage of the cementation bonds created by lime treatment.

Figure 8 shows plots of peak and ultimate deviator stress versus matric suction. The figures are complemented with peak stress results for the saturated lime-treated soil, 29

ACCEPTED MANUSCRIPT based on results presented elsewhere (Mavroulidou et al. 2011 and 2013b). Figure 8 clearly illustrates the dependence of the peak deviator stress on the suction level and

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suggests a non-linear relationship between peak shear strength and matric suction,

SC R

whereas the ultimate deviator stress is shown to be practically unaffected by the levels of suction considered in this study (there is some variation and one anomaly for the 200 kPa suction-200 kPa net stress results -noted earlier- but other than this, the

MA

usual scatter of the experimental results).

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variation in the rest of the results does not appear to be considerable, in view of the

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Fig. 8 Peak and ultimate deviator stress envelopes for various suction levels

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5. Mathematical description of the results 5.1 Wetting induced swelling: Identification of the Suction Decrease (SD) yield line

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The analysis refers to the results of the suction unloading (suction decrease due to wetting) tests of the lime-treated soil plotted in Figure 4(b). The results will be analysed using the concept of a suction decrease (SD) yield locus postulated in the well-known BBM model (Alonso et al. 1990), to refer to the occurrence of plastic deformation due to wetting. From Figure 4(b) it can be seen that for low mean net stresses the resulting curve tends to be bilinear; the ‘turning point’ between the two lines can be defined as the yield suction (Zhan, 2003). The bilinear relationship was identified as: v  vs   s ln(

s  Pat ) within the elastic zone pat

(Eqn. 4a) 30

ACCEPTED MANUSCRIPT

s  Pat ) within the plastic zone pat

(Eqn. 4b)

T

v  N s  s ln(

IP

where  s and s are stiffness parameters with respect to a change in matric suction in

SC R

the elastic and plastic zone, respectively; vκs is the specific volume corresponding to

NU

the yield suction and Ns is the specific volume at zero suction.

Combining the yield suction at different mean net stress obtained from the various

MA

tests (see Fig. 4(a) and (b)), the SD yield line s  p  200 was plotted in Figure 9.

D

The inclination of the SD yield line shows the effect of the mean net stress on the

TE

yielding of the lime treated London Clay upon wetting. The assumption of zero yield at 200 kPa net stress gives a straight SD yield line inclined at 45° in the s: plane as

CE P

initially suggested by Gens & Alonso (1992), and unlike data for expansive soils, which can give inclination angles of the SD yield line larger than 45° (e.g. Zhan,

AC

2003). This is consistent with the improvement brought about by lime in terms of volumetric stability upon wetting noted in section 4.1 above, which showed that the nature of the initially shrinking/swelling London Clay soil changed to that of a soil of low swelling potential due to lime treatment.

Fig. 9 SD yield line for lime-treated London Clay 5.2 Isotropic compression: Identification of the Loading Collapse (LC) yield curve Consistently with the analysis presented in the previous section, using the concept of the SD yield curve, the isotropic compression results will be analysed using the 31

ACCEPTED MANUSCRIPT concept of the Loading- Collapse (LC) yield curve, postulated in the BBM model (Alonso et al. 1990). This curve is defined from the yield points of specimens with

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identical stress histories subjected to isotropic loading under constant suction. The LC

hardening behaviour (Gens & Alonso, 1992).

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will expand once the load applied exceeds the yield stress, reflecting the suction

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The Loading Collapse (LC) yield curves of the lime treated soil based on the isotropic

MA

loading tests were identified according to the approach proposed by Alonso (1990) (see Fig 10). The shape of the resulting LC yield curve was consistent with that

D

proposed by Alonso et al. (1990). Although only three points were available for the

TE

untreated soil these were also used to draw a tentative LC curve for the purposes of

CE P

comparison. The two curves reflect the considerable increase in the yield stress of the lime treated London Clay compared to that of the untreated London Clay and for this

AC

reason the lime-treated LC curve moved to the right-hand side of the untreated soil LC. The increased yield stress of the lime-treated soil indicates that the elastic range is enlarged, which can be attributed to the effect of the lime on the soil structure /cementation bonding.

