Shear strength of a pyroclastic soil measured in different testing devices

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1 Volcanic Rocks and Soils Rotonda et al. (eds) 2016 Taylor & Francis Group, London, ISBN Shear strength of a pyroclastic soil measured in different testing devices Sabatino Cuomo, Vito Foresta & Mariagiovanna Moscariello Laboratory of Geotechnics, Department of Civil Engineering, University of Salerno, Italy ABSTRACT: The paper deals with the shear strength of a Vesuvian pyroclastic soil (southern Italy), sampled in a district, where flowslides frequently occur. Direct shear tests, drained and undrained triaxial tests, as well as simple shear tests were carried out. The use of different testing devices was motivated by an extensive literature, which outlines some differences among the measured shear strength due to different failure mechanisms. The interpretation of the experimental results was made through distinct procedures, specific for each type of device. As main insight, some differences were pointed out for the effective friction angle (ϕ ) and for the stress ratio (M) at critical state, as computed in the deviatoric (q p ) plane for triaxial conditions. Particularly, the results of the simple shear tests, whether interpreted in the (t s ) plane, well agree the direct shear and triaxial tests, and the differences were related to the normalized displacement of the specimen at failure. 1 INTRODUCTION The pyroclastic soils are produced by the volcanic explosive eruptions, wind transportation and air-fall deposition along the hillslopes located nearby the volcanoes (Cascini et al., 2010). Pyroclastic deposits are layered (Cascini et al., 2011), and made of different coarse-grained soils (Cascini et al., 2013), most of which have a negligible effective cohesion and a considerable suction-related cohesion, especially in Mediterranean Countries. The (matric) suction is the difference of the air pressure (u a ) minus the pore water pressure (u w ). The latter is negative for a soil in unsaturated conditions. Is a matter of fact that the mechanical properties and the unsaturated conditions of the Campania pyroclastic soils allow stable slope conditions in steep areas. Nevertheless, rainfall infiltration reduces the soil suction, increases the soil water content, eventually up to full saturation. Once the soil reaches full saturation conditions, this extra suctionrelated shear strength vanishes and slope failures may be triggered. Particularly, the metastable structure of pyroclastic soils (Bilotta et al., 2005; Lancellotta et al., 2012) causes the pore water pressure to build-up in the post-failure stage. This corresponds to the onset of flow-type landslides, which result in casualties and damage to property (Cascini et al., 2010). The soil saturated shear strength envelope serves as an input for landslides triggering analysis, and it is usually expressed in terms of effective friction angle (ϕ ) and effective cohesion (c ). The latter is negligible for the pyroclastic soil tested in this paper. The soil friction angle is used in Limit Equilibrium Method (LEM) slope stability analyses, which schematize the soil as a rigid perfectly-plastic medium through a Mohr Coulomb, or Drucker-Prager failure criterion. Alternatively, stress-strain analyses may be used and advanced constitutive models must be adopted, which properly simulate the soil stress-path upon deformation. In this case, the soil shear strength envelope is more conveniently defined in the deviatoric (q p ) plane, through the stress ratio (M = q/p ) at critical state in triaxial conditions. In southern Italy, a large area (3,000 km 2 ) surrounding the Vesuvius volcano is characterized by steep carbonate reliefs covered by pyroclastic soils. In the paper, the shear strength of a Vesuvian pyroclastic soil, sampled in the Sarno-Quindici area (Cascini et al., 2008), was investigated. Purposely, direct shear (DS) tests, triaxial (TX) tests, and simple shear (SS) tests were carried out. The use of different testing devices was motivated by an extensive literature, which outlines differences among the results achievable from those equipments due to different failure mechanisms of the specimens (Randolph and Worth, 1981). However, further research is still needed about this topic. This paper firstly describes the equipments, the material tested, the experimental procedures and the testing programme developed. Then, the theoretical framework used for the results interpretation is illustrated, and the specific results of the DS, TX and SS tests are outlined and compared. Some experimental results were taken from former publications and reinterpreted, and the remaining tests were performed for this paper. The collection and the comparison of all the experimental results in a unitary theoretical framework guarantee the originality of this paper. 231

