The large displacement shear characteristics of granular media against concrete and steel interfaces

Barmpopoulos, I.H. Mott MacDonald Ltd London, formerly Imperial College London

Ho, T.Y.K. Geotechnical Engineering Office, Civil Engineering and Development Department, Hong Kong,

formerly Imperial College London

Jardine, R.J. Imperial College London

Anh-Minh, N. Atkins Geotechnics, Atkins Global, formerly Imperial College London

Keywords: ring-shear apparatus; granular media; steel and concrete interface; shaft friction; driven pile design

Modern methods for estimating the axial capacity of piles driven in granular media rely on accurate interface

shear failure models. While earlier studies have focused on determining shaft friction failure parameters from

small displacement laboratory shear box experiments, large displacement ring-shear interface tests provide a

better representation of conditions adjacent to the shafts of driven piles. This paper describes systematic

studies in which granular quartzitic media, ranging from angular rock flour to sub-rounded coarse sand, were

sheared against concrete and steel interfaces in ring shear experiments that involved several metres of shear

slippage. The study included an examination of how the large displacement processes involve grain crushing

and modify the texture of the interfaces. Conclusions are drawn regarding the constant volume angle of

interface shearing resistance that may be applied in pile design, the soils’ particle size distributions and the

roughness of the interfaces tested, before and after the ring-shear tests.

1 INTRODUCTION

Field research in clays and sands with highly instrumented displacement piles has shown that (local) shaft failure is governed by the Coulomb effective stress interface friction criterion; Jardine and Bond (1989), Lehane et al (1993). New design procedures have been developed that incorporate recommendations for evaluating the local shaft resistance by applying (i) functions that predict the expected normal effective stresses from appropriate site investigation tests and (ii) site specific laboratory interface shear experiments; Jardine and Chow (1996) and Jardine et al (2005). With granular media, the constant volume friction value, δ΄cv, has been shown to be the controlling shear parameter: local shaft failure takes place during loading tests when the soil at the interface ceases to dilate.

Preliminary correlations with field tests led Jardine et al (1992) to propose that small displacement direct shear tests involving interfaces of appropriate stress levels, materials and roughnesses might be sufficient to measure the operational δ΄cv of sands and silts. However, CUR (2001) argued that pile driving produces continuous changes in both the particle size distributions of sand in contact with the pile shaft, and the surface roughness of the interface. They concluded that operational friction angles would tend to a unique value for silica sands and suggested that

steel piles may be designed safely by simply assuming an invariable δ΄cv = 29

o.

This paper summarises how the CUR (2001) hypothesis has been investigated by Anh-Minh (2005); Barmpopoulos (2006) and Ho (2007) through large displacement tests on a suite of standard granular silica test media sheared against annular interfaces made from both mild steel and concrete. Measurements on the roughness of the interfaces and the particle size distributions of the soils, before and after each test, indicated how grain crushing, surface roughness changes and displacement magnitude may affect the ultimate interface friction angles, δ΄cv.

The interface shear experiments were performed in a modified Bishop et al (1971) ring shear apparatus that allows shear displacements of potentially many metres to be imposed between the soil and the interface, replicating aspects of pile driving. A generic ‘multi-stage’ testing approach was selected to address the influences of interface material and roughness, normal stress level, shearing history and initial particle size distribution.

Ho (2007) and Ho et al (2008) discuss the influence of the interface position or the modes of shearing. They also review earlier contributions by Potyondy (1961), Kulhawy & Peterson (1979), Yoshimi & Kishida (1981), Kishida & Uesugi (1987), Paikowsky et al (1995), Subba Rao et al (1998), Dove and Frost (1999), DeJong and Frost (2002), Frost et al (2004), Lings & Dietz (2005) and Dietz

and Lings (2006). We do not include any such review here, but note that the above studies primarily involved relatively small displacement tests performed in modified direct or simple shear cells.

2 TESTING PROGRAMME

2.1 Ring shear apparatus

The Bishop apparatus, which is illustrated schematically in Figure 1, allows notionally unlimited shear displacements to be applied continuously, without having to stop and reverse the shearing movement. Any friction developed between the soil and vertical walls of the soil confining rings can be measured and accounted for in this equipment. The sets used at Imperial College have been modified to allow shearing rates to be applied that are comparable to those mobilised during pile driving.