Fig. 10 LC yield curves of lime-treated and untreated London Clay based on isotropic compression tests

To express the LC curve mathematically, both the BBM model (Alonso et al., 1990) 32

ACCEPTED MANUSCRIPT and Wheeler and Shivakumar’s (1995) model involve the parameter λ(s). However as discussed above the value of λ(s) was not easily identifiable due to the non-linearity

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of the NCL of the lime-treated soil and due to the fact that the slopes λ(s) for lime-

SC R

treated soil found in this study are likely to be related to initial yield rather than yield associated with the complete breakage of cementation bonds. An expression that would not involve the post yield parameter λ(s) was therefore introduced for the

NU

description of the presented experimental data. This reflects the fact that the isotropic

MA

yield behaviour is associated with both the loading and wetting paths. Therefore both paths were considered in the suggested alternative mathematical expression of the LC

D

curve. To demonstrate the procedure, data of the wetting path under a mean net stress

TE

of 20 kPa was used; first the relationship of wetting-induced swelling with respect to

CE P

the decrease in suction in the v:s space from data plotted in Figure 11(a) was established (in this figure the plotted v values represent the final specific volumes

AC

recorded during wetting at each suction level). To express this relationship mathematically two different fitting expressions were tried, namely a bilinear and a double exponential expression: 

Case A: bilinear fitting v  A1  B1 ln(

s  pat s  pat ) , ln( )  sd pat pat

v  A2  B2 ln(

s  pat ), pat

ln(

s  pat )  sd pat

(Eqn. 5a)

(Eqn. 5b)

where A1 and A2 are the intercepts at suction s=0 kPa of the lines defined in Eqn 5a and 5b respectively; B1 and B2 are the slopes of the lines defined in Eqn 5a and 5b 33

ACCEPTED MANUSCRIPT respectively; sd is the yield suction factor during the wetting where pat is the

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atmospheric air pressure (100 kPa).

Based on the experimental data, the resulting curve-fitting equations were as follows: ln(

s  pat )  1.03 pat

s  pat s  pat ) , ln( )  1.03 pat pat

(Eqn. 6a)

(Eqn. 6b)

Case B: double exponential function fitting:

s  pat s  pat )  a2 exp(b2 ) pat pat

(Eqn. 7)

TE

v  a1 exp(b1

D



MA

NU

v  2.045  0.092 ln(

s  pat ), pat

SC R

v  1.9876  0.03 ln(

CE P

where a1 , a2 , b1 and b2 are curve-fitting parameters.

Using units of kPa for the suction and atmospheric pressure, the curve-fitting gave the

AC

following expression:

(Eqn. 8) The final specific volume recorded during wetting is the initial specific volume prior to the isotropic compression. This can now be plotted against mean net stress. The yield points identified using Casagrande’s method presented in Table 4 are also added on the graph (Fig. 11(b)). From this figure, it can be noted that in the v:

plane, the

yield points can be reasonably assumed to fall on a straight line. This line can be fitted as: 34

ACCEPTED MANUSCRIPT  p   v  N y   y ln   pat 

(Eqn. 9)

T

where Ny and λy are respectively the intercept on the yield compression line at a

IP

reference pressure of pat and is the slope of the yield compression line in v: plane.

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For the case of the lime-treated soil the fitted line was found to be:  p   v  2.0625  0.086 ln   pat 

(Eqn. 10)

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Note that before yield the change in specific volume of the lime-treated soil was very

MA

small and was therefore ignored. Thus the specific volumes represented in Figure

D

11(a) (end of wetting stage volumes) are also assumed to be the same until the yield.