2 Table 1. List of the strain-controlled laboratory tests. Type # σ yy (kpa) p c (kpa) displ. rate (mm/min) DS / TX 8 / SS / DS: Direct Shear; TX: Triaxial; SS: Simple Shear; σ yy : effective vertical consolidation stress; p c : mean effective consolidation stress; displ. rate: displacement rate imposed during the test. 2 MATERIALS AND METHODS 2.1 Experimental programme The soil investigated is an air-fall volcanic (pyroclastic) soil, produced by the explosive activity ofvesuvius volcano. Specifically, this soil was formerly studied by Bilotta et al. (2005), and classified as class A ashy soil. It consists into a very porous, light, non-plastic sandy silt, which is unsaturated in the field, for most of the hydrological year. This soil is collapsible upon wetting, as during heavy or prolonged rainfall (Bilotta et al., 2008). The grain size distribution consists in 1.1 to 6.4% Gravel, 40.6 to 51.3% Sand, 40.9 to 53.61% Silt, 1.1 to 6.4% Clay. The soil specific gravity (G s ) is equal to 2.55; the specific volume (v) ranges from (undisturbed) to (remoulded); the saturation degree (S r ) is comprised between 74.8% (undisturbed specimens) and 92.1% (remoulded) (data from Bilotta et al., 2008); the dry unit weight (γ d ) is 6.93 to 8.65 kn/m 3 (respectively, undisturbed and remoulded). The average water content (w) is 51.9%, while the liquid limit (w L ) is 53.8% and the plastic limit (w P ) is 49.3%. The specimens used for the experimental programme were prepared from the remoulded soil. Specifically, the preparation included the mixing by hand of the soil at natural water content with distilled water up to produce a slurry with an initial water content equal to 1.5 times the liquid limit (w L ). Then, the slurry was statically compressed in a large consolidometer under an effective vertical stress equal to 10 kpa for simple shear tests, 80 kpa for triaxial tests, 100 kpa for direct shear tests. The consolidated soil was dried on air for one day and then used to sample the specimens. Doing so, the specimens were obtained at nearly saturated conditions (S r = 92.1%). The laboratory equipments used for soil mechanical characterization spanned over three distinct testing devices, and all of the tests were performed under displacement control (Tab. 1). All the Direct Shear (DS) tests were performed at a strain rate of mm/min, on specimens mm, for effective vertical stresses (σ yy ) ranging from 20 to 80 kpa. The latter corresponds to about twice the maximum in situ vertical stress. The triaxial (TX) tests were carried out at a strain rate of mm/min on cylindrical specimen mm (respectively, diameter and height). The mean effective isotropic compression stress was from 30 to 100 kpa. The TX tests were carried out in drained (7 tests) and undrained (1 test) conditions. Simple shear (SS) tests were carried out in a new equipment, recently designed at the University of Salerno (Sorbino et al., 2011; Cuomo et al., 2015), which allows testing cylindrical specimens mm (respectively, diameter and height) sized. The tests were performed in drained conditions, at strain rate of about mm/min on specimen consolidated at effective vertical stresses (σ yy ) ranging from 50 to 100 kpa, which are close to those adopted for DS tests. 2.2 Theoretical framework The experimental results were interpreted differently, depending on the type of test, while computing for all of them the corresponding soil friction angle. The effective cohesion is negligible for this soil and was disregarded. Particularly, the effective vertical stress (σ yy ), the shear stress on the horizontal plane (τ xy), and the effective principal stresses (σ 1, σ 2, σ 3 ), were considered. Doing so, the following stress variables can be defined: where σ is the effective stress tensor, σ is total stress tensor, p w is the pore water pressure, and I is the identity tensor of second order (eq. 1). Later on, q is the deviatoric stress, p is the mean stress acting on solid skeleton, t and s are two stress invariants independent on the intermediate stress (σ 2 ). For DS tests, the stress ratio (R), i.e. the shear stress on the horizontal plane (τ xy ) to the effective vertical stress (σ yy ), was used to individuate the shear strength envelopes. It reads: For the TX tests, the stress invariants q p (eq. 2 3) or the stress variables t s (eq. 4 5) can be equivalently used as σ 2 = σ 3. At failure it entails: 232