Figure 1. Cross section of Bishop ring shear apparatus used in the study (after Bishop et al (1971)).

2.2 Testing procedure

The authors’ multi-stage ring shear tests were designed to consider the interface shear behaviour of piles driven deeply into layers of sands or silt having constant or increasing relative density. Step increases were applied to the normal stresses, with controlled phases of large deformation shearing taking place between each loading stage.

The majority of the concrete interface tests comprised three stages conducted with normal stresses of 100, 200 and 400 kPa; two single-stage tests at 800 kPa were also carried out. A larger programme of tests was conducted with the steel interfaces with stages at 100, 200, 400 and 800 kPa. Each test involved a fresh interface, and each loading stage involved three sub-stages. In the first and last sub-stages, shearing was conducted at a slow rate of 0.9 mm/min for an annular displacement interval of about 50 mm. During these intervals, the horizontal gap incorporated in the Bishop apparatus between the

upper and lower shearing stages was kept open; the recorded stresses represent the average interface contact stresses. Fast shearing (250 mm/min) was applied to develop a 2 m annular displacement in each intervening (second) sub-stage with the gap kept closed. Sample loss was minimised, but the stress measurements were affected. Changes in the frictional response associated with particle breakage and interface roughness changes were best gauged by comparing the ‘before-and-after’ slow sub-stage measurements. The key features of this testing programme are summarized in Figure 2 where indicative tests results on two three-stage tests on sand – concrete interface are illustrated.

Overall, the total lengths of shear displacements experienced in the three- and four-stage tests were around 6 m and 8 m respectively. The two single-stage tests performed on concrete interfaces developed total shear displacements of 10 m and 15 m. Some of the experiments were repeated, but with shearing being halted at different stages so that checks could be made of the progressive changes in interface roughness and particle size distributions.

Figure 2 Indicative three-stage tests results illustrating the key features of the testing programme (sand 16/30 against concrete interfaces).

2.3 Soils tested

The soils tested were (i) four graded mortar test sands produced by the David Ball Co., UK denoted 7/14, 14/25, 16/30 and 52/100; (ii) TVS (also known as Ham River Sand) from the Thames Valley in the UK, (iii) FS a fine siliceous uniform sand from Fontainebleau, France and (iv) HPF4, an industrial rock flour silica silt. The sands were all sub-rounded to sub-angular while the silt particles were angular. Particle size distributions were determined by wet sieving; the mean (D50) values are shown in Table 1.

The shear tests were carried out dry. Since the operational constant volume δ΄cv is not affected by the initial relative density in sands (Jardine et al 1992), a single nominally uniform initial density (equal to about 1600 kg/m

3) was used in all the tests.

Table 1. Properties of soil tested and types of interface used.

2.4 Particle size distributions before and after tests

In the tests run by Anh-Minh (2005) and Barmpopoulos (2006) global assessments were made of the changes in particle size distribution based on the whole sample left at the end of testing, while Ho (2007) sub-sampled material from areas of the specimen that had experienced different degrees of shear distortion. We report here only the global data from the concrete interfaces tests, which are presented in Figure 3, noting that the fine material was concentrated in shear zones formed close to the interfaces (Figure 4).

2.5 Preparation for the interfaces tested

Concrete driven piles are often cast in stainless steel moulds that provide a relatively smooth finish along their shaft. The roughnesses of specimens cut from industrial piles (provided by Centrum Pile Limited, UK) were measured to define the finish required for the annular ring shear concrete test specimens.

The concrete interfaces were made with mortar sand (sieved to remove grains > 1.2 mm), cement and water mixed in a ratio of 2:1:0.5 by weight and cast

3(a)

3(b)

3(c)

3(d)

Figure 3. Particle size distributions of the soil samples tested against the concrete interfaces before and after the tests.

against a smooth Perspex surface, taking care to remove air voids. These mortar or ‘concrete’ specimens were cured in their moulds for 1 day and then under water for not less than 27 days before testing. Unconfined compression tests showed equivalent cube strengths exceeding 37 MPa within seven days of casting.