TE

Working with the above mentioned relationships in the v:s and v:

planes, an

CE P

expression relating directly suction and yield stress can be obtained and hence, the LC yield curve (Fig 11 (c)), thus described without the use of the parameter λ(s). For

AC

example using the double-exponential function (Eqn 7) and Eqn (9), and equating the right hand side parts of the two equations gives: a1 * exp(b1

 p  s  pat s  pat  )  a2 * exp(b2 )  N y  y ln  pat pat  pat 

(Eqn. 11)

Note that following the above procedure a relationship can also be determined between suction and isotropic compression yield mean net stress shown in Figure 11.

Fig. 11 Demonstration of an alternative procedure to mathematically describe the LC yield curve based on the wetting and isotropic loading paths

35

ACCEPTED MANUSCRIPT The predicted specific volume and suction relationship using the two different descriptions of the curve (bilinear and double exponential respectively) is shown in

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Figure 12. The predicted LC curves using the above procedure and the two different

SC R

descriptions of the v:s curves are shown in Figure 13. From the figures it can be seen that both expressions fit the available results very well; there are some small observed differences in the fitted curves based on the two different expressions at suctions

NU

higher than 100 kPa; more experimental data would be needed to confirm which

MA

expression gives the closest fit in this range.

TE

D

Fig. 12 Predicted specific volume vs. suction

CE P

Fig. 13 Predicted LC yield curve

AC

5.3 Shear behaviour of the lime-treated soil For the sake of convenience the shearing testing results will be interpreted by considering separately the contributions from mean net stress and suction. The peak strength and ultimate (end of test) strengths identified in section 4.3 above will be used for the analysis, together with the critical state line of the untreated London Clay soil.

Based on the values of Table 5, the Peak Strength Line (PSL) and Ultimate Strength Line (USL) were therefore plotted in Figure 14. On the same graph, the results of 36

ACCEPTED MANUSCRIPT saturated lime treated and untreated London Clay specimens published elsewhere (Mavroulidou et al, 2011) were also included for the sake of comparison between

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saturated and partially saturated states. According to Mavroulidou et al (2011), the

SC R

saturated lime-treated soil specimens also showed a strain softening behaviour whereas the untreated soil presented a typical hardening behaviour. The volumetric strain behaviour upon shearing confirmed this. For the saturated lime-treated soil it

NU

was found that the peak angle of friction was  p =26.5°, whereas the results for both

MA

soils (treated and untreated) could be expressed by a unique Critical State line in the q:p' plane, with a slope of M=0.88 corresponding to an angle of internal friction of  c

D

=22.5˚ (Mavroulidou et al, 2011). This value is consistent with values reported in the

TE

literature for London Clay (Schofield and Wroth, 1968). This implies that lime did not

CE P

affect the frictional characteristics of the clay soil and that the observed improvement in the shear strength manifested in terms of peak strengths was thus due to the

AC

differences in the soil structure before and after lime treatment (including bonding due to cementing compounds induced by lime).

For the partially saturated lime-treated soil (Fig. 14) the ultimate strength lines according to the results appear to be slightly different than the CSL of the saturated soils; from linear regression the slopes of the USLs were found to be 0.82, 0.86 and 0.95 for suctions s, of 100 kPa, 200 kPa and 300 kPa respectively (presumably due to suction effects) giving however an average slope of 0.88 which is the value of the

37

ACCEPTED MANUSCRIPT critical state parameter M of the saturated untreated (and treated) soil1. Note that in the original BBM model, Alonso et al (1990) suggest that a critical state line (CSL)

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for non-zero suction will represent the increased strength induced by suction in terms

SC R

of an increase in the apparent cohesion, while maintaining the slope M of the CSL for saturated conditions. The intercepts of the best fit lines through the ultimate stress points for each suction level are found to decrease with suction i.e., about 50 kPa for

NU

s=100 kPa, 46 kPa for s=200 kPa (which is actually quite close to that of 100kPa and

D

kPa for s= 300 kPa (see Fig. 14).

