analysis of interface friction effects on microtunnel jacking forces in coarse-grained soils
TRANSCRIPT
ABSTRACT: Microtunnelling is an important trenchless construction technique that is used to successfully install essential utility pipelines in increasingly congested urban centres around the world. An important consideration for a microtunnelling project is the magnitude of the jacking force that will be required to advance the microtunnelling shield and the string of product pipes from the starting shaft to the receiving shaft. Frictional resistances along the surface of the pipeline have a major contribution to the total jacking force. This paper considers the frictional resistance mechanism involved in advancing concrete pipes through a coarse-grained soil and describes laboratory testing carried out with the aim of physically modelling the process. Comparisons are made with case histories from microtunnelling projects recently completed in coarse-grained soils. Recommendations are made on predicting likely jacking forces in advance of future projects.
KEY WORDS: Microtunnelling; Pipe-jacking; Interface friction; Coarse-grained soils.
1 INTRODUCTION The use of microtunnelling as a cost effective, safe and environmentally sound means of providing utility pipelines is slowly being accepted in Ireland [1], while the technologies have been in extensive use elsewhere in the world for a great many years [2]. The prediction of jacking forces during a microtunnelling project is important for reasons of economy and to establish confidence in the proposed method of works, but these predictions are quite difficult due to the complex soil-structure interactions present [3]. The ground resistances, which the jacking force must overcome, are a combination of face resistance at the front of the shield and frictional resistance along the surface of moving pipeline, which increases as more pipes are added. The frictional resistance tends to be much greater in magnitude than the shield face resistance. It is more readily related to quantifiable pipe-soil interactions than the face resistance, which can be heavily operator-dependent [4]. As lengths demanded of microtunnels get longer [5, 6], and curved drives become more common [7], the mechanism of frictional resistance is an obvious area for further study with the objective of proposing suitable methods to predict and reduce it. Several factors effect the generation of skin frictional resistances during a pipe jacking operation [8, 9]:
• Type of soil and variation along the pipeline • Normal stresses acting on pipeline • Roughness of pipeline surface • Contact conditions between pipe and soil • Overcut during excavation • Position of water table • Type and consistency of lubricant (if any) • Duration of stoppages during driving • Pipeline misalignment and/or curvature.
All of these factors are in some way related to the properties of a thin zone of intensely shearing material in the contact zone between a jacked pipe and the soil. The soil in this zone will be extensively reworked and remoulded as the pipeline advances at rates upwards of 30mm/minute. Much research has been undertaken on the shearing mechanism between soil and construction materials [10-15]. The main factors shown to affect interface friction behaviour under direct loading are interface roughness, soil density, mean particle size, particle angularity and normal stress across the interface. Uesegi et al. [12] performed simple shear tests between air-dry sand and concrete which showed that the peak coefficient of friction depends on the maximum peak-to-valley surface roughness (Rmax) of the concrete and the mean particle size (D50) of the sand. It was found that as the normalised roughness increases, a critical roughness is reached after which the coefficient of friction ssumes a peak value, as shown in Figure 1.
Many studies have attempted to model the behaviour of the soil being sheared between a jacked pipe and the soil. Zhou [16] carried out numerical modelling using interface elements and found that in coarse soils the pipe separates or almost separates from the surrounding soil over a large proportion of the upper pipe surface area, and that normal stresses between the pipe and the soil over this part are low. Phelipot et al. [17] used a device similar to an annular shear apparatus to meaure the effects of overcut and lubrication on the advance of a model micro tunnel boring machine and pipe string. Iscimen [18] and Staheli [4] describe tests carried out using a novel shear testing apparatus capable of accepting curved sections of pipe. Shou et al. [19] use a simple apparatus to quantify the impact of different lubricant mixes on the jacking forces in sandy laterite gravel. In all the studies referred to above, large variations were reported between the experimental modelling results and field measurements. A
Analysis of interface friction effects on microtunnel jacking forces in coarse-grained soils
Reilly, Ciaran C.1and Orr, Trevor L.L.1
1Department of Civil, Structural and Environmental Engineering, University of Dublin, Trinity College, Dublin 2, Ireland email: [email protected], [email protected]
common uncertainty is the contact area between the pipeline and the soil.
Figure 1 - Coefficient of friction reaches a peak value at critical normailsed roughness of steel, Rn = Rmax/D50 [11]
This paper gives preliminary results from physical modelling carried out using coarse-grained soils in order to better understand the pipe-soil interface present in pipe jacking. These results are compared to the frictional stresses observed during a number of pipe jacking operations, and an attempt is made to relate measured field values to the results of physical modelling.