Interface Soil Description D50 (mm)

Steel, Concrete

HPF4 crushed industrial rock flour; Angular

0.04

Steel FS uniform sand; Angular 0.21

Steel, Concrete

52/100 uniform sand; sub-rounded 0.26

Steel, Concrete

TVS river sand; sub-rounded to sub- angular

0.32

Steel, Concrete

16/30 uniform sand; sub-rounded to sub-angular

0.72

Steel, Concrete

14/25 uniform sand; sub-angular 0.90

Steel, Concrete

7/14 uniform sand; sub-angular 1.60

Figure 4. Shear zone formed close to the interface. Sand 7/14 against steel interface (Ho 2007).

To achieve similar finishes to the industrial piles, any imperfections on the concrete ring surface were rectified by gentle hand polishing with fine sand paper. As noted below, measurements were made that indicated initial (pre-test) centre line average roughness, Ra, between 14 and 15 µm for the concrete interfaces.

The steel interfaces were made from mild steel and were shot-blasted prior to each use. The resulting Ra values of about 4 to 5µm fell towards the lower range of field measurements reported by Jardine et al (1992). Earlier measurements on uncoated weathered industrial steel piles have shown Ra of around 5 to 10 µm prior to installation.

Figure 5. Sampling length on concrete interface.

2.6 Roughness measurements before and after tests

Surface roughness measurements were taken at several locations, and over appropriate shear trajectory lengths, for all interfaces, using a Rank Taylor Hobson Talysurf profilometer. A typical sampling length on concrete interface is illustrated in Figure 5. The characteristic patterns of the concrete and steel prototype profiles are shown in Figure 6.

The roughnesses of the concrete specimens are summarised in Table 2. Measurements made on interfaces after shearing are denoted S, while the average industrial pile specimen values are indicated as I and measurements made on the concrete test interface prototypes are denoted P. Figure 6. Characteristic patterns of the concrete and steel prototype profiles.

Table 2. Roughness measurements for the concrete interfaces.

3 TEST RESULTS AND DISCUSSION

3.1 Friction angle measurements for concrete and

steel interfaces

The large displacement interface shear strength data are summarised for both concrete and steel interfaces in Table 3 and Figures 7 and 8 respectively.

The δ΄cv values were displacement-dependant: the values developed after 50 mm and 6 m of shearing differed most markedly with the concrete tests on coarse sands. The data plotted represent average values determined after shear displacements of 2, 4, 6 and 8m. Figures 7(a) and 8(a) show the full normal stress range, while Figures 7(b) and 8(b) focus on results obtained with normal stresses up to 400 kPa.

The results presented in Table 3 correlate closely with the sands’ initial D50 values and to clarify the results, linear regression analyses were made of all the measurements made considering three sets of relatively narrow initial D50 ranges, which are annotated as 1, 2 and 3 on the Figures. The relatively high R

2 regression coefficients shown in Figures 7

and 8 confirm the strong linear correlation between

Ra (µm)

D50 0.04 0.26 0.32 0.72 0.90 1.6 1.6* 1.6#

S 12.4 29.6 23.0 40.8 45.6 55.6 93.4 118

P 14.2

I 12.7

I: Industrial, P: Prototype, S: Sheared profile, D50: Initial size

*10m of shearing, #15m of shearing

10mm

normal and shear stresses (with c΄ = 0), indicating that the large displacement values of δ΄cv did not depend significantly on normal stress level. Note that the results of the two large displacement single-stage concrete interface tests have been excluded from the Group 3 regression analysis summarised in Figure 7(a). Table 3. Concrete interfaces, defined after shear displacements of 4 to 8m unless otherwise stated.