MA

could be considered the same due to the usual scatter of experimental data), and 17

CE P

TE

Fig. 14 q-p' plots (drained shearing / CS tests)

The critical state of the partially saturated soil in the q-(p-uα) plane can be written as

AC

(Toll, 1990):

q= Mα (p - uα)+Mb(uα- uw)

(Eqn. 12)

where Mα is the critical state stress ratio with respect to the mean net stress (p - uα) and Mb is the critical state stress ratio with respect to the matric suction (uα- uw). Based on Alonso et al (1990) assumptions that Mα=M=0.88 (where M is the critical state parameter of the saturated soil) the values of Mb were calculated as shown in 1 The anomalous point of the LTLC-CS5 was excluded from the analysis of the results (curve fitting and determination of the parameter Mb) 38

ACCEPTED MANUSCRIPT Table 6. To fit the presented data, it was found that Mb was not a constant and generally decreased with suction (with one irregularity in the trends for the LTLC-02

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specimen, probably because the test terminated earlier than the others). This implies

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that there is a drop in Mb with the gradual desaturation of the material as reported elsewhere for untreated partially saturated soils (Vanapalli et al, 1996; Toll and Ong, 2003). Unlike the findings in Toll et al (2008) for an artificially bonded sand, showing

NU

that for  a =   (equivalent to Mα=M) the angle  b exceeded   which is unrealistic

MA

(hence Toll et al suggested that a value of  a >   should have been used to overcome this issue), here the value of Mα set equal to M gave sensible values of Mb, clearly

TE

D

lower than Mα.

AC

variables

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Table 6. Partially saturated lime-treated specimens: critical state values of the state

From Figure 14, linear regression gave the slopes of the PSLs as 0.73, 0.69 and 0.64 for suctions of 100 kPa, 200 kPa and 300 kPa respectively, i.e. the PSL slope decreased with suction. The PSL was much higher than the USL in each individual suction level, suggesting that the additional peak strength is mainly dependent on the lime-induced bonding.

39

ACCEPTED MANUSCRIPT For the mathematical description of the peak strength of partially saturated lime treated specimens we adapted Toll’s (1990) critical state line expression for partially

IP

T

saturated soils. It was thus assumed that the peak strength of artificially cemented

SC R

specimens could be associated with three components: (a) a mean net stress component, (b) a suction component and (c) a cementation bonding component.

NU

Therefore, the PSL was represented by the following equation:

q  n ( p  u )  n (u  u )  f (c) a

a

b

a

w

(Eqn. 13)

MA

where the stress ratio parameters na and nb describe the components of shear strength associated respectively with the uncoupled stress state variables (p-ua) and (ua –uw) in

TE

D

a similar fashion as in Toll (1990) and Toll and Ong (2003); f(c) is the additional cementation bonding effect, which is a function of suction (because the difference in

CE P

peak values between the treated and untreated partially saturated soil was not constant for different suctions); f(c) can thus be expressed as:

f (c)  n (u  u )  q

AC

c

a

w

(Eqn. 14)

c

where nc is the suction-dependent cementation bonding component, applicable for partially saturated conditions, whereas qc is the cementation bonding (‘true’ cohesion) at saturated conditions and can be taken as 80 kPa, i.e., the value of the intercept of the PSL of the saturated lime treated soil in Figure 14.

Substitution of Eqn 14 into Eqn 13 yields:

q  n ( p  u )  (n  n )(u  u )  q a

a

b

c

a

w

c

(Eqn. 15)

40

ACCEPTED MANUSCRIPT where (ηb+ηc) is the stress ratio incorporating chemically-induced bonding and suction, whose effect is likely to be coupled 2. When suction is zero, the above

IP

T

equation reduces to the PSL expression for saturated soils. It should be noted that the

SC R

value of (ηb+ηc) is expected to vary for different lime percentages.