2 JACKING FORCE PREDICTION In microtunnelling, a force Ptotal needs to be provided to advance the microtunnelling shield and the string of product pipes behind it through the ground. Ptotal needs to overcome the combination of the face resistance at the front of the shield, PF and the skin friction resistance along the pipeline, PS:
!!"!#$ = !! + !! (1)
Face resistance 2.1Face resistance has been found to be dependent on operator influence, and as such is hard to predict. The face resistance must be maintained higher than the active earth pressure and lower than the passive pressure in order to prevent settlement or heave on the surface, and must be adjusted on site to suit the conditions encountered [20].
Frictional resistance 2.2Depending on the stability of the soil during pipe jacking, the pipeline may slide along in an open bore, or the soil may collapse onto the pipe barrel. These two cases will produce markedly different frictional resistances, although even when the soil has collapsed onto the pipe, arching in the ground above the collapsed soil will reduce the normal stress acting on the pipe barrel below the initial in-situ value [21].
Figure 2 - Stable and unstable bore situations
As the effective angle of friction, ϕ’, defines the shearing resistance of a soil, the angle of interface resistance, δ, defines the shearing resistance of an interface. The interface friction angle is usually less than the effectice angle of friction. For convenience, an interface friction coefficient, µ, is defined:
! = tan ! (2) Eventual collapse of an excavated face is inevitable in a coarse-grained soil due to lack of undrained shear strength, although there is some evidence that suction effects in partially saturated coarse-grained soils may allow for some stand-up time, and short term stability. Stability can also occur if correct lubricant pressure is maintained during pipe jacking. The case histories in Section 4 will elaborate on this.
2.2.1 Stable ground For coarse-grained soils, the frictional resistance acting on a pipe string sliding in a stable bore is a function of the weight of the pipes and the frictional properties of the pipe-soil interface:
!! = ! ∙! (3)
where W is the weight of the pipe string per linear metre and L is the length of drive. If the bore is located in ground water, the bouyant weight of the pipeline should be used:
!! = ! ∙ ! − !!!!!!
! (4)
where γw is the unit weight of water and DP is the outside diameter of the pipeline, shown in Figure 2.
2.2.2 Unstable ground The frictional resistance acting on a jacked pipeline in unstable ground is a function of the normal force acting on the pipeline, N, and the coefficient of friction, µ:
!! = ! ∙ ! (5)
The normal force is the integral of normal stress, σN, over the pipeline surface. The normal stress, σN, is the stress due to soil and groundwater loading acting perpendicular to the pipeline surface. It has been shown that there is good agreement between the vertical stress acting at the pipe crown level, σEV, and σN [4], due to the effects of arching. This allows for
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estimation of the normal force directly from arching theory, as follows [22], with reference to Figure 3:
!!" =
!"!! ∙ !"#!
1 − !!! ∙ !"#! ∙ !/! (6) where K is the coefficient of lateral earth pressure above the microtunnel (taken as 1 following recommendations in the literature, [3]), H is the height of cover over the roof of the pipe and B is the width of affected ground, defined as follows:
! = !! 1 + 2 tan
!!− !
! (7)
Figure 3 - Silo pipe loading model after Terzaghi [25]
3 TESTING Tests using a standard small shear box were carried out at TCD’s Geotechnical Laboratory to establish the interface friction characteristics of two types of sand against a rough concrete surface, approximating the surface of a jacking pipe.
3.1.1 Soil properties Two different fine to medium silica sands were tested – Glenview sand and Irish Glass Bottling (IGB) sand. Glenview sand is manufactured by crushing sandstone while IGB sand is a Belgian silica sand that was obtained from the Irish Glass Bottle plant in Ringsend, Dublin. The properties of each sand are listed in Table 1. Both sands are uniformly graded. The biggest difference observed between the sands is the shape of particles; this difference is shown clearly in Figure 4. Soils were tested dry. A loose density was achieved by air pluviating the sands into the shear box. A loose density was considered appropriate to model the effects of the overcut, where the TBM excavates a space with diameter DS which is larger than the diameter of the pipes, DP.
Table 1 - Soil properties
Sand Glenview IGB D10 0.08mm 0.164mm D50 0.2mm 0.23mm CU 2.85 1.60 CZ 1.23 0.885 γd,max 16.0 kN/m3 15.7 kN/m3 γd,min 13.8 kN/m3 14.2 kN/m3
Shape Angular to very angular
Subrounded to subangular
Figure 4 - IGB sand (top) & Glenview sand (bottom)
3.1.2 Concrete
Concrete jacking pipes used in Ireland are typically manufactured using the dry cast concrete process, where the pipes are formed by mechanically placing low-slump concrete between a vibrating core former and a stationary outer form. The resulting pipes are strong and impermeable. The outer finish is generally quite rough, as illustrated in Figure 5 (a) & (b) below, but can vary to quite smooth in some cases due to variations in the manufacturing process (Figure 5 (c)). To attempt to model the shearing behaviour of coarse-grained soils against concrete jacing pipes, a concrete sample was manufactured. Its dimensions were 60mm x 60mm x 20mm and a strong concrete mix was used. The main considerations in making the slab were durability during repeated shearing and achieving a surface finish close to what would be expected for jacking pipes.
