Review:SupportDesigninSqueezingGround
byErikEberhardt,Mar.2017
In weak, tectonically disturbed rock under high in‐situ stresses, a ductile rock mass response can be expected resulting in large plastic shear deformations and squeezing ground conditions. Similar rock mass conditions have been recently reported by Hoek & Guevara (2009) for the Yacambú‐
Quibor tunnel in Venezuela. This 5.2 m diameter water conveyance tunnel passes through highly variable, tectonically disturbed metamorphic rocks (e.g. graphitic phyllites) and a large regional fault related to the intersection of three major crustal plates. The 25 km long tunnel involves overburdens of more than 1000 m. Based on their experiences on the Yacambú‐Quibor tunnel, for which squeezing ground conditions have led to a long history of ground control problems and construction challenges, Hoek & Guevara (2009) outline a methodology for the design of tunnel linings under extreme squeezing conditions. These include several points of consideration: • The key to liner design is to sequence the installation of support to avoid overload while still maintaining a safe working environment at and near the tunnel face. • As the rock mass is already in a tectonically deformed (residual) state, it can be assumed to act plastically in response to stress change. • The long‐term strength of the rock mass is assumed to correspond to moderate disturbance according to the GSI system with a Disturbance factor D = 0.2. • The deformation modulus of the rock mass can be estimated based on the methodology of Hoek & Diederichs (2006). The analysis recommended by Hoek & Guevara (2009), and also discussed in Hoek et al. (2008), for estimating the support capacity of the lining installed at Yacambú‐Quibor involved the following steps: 1) From the available geological data, estimate the characteristics of the rock mass along the tunnel. 2) On the basis of estimates of the Geological Strength Index (GSI) and intact rock strength (ci), estimate the strength of the rock mass. 3) Using the ratio of rock mass strength to in‐situ stress (calculated from the height of overburden), estimate the potential for overstressing of the tunnel (Hoek & Marinos 2000). 4) For those sections where the tunnel is highly stressed, estimate the support capacity of the installed tunnel lining by numerical analysis. 5) Where the support capacity is found to be inadequate, investigate the measures that are required in order to increase the capacity of the lining to an acceptable level. The procedure recommended for analyzing the tunnel support capacity is provided by Hoek et al. (2008) and Hoek & Guevara (2009) and involves using the Rocscience finite element program Phase2. This program is available for use in the Earth & Ocean Sciences student computer labs. The model input parameters can be derived using the GSI and Hoek‐Brown systems estimates for rock mass properties, as described the Part 1 Supplementary Notes, with the long‐term residual strength calculated using a Disturbance Factor of 0.2 as recommended by Hoek et al. (2008). The Phase2 finite‐element models can first used to calculate the characteristic curves for the rock mass. This is done by setting up the model geometry to represent the tunnel without support, specifying an internal support pressure Pi inside the tunnel equal to the far‐field in‐situ stress Po, and progressively reducing the internal support pressure to produce a plot of tunnel wall displacement versus support pressure. To design the appropriate timing for the installation of the support or when optimizing the installation of the support with specific displacement capacity, the longitudinal closure profile for the tunnel must first be determined so that the amount of closure at the time of support installation can be assessed (Hoek et al. 2008; Hoek & Guevara (2009)). In other words, a portion of the maximum tunnel closure will begin to take place once the opening is drilled, blasted and mucked out. The tunnel boundary will then continue to displace inwards as the tunnel is advanced, eventually reaching its maximum value. Although analytical solutions for calculating the longitudinal displacement profile exist, for example that by Chern et al. (1998), Hoek et al. (2008) caution that these may not be adequate if the radius of the plastic zone around the tunnel exceeds two tunnel radii and/or if the tunnel face interacts with the developing yield zone around the tunnel. Chern et al.'s (1998) empirical fit is given by: ⁄
.
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(14) where ur and umax are the radial and maximum radial closure, respectively, x is the longitudinal position from the tunnel face, and a is the tunnel radius. Alternatively, Hoek et al. (2009) suggest that the longitudinal displacement profile can be calculated using a 2‐D axisymmetric finite element model, for example using Phase2, assuming uniform or isotropic initial stress conditions and a circular tunnel cross section. Where complex loading and geometric conditions (i.e. a multi‐benched cut) are involved, 3‐D numerical modelling is required. From the calculated longitudinal displacement profiles, the amount of tunnel closure based on the distance from the tunnel face at which the support is installed can be estimated. In general, it is assumed that one tunnel advance cycle (drill, blast and muck out) is one half tunnel diameter. This distance can be modified according to the specific blast round length and/or logistical requirements for support installation if required. From the wall displacement at the time of support installation, calculated from the longitudinal displacement profiles, the support pressure pi can be estimated from the ground characteristic curves calculated earlier. It should be noted that if the values calculated indicate complete closure before support installation is possible, then a benched excavation to reduce the size of the unsupported tunnel section together with forepoling to support the tunnel face may be required to significantly reduce and manage the amount of closure preceding installation of the support ring. Otherwise, the final step in the procedure by Hoek et al. (2008) is to use the estimation of internal support pressure from the ground characteristic curves and determine the contribution of each of the support elements (e.g. where steel sets are embedded in shotcrete), to be able to adjust the number and dimensions of each component to accommodate the loads imposed on the lining. Consideration must be given to how these are treated in a numerical analysis where beam elements, constituting tunnel support, are attached to the tunnel boundary and the axial thrust, bending moments and shear forces induced in these elements are computed directly. These differ from rockbolts, which act as tunnel reinforcement in that they change the mechanical properties of the rock mass surrounding the tunnel. Thus, in the numerical analysis, the loads imposed on the lining will be reduced by the reinforcement and the composite lining will respond to these reduced loads (Hoek et al. 2008). [Recommendation: For this assignment, it is suggested that the support package involves rockbolt reinforcement and a composite wire mesh/fibre‐reinforced shotcrete liner. The reinforcement is assumed to involve pattern bolting of 4‐m long super Swellex bolts. The composite liner is assumed to involve welded wire mesh #6 and 150 mm of steel‐reinforced shotcrete.] Using Hoek et al.'s (2008) numerical procedure, capacity plots can be generated that indicate the various Factor's of Safety for thrust and bending moments from when the support is installed to when the final distribution of stress and displacement are achieved. The value of this analysis is that the results can provide important insights into the complex short‐ and long‐term liner loads placed on the various individual components of the support package. References
Chern, J.C., Shiao, F.Y. & Yu, C.W. (1998). An empirical safety criterion for tunnel construction. In Proceedings of the Regional Symposium on Sedimentary Rock Engineering, Taipei. pp. 222–227. Hoek, E., Carranza‐Torres, C., Diederichs, M.S. & Corkum, B. (2008). The 2008 Kersten Lecture: Integration of geotechnical and structural design in tunnelling. In Proceedings, University of
Minnesota 56th Annual Geotechnical Engineering Conference, Minneapolis. University of Minnesota, pp. 1‐53. Hoek, E. & Diederichs, M.S. (2006). Empirical estimation of rock mass modulus. International Journal of Rock Mechanics & Mining Sciences 43(2): 203‐215. Hoek, E. & Guevara, R. (2009). Overcoming squeezing in the Yacambu´‐Quibor tunnel, Venezuela. Rock Mechanics and Rock Engineering 42: 389‐418. Hoek, E. & Marinos, P. (2000). Predicting tunnel squeezing problems in weak heterogeneous rock masses. Tunnels and Tunnelling International 32(11): 45‐51.