For the constant suction tests ηa was found from regression analysis of the

NU

experimental data for each specimen. The value of (ηb+ηc) was then determined from

MA

Eqn 15. The average value of (ηb+ηc) at each suction level was then calculated and used for the prediction of the peak strength according to Eqn 15 (note that ηa was not

D

averaged). The calculated parameters ηa and mean (ηb+ηc) and resulting predictions of

CE P

TE

deviator stresses for the constant suction tests are given in Table 7.

For the constant water content tests the parameters ηa and average (ηb+ηc) were

AC

determined as a function of suction (and not a function of the degree of saturation, Sr , as in Toll, 1990 due to the narrow ranges of Sr considered in this study). The experimental data and fitting curves for ηa and mean (ηb+ηc) are plotted in Figures 15 and 16, respectively and their values shown in Table 7. The mathematical expression of ηa and mean (ηb+ηc) according to the experimental curve-fitting of the results are respectively:

n  0.0005 * (u  u )  0.7873 a

a

w

(Eqn. 16)

2 To determine each one of these two parameters nb and nc individually, a relationship between cementation bonding component and suction is needed, which would necessitate a large number of data on the untreated partially saturated soil, in addition to those of the lime-treated partially saturated soil. 41

ACCEPTED MANUSCRIPT n  n  55.42 * (u  u ) b

c

a

0.606

(Eqn. 17)

w

Fig. 16 Variation of term (nb + nc) with suction

SC R

IP

T

Fig. 15 Variation of parameter nα with suction

NU

Using the above expressions, the peak deviator stresses qp of the three CW specimens

MA

were successfully predicted (with a maximum error of -1.37%), (see Table 8 and

TE

D

Figure 17 depicting the predicted and the measured peak deviator stresses).

CE P

Table 7. Values of ηα, (ηb+ηc) and predicted peak deviator stresses qp for lime treated

AC

specimens (CS tests)

Table 8. Predicted peak deviator stresses qp for lime treated specimens (CW tests)

Fig. 17 Predicted peak deviator stress qp

6. Conclusions

The mechanical behaviour of lime treated London Clay during wetting, compression and shearing was investigated via suction-controlled triaxial testing. The paper 42

ACCEPTED MANUSCRIPT provided valuable data of suction controlled testing of lime-treated soils and addressed the corresponding behaviour of the lime-treated soil, within an unsaturated

IP

T

soil mechanics framework. Although the results presented are specific to London

SC R

Clay, experience indicates that the behaviour of other highly plastic clays treated with lime will follow similar trends.

NU

Overall the experimental results showed the beneficial effect of the lime on the

MA

volumetric properties (reduced swelling upon wetting, increased yield stress and reduced compressibility upon compression) and shear strength within ranges of strain

D

relevant to engineering design (despite the observed strain softening behaviour). They

TE

also indicated that for the range of suctions considered in this study chemically

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induced bonding had a major effect on the properties and behaviour of the soil.

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It was also shown that although the properties of the soil have obviously changed upon lime treatment, the behaviour of the partially saturated lime treated soil presented behaviour trends consistent with the reported behaviour of uncemented partially saturated soils regarding the effect of suction or mean net stress, with the exception of the compressibility behaviour of the lime-treated soil. It was thus shown that the partially saturated lime treated soil could be described within the framework of constitutive models commonly used for untreated partially saturated soils with only occasional modifications. Thus, some simple procedures for the analysis of the results of this type of soil were suggested and successfully applied to the mathematical 43

ACCEPTED MANUSCRIPT description of the results. These consisted in (a) an alternative, simple method of determining the LC yield equation combining the wetting induced swelling curve and

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the LC yield compression lines and (b) an expression used to interpret the peak

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strengths of the partially saturated lime treated soil to account for the combined effect of mean net stress, suction and cementation bonding.