3 For the SS tests, the fundamental contributions of Roscoe (1967) and Wood et al. (1980) were referred for the interpretation of the results. According to Roscoe (1967), the principal axes of stress and strain increments coincide, when the applied shear strain is sufficiently high (i.e. the plastic flow should be developed). The rotation of the principal stress axis occurs when the stress ratio (R) increases. Budhu (1979) and later literature showed the two possible mechanisms associated to the failure planes: translation along a series of horizontal planes or translation along vertical planes with an associated rotation of those planes. Soil failure is associated to the mechanism which requires the lowest external applied load. In the first case, the failure is associated with horizontal failure plane and the stress ratio (R) is given by the eq. 6, like for direct shear tests. In the second case the failure might be associated with a vertical rupture and the stress ratio (R) at failure is given as: Table 2. List of the Direct Shear (DS) tests. σ yy (kpa) displ. rate # v 0 consolidation (mm/min) TAL TAL TAL TAL TAL In addition, for simple shear tests, Wood et al. (1980) demonstrated that it is possible determining the principal stresses (σ 1, σ 2, σ 3 ) acting on the solid skeleton as follows in the eqs : Figure 1. Experimental results of the Direct Shear tests. where the earth pressure coefficient at rest (k 0 ) can be estimated from Jaky s expression k 0 = 1 sin ϕ, and the stress ratio (R) at the critical state (i.e. continued shearing can occur without change of volume or stress) is noted as R cv. 3 DIRECT SHEAR TESTS The direct shear tests were performed on five specimens with initial void ratio ranging from 2.55 to The specimens were normally consolidated in k 0 conditions at 20, 34.8, 49, 65.4 and 79 kpa of effective vertical stress (σ yy ). The tests were carried out with a conventional direct shear apparatus, so it was not possible to measure the pore pressure and to control if any excess of pore water pressure arose during the failure. For these reasons the displacement rate was chosen low, and as a function of the consolidation time. In particular, the minimum necessary time to reach the failure (t R ) for Figure 2. Saturated shear strength envelope from DS tests. each test was estimated according to the equations proposed by ASTM (1979), Bowles (1970), Gibson and Henkel (1954). Then, the minimum displacement rate was calculated as the ratio between the displacement which correspond to the stress peak (S R ) and t R. The adopted displacement rate varied from to mm/min. All the tests exhibit a stress peak and then a strain softening (Fig. 1). The higher the effective vertical consolidation stress, the lower is the stress peak. Saturated shear envelope from the direct shear tests (Fig. 2) is drawn for the stress-strain state corresponding to the steady-state shear strength. The 233

Table 3. Details of the Triaxial compression (TX) tests. # test v 0 p cons (kpa) v cons BIS24_06 D 2.901 100 2.826 BIS26_06 D 2.934 50 2.885 BIS28_06 D 2.904 30 2.873 BIS29_06 D 2.916 30 2.

4 Table 3. Details of the Triaxial compression (TX) tests. # test v 0 p cons (kpa) v cons BIS24_06 D BIS26_06 D BIS28_06 D BIS29_06 D BIS25_06 U D: Drained TX tests, U: Undrained triaxial test Figure 4. Table 4. Saturated shear strength envelope from TX tests. Details of the Simple shear (SS) tests. Figure 3. Results of the triaxial compression (TX) tests. # test v 0 p cons (kpa) v cons SSP0115 CL SSP0225 CV SSP0315 CL CL: Constant Vertical Load, CV: Constant Volume. ( ) p varies during the test up to about 50 kpa at failure. formulation of Mohr-Coulomb is utilized to draw the shear envelope in τ xy σ yy plane. 4 TRIAXIAL TESTS The triaxial tests were performed both in drained and undrained conditions. All of the tests were consolidated in isotropical conditions at 30, 50 and 100 kpa (Tab. 3). After the consolidation, the tests were conducted in strain-controlled conditions. The axial stress was increased at a constant displacement rate of about mm/min. The drained triaxial tests exhibit a moderate strain softening after the peak stress. The test BIS_25_06 in undrained conditions showed a hardening behaviour. None of the tests reached the critical state within the investigated deviatoric strains (Fig. 3). Thus, the saturated shear envelope was drawn extrapolating (Fig. 4) the experimental stress-strain curve to a deviatoric strain equal to 40%, which corresponds for this soil to critical state conditions (Migliaro, 2008). 5 SIMPLE SHEAR TESTS The simple shear tests were carried out in drained conditions on saturated remoulded specimens. The tested specimens were firstly consolidated in k 0 -conditions at vertical stress equal to 69.8, 75.0 and kpa (Tab. 4). Then, the shearing of the specimen was performed in strain-controlled condition at a shear rate Figure 5. tests. Experimental results of the Simple Shear (SS) equal to about mm/min. This value of shear rate was imposed allowing the volumetric strains to develop during the shearing for CL tests. The CV test was performed controlling the vertical stress to inhibit the volume variation for the specimen.all of the results show a hardening behaviour. The shear strength envelope was obtained for the shear stress values corresponding to the maximum horizontal displacement reached during the tests. In this case, for this reason, steady state or critical state conditions cannot be directly checked within about 30% of shear strain. This circumstance can 234