(a) Rough finish
(b) Medium finish
(c) Smooth finish
Figure 5 (a) to (c) - Typical dry cast concrete jacking pipe
Figure 6 - Concrete test surface
Testing 3.2Tests were conducted on a small shear box apparatus on dry samples of each type of sand. Sand was pluviated to unit weights as specified in Table 2. A shearing rate of 1.2mm/min was chosen for convenience. Other work on sand/concrete interface testing [23] showed that while both δ and ϕ increase with shearing rate, the ratio δ/ϕ remains the same. Sufficient travel was allowed for the sand to reach critical state values of ϕ and δ, ϕc and δc. Results of interface friction testing are set out in Table 3. Figure 7 shows plots of shear stress against horizontal displacement for some of the tests on IGB sand. It is notable that the shear stress plot isn’t smooth, as would be expected, but quite rough. This may be because of the roughness of the interface. While δc/ϕc decreases with normal stress for Glenview sand, it remains constant for IGB sand.
Table 2 - Testing conditions for each sand
Sand Glenview IGB γd,test 13.8 kN/m3 14.6 kN/m3 ϕc 36.5º 33º
Table 3 - Results of interface friction tests between rough concrete and two types of sand
Glenview sand IGB sand σN
(kPa) 18.6 25 50 100 18.6 25 50
τ (kPa) 18.7 20.2 38.8 68.8 12.4 16.5 33
δc 45º 39º 38º 35º 34º 33º 33º µc 1 0.81 0.78 0.69 0.67 0.66 0.66 δc/ϕc 1.23 1.07 1.04 0.96 1.03 1.0 1.0
Figure 7 - Testing results for IGB sand sheared against the
concrete test surface in direct shear
60 1 2 3 4 5
35
0
5
10
15
20
25
30
Horizontal displacement (mm)
Shea
r str
ess ! (
kPa)
IGB sand
Loose
σn=18.6kPa
σn=25kPa
12.4kPa
16.5kPa
σn=50kPa 33kPa
4 CASE HISTORIES The findings from the authors’ laboratory tests were compared to measurements made during two microtunnelling schemes in sandy soil; one in Ireland and one in France reported as part of the French national project “Microtunnels” [3].
4.1.1 Case History 1 In Howth, Co. Dublin, three 1000mm internal diameter microtunnel drives of 60m to 110m in length were carried out in slightly silty estuarine sand and sandy gravel as part of a wastewater network improvement scheme [24]. The effective angle of friction from shear box testing was in the range 37º to 42.5º. Average saturated unit weight was 21.2kN/m3. The tunnelling shield used was a Herrennecht AVN 1000 with diameter, DS, of 1295mm. The pipes had an outer diameter, DP, of 1200mm and a cover depth of 2.5m to 3.9m at the start of each drive. A 10m section at the start of each drive was unlubricated for operational reasons, while the remainder of each drive was lubricated with an average of 60 litres/linear metre of bentonite lubricant. Table 4 shows the relevant parameters for both the initial unlubricated section and the remainder of the drive which was lubricated, along with the reduction in resistances observed due to the effects of lubrication.
Table 4 - Frictional stress measured & predicted for Howth
Unlubricated Lubricated Reduction Normal stress σN (kPa) 16.5 – 19.1 16.5 – 19.1 No change
Measured friction (kPa) 2.3 – 6.0 1.4 – 2.1 65 – 185%
Calculated friction (PS = µN) 8.1 – 24 5.6 – 8.7 31 – 64%
Calculated friction (PS = µW) 0.11 – 0.35 0.08 – 0.12 33%
Calculated δ 6.8º – 19.8º 4.6º – 7.4º 32 – 62% Calculated µ 0.12 – 0.36 0.08 – 0.13 33 – 62% Calculated δ/ϕ 0.16 – 0.46 0.11 – 0.17 31 – 63%
Figure 8 - Unlubricated section of Howth drive 1 in sand,
showing predicted and measured total jacking forces
“Calculated” indicates values back calculated from site records and based on the assumption of uniform all-round contact between the soil and the pipe. This approach is useful and common in practice [20]. Figure 8 outlines how the measured jacking forces compare to both the stable bore (F=µW) and unstable bore (F=µW) models discussed earlier, using a value for µ obtained for a fairly similar soil through laboratory testing at the appropriate density and normal stress. It is noted that measured total jacking force lies between both sets of predictions, indicating that neither entirely stable nor unstable bore conditions prevailed. A possible explanation for this is that a combination of the large overcut created (42.5mm on the radius) and the slight siltiness of the soil allowed the overcut to remain partly open during pipe jacking.