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Acknowledgements

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The work presented in this paper was carried out at London South Bank University during the doctoral studies of the first author, funded by the UK Engineering and

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Physical Sciences Research Council (EPSRC) through grant EP/E037305/1.

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ACCEPTED MANUSCRIPT Table 1 Composition of London Clay soil used in this study

Clay content %

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50 26 15 9 4% 45%

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Illite (%) Smectite (%) Kaolinite (%) Chlorite (%) Sand Silt

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of which:

(%)

(%)

64

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Lime-treated

wP

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London Clay

wL

26

Standard Proctor wopt (%)

Standard Proctor ρdmax (g/cm3)

2.75

26

1.43

2.73

30

1.26

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GS

38

35

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Soil Type

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Table 2 Physical characteristics of untreated and 4% lime treated London Clay soils

89

54

London Clay

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(4% lime)

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Specimen Dia. mm

Suction level, kPa

p' or (p-ua), kPa

Stress Path

Initial conditions

Type

w (%)

e

Sr (%)

0%

50

0

100

δσ3=0

Drained 24.76 0.930

73.2

LC-SAT2

0%

38

0

200

δσ3=0

Drained 25.07 0.921

74.9

LC-IS1

0%

38

0

5-570-5

N/A

N/A

24.95 0.928

73.9

LC-IS2

0%

38

300

200

N/A

N/A

24.80 0.923

73.9

LC-CS1

0%

38

200

100

δσ3=0

Drained

LC-CS2

0%

38

300

200

δσ3=0

Drained 24.8

LTLC-IS1

4%

38

0

300

N/A

LTLC-IS2

4%

38

50

20-400-20

LTLC-IS3

4%

38

100

LTLC-IS4

4%

50

LTLC-IS5

4%

50

LTLC-CS1

4%

38

LTLC-CS2

4%

38

LTLC-CS3

4%

LTLC-CS4

4%

LTLC-CS5

4%

LTLC-CS6

4%

LTLC-CS7

4%

LTLC-CS8

4%

LTLC-CS9

4%

LTLC-CW1

4%

LTLC-CW2

4%

38

LTLC-CW3

4%

38

LTLC-W1

4%

50

LTLC-W2

4%

50

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LC-SAT1

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Lime % Specimen ID

Shearing

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Consolidation/ Compression

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Table 3 List of tests and specimen characteristics after compaction

73.8

0.923

73.9

N/A

27.00 0.918

80.2

N/A

N/A

26.87 0.931

78.6

20-530-23-100

N/A

N/A

26.91 0.941

77.9

200

20-1000-20

N/A

N/A

26.97 0.933

78.8

300

20-1200

N/A

N/A

26.92 0.925

79.3

100

δσ3=0

CS

26.93 0.930

78.9

200

δσ3=0

CS

26.82 0.928

78.8

38

300

δσ3=0

CS

26.97 0.932

78.9

38

100

δσ3=0

CS

26.82 0.931

78.5

200

δσ3=0

CS

26.69 0.938

77.5

300

δσ3=0

CS

26.82 0.931

78.5

100

δσ3=0

CS

26.78 0.928

78.6

200

δσ3=0

CS

26.89 0.931

78.7

38

300

δσ3=0

CS

26.78 0.931

78.4

38

100

δσ3=0

CW

27.00 0.940

78.3

200

δσ3=0

CW

26.80 0.931

78.4

300

δσ3=0

CW

26.60 0.931

77.9

100

N/A*

N/A

26.60 0.945

76.7

200

N/A*

N/A

26.90 0.938

77.0

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0.931

100

200

38 38 38

300

300

500-0 (wetting path)

25

*These were Soil Water Retention Curve tests demonstrating the effect of stress state (see Fig 4a and b)

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 (s)

N (s)

0 200 300 0 50 100 200 300

35 80 180 130 180 260 340 400

0.139 N/A 0.047 0.064 0.027 0.031 0.058 0.065

2.27 N/A 2.16 2.041 2.028 2.014 2.048 2.049

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s (kPa)