5 Figure 6. Saturated shear strength envelope from SS tests. reflect on underestimated values of critical shear strength of the tested soil. 6 DISCUSSION The experimental results described in the previous sections allow some considerations. The different stress-strain behaviour observed in the DS,TX and SS tests is caused by different values of initial void ratio of the specimens. This last circumstance relates to different vertical stresses imposed during the consolidation stage of the soil samples preparation (see section 2.1). In particular, the target vertical stress consolidation values imposed to the soil were 100, 80 and 10 kpa, respectively for DS, TX and SS samples, at a later stage utilized to obtain the test soil specimens. As for the attainment of critical state conditions for each testing technique, different criteria were utilized, as described before. In particular, steady-state condition could not be directly checked within about 30% of shear strain for SS specimens, and this reflects on an underestimation of the shear strength. Independently on the testing device, the experimental tests are fully reproducible as it concerns both the specimen preparation procedure and the testing technique. This is shown by the excellent interpretation of the results through standard procedures, as reported in the figures 2, 4 and 6. The mean square error (R 2 )is higher than 0.93 for all of the testing devices. The comparison among the results achieved through different devices was the principal goal of the paper, and this issue was here tackled computing both the effective friction angle (ϕ ) and the stress ratio at critical state (M), as computed in the deviatoric (q p ) plane for triaxial conditions (Eq. 7). It is worth noting that the soil critical state is reached at different displacements (or strain, if definable) depending on the type of the testing device. For this reason, a specific normalized displacement was individuated for each type of device. For the direct shear device, the external horizontal displacement imposed to half-a-specimen was normalized to the size of the shear box (B = 60 mm). For the triaxial apparatus, the Figure 7. Values of the effective friction angle obtained from different devices (DS, TX, and SS) and through different interpretation of the results (the normalized displacement is assumed equal to δ h /B at steady-state strength for DS tests, δ a /H c at critical state for TX tests, and δ h /H at the final stage of SS tests). Figure 8. Values of the stress ratio (M = 6 sin ϕ / (3 sin ϕ )) obtained from different devices (DS, TX, and SS) and through different interpretation of the results (the normalized displacement is assumed equal to δ h /B at steady-state strength for DS tests, δ a /H c at critical state for TX tests, and δ h /H at the final stage of SS tests). external vertical displacement was normalized to the specimen height at the end of the consolidation stage (H c ). For the simple shear device, the external horizontal displacement was normalized to the specimen height (H = 22 mm). The results of the direct shear (DS) and triaxial (TX) tests were univocally elaborated and the results were close each other, both for ϕ (Fig. 7) and for M (Fig. 8), with an increasing trend related to the normalized displacement reached at critical state. The experimental evidence of the simple shear tests can be interpreted in different ways. Thus, the interpretation in the t s plane provided values for ϕ (Fig. 7) and M (Fig. 8), which were close to those of the other testing devices. Conversely, the interpretations of the SS results in the q p plane provided the lower values 235