The concrete interface tests indicate average ‘6_m’

displacement δ΄cv values for Groups 1, 2 and 3 of 25.6

o, 27.3° and 30.9

° respectively, showing trends

for the ‘6_m’ displacement δ΄cv values to be higher for the finer soils and for the sand-concrete tests to develop slightly lower δ΄cv values than the sand-steel cases. The trend for tan δ΄cv against initial mean particle size D50, is shown in Figure 9. Also shown, is the spread of design recommendations noted by Jardine et al (2005); the concrete ‘6_m’ trend falls between the CUR (2001) recommendation and the variable trend seen in the short displacement direct shear tests reported by Jardine et al (1992).

The steel interface tests showed a weaker

dependency of tan δ΄cv on initial D50. The angles for

the Group 1 and 2 sand samples fell within a

relatively narrow range (27.5° to 29.0°) although the

angular silt samples gave angles about 2° higher.

Generally, the results for the steel interface tests are

slightly higher than those for concrete and are 0° to

1.5o lower than the CUR (2001) recommendation of

δ΄cv = 29°. The differences between the small and

large displacement shear test trends reflect the

gradual growth with displacement of particle

breakage and, for the steel interfaces, surface

smoothing.

3.2 Particle size distributions and interface surface

roughnesses for the concrete tests

It is immediately clear that the coarser sands (e.g. Sands 7/14, 14/25 and 16/30), experienced the most substantial degrees of soil grain crushing and fines generation during shearing.

(a) Range up to 800 kPa (full scale)

(b) Range up to 400 kPa (re-scaled for clarity) Figure 7. Results on concrete interfaces

(a) Range up to 800 kPa (full scale)

(b) Range up to 400 kPa (re-scaled for clarity)

Figure 8. Results on steel interfaces

Friction angles δ’cv for the concrete and the steel interfaces

Sand

Tested HP

F4

FS

52

/100

TV

S

16

/30

14

/25

7/1

4

Concrete

after 6m 31.0 n/a 27.6 26.9 25.5 25.7

25.9

27.0*

29.5#

Steel

after 8m 30.5 27.5 28.0 28.6 28.7 27.4 29.0

Note: Above values obtained from linear regression analysis

* δ΄cv after 10m, # δ΄cv after 15m

Figure 9. Comparison of the results of interface friction angle δ΄cv for the concrete and steel interfaces with various published trend lines (modified after Jardine et al (2005))

Figure 3(a) considers data from identical tests that were halted after different stages, showing how the relative volume of the fine soil (developed near the interface) grows sharply over the first few metres of displacement and stabilises after around 10 to 15m. The finer sands (e.g. TVS and Sand 52/100) showed far less significant crushing and variations in their global particle size distributions.

Regarding the roughness of the interfaces (Table 2), in all tests, except those on HPF4 silt (D50 = 0.04 mm), the concrete interfaces became rougher after shearing against the sands under a final normal stress of 400 kPa. It is possible that industrial piles would have harder surfaces and might experience less marked increases during driving. As with particle crushing, surface roughening increased with total shear displacement. Shearing against HPF4 silt reduced the surface roughness, suggesting a polishing action against the concrete interface.

The observation of marginally higher δ΄cv values in the concrete-sand tests than in the steel-sand experiments may appear to contradict the classical expectation that interface friction should increase as surface hardness falls and surface roughness increases: see for example Bowden and Tabor (1967), Frost et al (2002). However, the anomaly may relate to the different surface morphology of the two interface types. The concrete interface surface can be described as being undulated with gentle peaks and valleys (Figure 6) that may encourage the coarse grain soils (i.e 7/14, 14/25, 16/30) towards a slipping failure mode, resulting in lower δ΄ values. With increasing displacement, crushing of particles and ploughing of the interface occurs and fines concentrate at the interface resulting in the trend for δ΄cv to rise as displacements increase.

In contrast, the steel interface has both sharp peaks and deep valleys (Figure 6). The grains of both coarse and fine soils are in contact with numerous sharp asperities. The greater hardness of the steel leads to less surface abrasion taking place; the

generation of fines by sand particle crushing dominates the trend for δ΄cv to grow with post-peak shear displacement. Ho et al (2008) provide further information regarding the particle size distribution and roughness trends for the steel interfaces; the latter generally became marginally smoother as a result of shearing.