4.1.2 Case History 2 Three drives were constructed in clean sand in Bordeaux, France, using a Markham 500 microtunnel boring machine [3]. The sand had a unit weight of 18 to 20 kN/m3 and an effective angle of friction was in the range 30 to 35˚. The outside diameter of the pipeline was 650mm, and a 10mm overcut was created. The cover depth was 7m and the average length of drive was 90m. Bentonite lubricant was used after 16m, and applied at a rate of 168 litres/linear metre. Table 5 shows the relevant data for this project, together with the reduction in resistances caused by lubrication.
Table 5 - Frictional stress measured & predicted for Bordeaux
Unlubricated Lubricated Reduction Normal stress σN (kPa) 18 18 No
change Measured friction (kPa) * 4.7 0.5* – 2.5 89%
Calculated friction (PS = µN) 5 – 5.4 2.7 – 2.9 86%
Calculated friction (PS = µW) 0.4 – 0.5 0.4 – 0.5* No
change Calculated δ 15.1 – 18.8 5.7 – 8.0 57 – 62% Calculated µ 0.27 – 0.34 0.1 – 0.14 50 – 70% Calculated δ/ϕ 0.43 – 0.63 0.16 – 0.27 57 – 63% * At some locations, average measured frictional stress was of the same magnitude as calculated self-weight friction stress.
There are similarities with the results obtained from Case History 1, in that values for the coefficient of friction µ for sliding contact between the soil and the pipe would not accurately predict the observed jacking forces. A back calculated coefficient of friction, µi, gives a reasonable prediction of observed values. It is notable that at times, a stable bore situation was achieved, indicated by * in Table 5 above. This, along with the consistent and large percentage reductions in frictional resistance force PS between unlubricated and lubricated driving, indicates that the lubrication regime was effective here.
5 DISCUSSION The interface shear tests performed produced results largely in line with literature. Potyondy [10] observed values of μ up to 0.97 in tests with dry sand and rough concrete. μ was
100 1 2 3 4 5 6 7 8 9
600
050
100150200250300350400450500550
Drive length (m)
Tota
l jac
king
forc
e (k
N)
Measured
μ=0.9 (testing)σ =19.1kPa (calculated)
Unlubricated
Predicted Ps=μW
Predic
ted Ps
=μN
N
observed to decrease as the normal force increased, in common with other works [10, 18]. It is notable that the grain shape of the sands leads to differences in the density achieved through pluviation into the shear box apparatus and differences in shearing behaviour observed, with the rounder sand achieving a lower shear strength but higher density, most likely as the particles can more easily and smoothly translate with respect to each other.
Neither the fully stable, nor the fully unstable models for pipe-soil contact in pipe jacking are found to be appropriate for the case histories presented, at least not using interface friction coefficients established from laboratory tests. If laboratory interface friction coefficients, ranging from 0.7 to 1.0, are utilised, the stable bore model underestimates the jacking forces, while the unstable bore model overestimates them. If an interpreted interface friction coefficient μi is utilized, ranging from 0.12 to 0.36, good agreement can be found. Clearly, there is a difference in the mechanism causing the interface frictional behaviour between small scale tests in the laboratory and full scale pipe jacking in the field. Field results analysed by Norris [25] showed highly localised and irregular contact between the pipe and the soil. In silty sand, soil was found to collapse onto the top of the pipe, while horizontal stresses varied from low to high. This is an area which requires further research.
6 CONCLUSION Physical modelling of soil-pipe interaction during pipe jacking and comparison to field results has raised some interesting questions.
1. It is clear that the interface friction during pipe jacking is influenced by parameters common to other soil-structure interface friction situations, including particle shape and size, soil density and the state of stress normal to the interface surface.
2. Due to differences in the magnitude of forces observed between physical modelling and the case histories, it is concluded that uniform contact conditions do not occur around a jacking pipe.
3. Possible reasons for this include short-term stability provided by suction in partially saturated soils, pipe misalignment and stoppages during jacking. These are areas that require further research.
4. Two methods have been examined which facilitate the prediction of jacking forces during pipe jacking. Neither method has been found reliable when combined with interface friction parameters derived from physical modelling.
5. Empirically derived interface friction parameters may be used to predict likely jacking forces with reasonable reliability, once good judgement is exercised.
ACKNOWLEDGMENTS The support of the Irish Research Council is gratefully acknowledged. Terra Solutions Ltd. are acknowledged for providing access to jacking force records for Case History 1.
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