6 than in the other equipments for both ϕ and M. This is probably due to the role played by the intermediate principal stress on the shear strength of the material. An underestimation of the soil shear strength in the τ xy σ yy is expected, as mentioned be, for SS tests. More in general, the full understanding of the deformation and failure modes in simple shear tests is an important issue, which still deserves further investigation. Related to this topic is the most appropriate approach for the interpretation of the experimental results of the SS tests. 7 CONCLUSIONS The paper dealt with the characterization of the soil shear strength of remoulded specimens of a Vesuvian pyroclastic soil, sampled in the Sarno-Quindici area (southern Italy), where huge flowslides frequently occur. Direct shear (DS) tests, triaxial (TX) tests, and simple shear (SS) tests were carried out to examine the differences in terms of friction angle (ϕ ) and the stress ratio at critical state (M). The paper showed that the experimental tests are fully reproducible as it concerns both the specimen preparation procedure and the testing technique. This was found independent on the testing device. Then, the comparison among the results achieved through different devices was tackled computing both the effective friction angle (ϕ ) and the stress ratio at critical state (M), in relation to a specific normalized displacement for each type of device: i) the external horizontal displacement normalized to the size of the shear box for direct shear device, ii) the external vertical displacement normalized to the specimen height, for the triaxial apparatus, iii) the external horizontal displacement normalized to the specimen height, for the simple shear device. The results of the direct shear (DS) and triaxial (TX) tests were univocally elaborated and the results were close each other, both for ϕ and for M, with an increasing trend related to the normalized displacement reached at critical state. Conversely, the interpretations of the SS results in the q p or in the τ xy σ yy planes provided lower values than in the other equipments for both ϕ and M. This issue require future experimental and theoretical investigations, to fully understand and characterize the failure mechanism occurring in the simple shear device, compared to those relative to direct shear and triaxial devices. Bilotta, E., Foresta, V., Migliaro, G., The influence of suction on stiffness, viscosity and collapse of some volcanic ashy soils. Unsaturated Soils: Advances in Geo- Engineering, Toll et al. (eds), ISBN , pp Budhu M., On comparing simple shear and triaxial test results. J. Geotech. Engrg : DOI: : Cascini L., Cuomo S., Guida D., Typical source areas of May 1998 flow-like mass movements in the Campania region, Southern Italy. Engineering Geology 96, DOI: /j.enggeo Cascini L., Cuomo S., Pastor M., Sorbino G Modelling of rainfall-induced shallow landslides of the flow-type. ASCE s Journal of Geotechnical and Geoenvironmental Engineering, No.1, pp DOI: /ASCEGT Cascini, L., Cuomo, S., Della Sala, M Spatial and temporal occurrence of rainfall-induced shallow landslides of flow type: A case of Sarno-Quindici, Italy. Geomorphology, Vol. 126, Issues 1 2, pp DOI: /j.geomorph Cascini, L., Sorbino, G., Cuomo, S., Ferlisi, S Seasonal effects of rainfall on the shallow pyroclastic deposits of the Campania region (southern Italy). Landslides, 11(4), , DOI: /s Cuomo S., Geomechanical modelling of triggering mechanisms for flow-like mass movements in pyroclastic soils. PhD Thesis, University of Salerno, pp Lancellotta, R., Di Prisco, C., Costanzo, D., Foti, S., Sorbino, G., Buscarnera, G., Cosentini, R.M., Foresta V., Caratterizzazione e modellazione geotecnica. In: Criteri di zonazione della suscettibilità e della pericolosità da frane innescate da eventi estremi (piogge e sisma)/ Leonardo Cascini. Composervice srl, Padova, pp ISBN (in Italian). Migliaro G., Il legame costitutivo dei terreni piroclastici per la modellazione di scavi in ambiente urbanizzato ed influenza della parziale saturazione. PhD thesis at University of Salerno (in Italian). Randolph M. F., Wroth C. P., Application of the failure state in undrained simple shear to the shaft capacity of driven piles. Géothecnique Vol. 31, No. 1, pp Saada A.S., Fries G., Ker C., An evaluation of LaboratoryTesting techniques in Soil Mechanics. Soils and Foundations, Japan Society of Soil Mechanics and Foundation Engineering, Vol. 23, No. 2, 1983, pp Thay S., Likitlersuang S., Pipatpongsa T., Monotonic and Cyclic Behavior of Chiang Mai Sand Under Simple Shear Mode. Geotechnical and. Geological Engineering., (2013) 31: DOI /s Wood D.M., Drescher A., Budhu M., On the Determination of Stress State in the Simple Shear Apparatus. Geotechnical Testing Journal, Vol. 2, No. 4, Dec pp REFERENCES Bilotta, E., Cascini, L., Foresta, V., Sorbino, G Geotechnical characterization of pyroclastic soils involved in huge flowslides. Geotechnical and Geological Engineering, 23: DOI: /S

SHEAR STRENGTH OF SOIL

Soil Failure Criteria SHEAR STRENGTH OF SOIL Knowledge about the shear strength of soil important for the analysis of: Bearing capacity of foundations, Slope stability, Lateral pressure on retaining structures,