3.3 Effect of particle crushing and relative

roughness on δ΄cv for the concrete interfaces

As noted above, prolonged shearing under high stresses modifies the particle size distribution (see Figure 3) and interface surface roughness (see Table 2). The effect of the accumulative displacement on tan δ΄cv is further illustrated for concrete interfaces in Figure 10. The δ΄cv values of the coarse sand increases progressively with shear displacement. Figure 10. Effect of shearing displacement on δ’cv for the tests on concrete interfaces

Data from short displacement interface shear tests

can show practically linear relationships between tan

δ΄ and normalised relative surface roughness Ra/D50;

see for example Uesugi and Kishida (1986) or

Jardine et al (1992). Considering first the δ΄cv angles

defined after just 50mm of displacement, when fines

generation would be minimal and the interfaces

relatively fresh, the soil-concrete data may be plotted

meaningfully in terms of the initial Ra values and

particle size distributions, as summarised by the D50

data. The 50mm results have been processed in this

way and the resulting tan δ΄cv trends given at the left

hand side of Figure 11 show good agreement with

the direct shear tests of Jardine et al (1992). Also

shown are trends from the large displacement

concrete interface shear tests, where Ra is taken from

the post-test interface shear roughness measurements

and D50 is estimated to fall in the range of the fine

soil isolated in the grading tests, as this dominates in

the shear zone. Ho et al (2008) give a closer analysis

of the equivalent trends for the steel interface tests.

The results presented in Figure 11 suggest a

general overall pattern for the sub-rounded to sub-

angular sands shearing against the concrete

interfaces. It appears that tan δ΄cv increases linearly

with relative roughness (from a minimum of around

0.425) until Ra/D50 ~ 60×10-3

when it reaches a

plateau with tan δ΄cv ~ 0.51 (or δ΄cv ~ 27o) that

applies, with some variation, to roughness ratios up

to 1200×10-3

. The angular silt is able to develop

marginally higher δ΄cv values, but none of the

interface shear tests is able to reach angles as high as

those normally associated with triaxial or direct shear

tests on the same soils under critical state shearing

conditions. The sub-rounded to sub-angular silica

sands tested typically show φ΄cv values of around 32o;

Jardine et al (1992).

Figure 11. Relationship tan δ΄cv for the concrete interfaces and normalised roughness, with trend lines

4 SUMMARY AND CONCLUSIONS

The experiments described above lead to six main conclusions: 1. Silica sands sheared against concrete interfaces of

similar initial roughness to industrial piles developed large-displacement constant volume interface friction angles δ΄cv of 25.5° to 31° that increased with decreasing initial grain size and generally fell below the sands’ φ΄cv values (the latter typically falling around 32

o).

2. The concrete interface δ΄cv values were also slightly (0

o to 1.5

o) lower that those applying in

equivalent tests on mild steel interfaces that had similar roughnesses to industrial piles.

3. Differences in the interfaces’ surface topographies might explain the otherwise surprising trend for δ΄cv to be higher in the tests involving hard steel interfaces than in those on softer concrete surfaces.

4. The steel interface tests showed lower sensitivity to initial particle size, their mean δ΄cv values fell marginally below the 29

o design recommendation

proposed by CUR (2001).

5. Tests on both interface types showed δ΄cv values that did not depend significantly on the normal stresses, over the ranges applied.

6. Large shear displacements permitted significant particle breakage to occur, causing considerable changes in surface roughness and increases in δ΄cv.

7. The studies confirm that δ΄cv is displacement-dependent. This has many important implications for practical geotechnical engineering. Due consideration must be given to the choice of test method when determining interface shear parameters for different applications. For example, limited displacement direct or simple shear test data are not applicable to the analysis of displacement pile foundations, for which ring shear testing is more appropriate. However, small displacement test data are more appropriate in applications such as shallow foundations, cast shafts or retaining walls.

ACKNOWLEDGEMENTS

The Authors acknowledge the generous assistance of

colleagues and technical staff at Imperial College

London. This paper is also published with the

permission of the Head of the Geotechnical

Engineering Office and the Director of Civil

Engineering and Development, the Government of

the Hong Kong Special Administrative Region, who

supported the second Author during his studies at

Imperial College.

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