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Specimen LC-IS1 LC-CS1* LC-IS2 LTLC-IS1 LTLC-IS2 LTLC-IS3 LTLC-IS4 LTLC-IS5

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*Specimen LC-CS1 was used to complement the data for the yield stress but the λ(s)

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and N(s) could not be determined due to the limited extent of the data

LTLC-CS0* LTLC-CS4 LTLC-CS5

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LTLC-CS1 LTLC-CS2 LTLC-CS3

Peak Strain range (%) for Mean net Ratio of peak Suction deviator peak strength stress /ultimate (kPa) stress (kPa) deviator stress (kPa) 3% 100 689 3.59 2% 100 200 791 2.95 1.62%, 300 880 2.11 4.2% 50 668 1.67

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Specimen ID

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Table 5. Results of suction-controlled shearing tests

784 849

LTLC-CS6

300

966

3% 2.66% 2.53%

LTLC-CS7 LTLC-CS8

100 200

836 957

5% 5.6

300

1032

3.96%

300

870 945 1015

1.6%, 3.1% 4.2%

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100 200

200

300

LTLC-CS9 LTLC-CW1

100

LTLC-CW2

200

LTLC-CW3

300

4.22 2.53 2.25 4.80 3.38 2.29 4.22 2.53 2.25

*Not plotted in Figures 6(a)-(b) for the sake of clarity in the presentation of the plots

Table 6. Partially saturated lime-treated specimens: critical state values of the state 53

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Mα

Mb

0.88 0.88 0.88 0.88 0.88 0.88 0.88 0.88

0.49 0.14 0.31 0.22 0.20 0.12 0.08 0.18

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(uα -uw) (kPa) 100 100 100 200 200 300 300 300

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(p-uα) (kPa) 162 289 439 162 443 158 295 450

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q (kPa) 192 268 417 186 429 174 283 450

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variables

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Table 7. Values of ηα, (ηb+ηc) and predicted peak deviator stress qp for lime treated

LTLC-CS0

(ua-ua)p (kPa) 50

LTLC-CS1 LTLC-CS2

100

LTLC-CS4 LTLC-CS5 LTLC-CS6

200

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LTLC-CS7 LTLC-CS8

na regressed

(nb+nc) calculated

(nb+nc) mean

300

LTLC-CS9

qp(kPa) predicted

668

423

0.78

5.166

5.166

668

689

330

0.73

3.683

3.564

677

784

461

0.73

3.672

3.564

773

836

579

0.73

3.336

3.564

859

791

364

0.69

2.300

2.243

779

849

483

0.69

2.179

2.243

862

957

619

0.69

2.249

2.243

956

880

393

0.64

1.828

1.822

878

966

522

0.64

1.840

1.822

961

1032

644

0.64

1.799

1.822

1039

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LTLC-CS3

qp(kPa) (p-ua)p Measured (kPa)

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Specimen ID

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specimens (CS tests)

Table 8. Predicted peak deviator stresses qp for lime treated specimens (CW tests) Specimen ID

(ua-ua)p (kPa)

qp(kPa) Measured

LTLC-CW1

276

868

389

0.65

1.891

856.125

-1.37

LTLC-CW2

274

945

515

0.65

1.899

936.935

-0.85

LTLC-CW3

247

1010

637

0.65

2.021

1000.070

-0.98

(p-ua)p (kPa)

na (nb+nc) regressed mean

qp(kPa) predicted

Error (%)

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ACCEPTED MANUSCRIPT Highlights We investigated the mechanical behaviour of lime treated London Clay



We performed suction-controlled tests (wetting path, compression, triaxial

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

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testing) We described the behaviour of the lime-treated soil



This was done within an unsaturated soil mechanics framework



We proposed simple expressions for the analysis of results for this soil type

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

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