Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
1. PURPOSE 
The purpose of this calculation is to analyze the stability of repository non-emplacement drifts 
during the preclosure period, and to provide a final ground support method for non-emplacement 
drifts for the License Application (LA). This calculation will provide input for the development 
of LA documents. 
The scope of this calculation is limited to the non-emplacement drifts including access mains, 
ramps, exhaust mains, turnouts, intersections between access mains and turnouts, and 
intersections between exhaust mains and emplacement drifts, portals, TBM launch chambers, 
observation drift and test alcove in the performance confirmation (PC) facilities, etc. The 
calculation is limited to the non-emplacement drifts subjected to a combined loading of in-situ 
stress, seismic stress, and/or thermal stress. Other effects such as hydrological and chemical 
effects are not considered in this analysis. 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
2. QUALITY ASSURANCE 
This calculation is developed in accordance with AP-3.12Q, Design Calculations and Analyses. 
In accordance with the Q-list (BSC 2003d, p. A-7), the ground control system for non-
emplacement drifts is designated as not important to waste isolation, and not important to 
safety, and the Safety Category (SC) is Non-SC. Therefore, this calculation is not subjected to 
the requirements of the Quality Assurance Requirements and Description (QARD) (DOE 2003). 

Title: Ground Control for Non-Emplacement Drifts for LA 19 
DI: 800-KMC-SSD0-00700-000-00A 
3. COMPUTER SOFTWARE AND MODEL USAGE 
Three commercially available computer programs, FLAC (Fast Lagrangian Analysis of 
Continua), FLAC3D (Fast Lagrangian Analysis of Continua in 3 Dimensions), and 3DEC (3Dimensional 
Distinct Element Code) are used in this calculation. Descriptions of these codes 
and their qualification status are provided in the following subsections. 
3.1 FLAC COMPUTER SOFTWARE 
FLAC Version 4.0 (STN: 10167-4.0-00) is a two-dimensional explicit finite difference code 
which simulates the behavior of structures built of soil, rock, or other materials subjected to 
static, dynamic, and thermally-induced loads (Itasca Consulting Group 2002). Modeled 
materials respond to applied forces or boundary restraints according to prescribed linear or nonlinear 
stress/strain laws and undergo plastic flow when a limiting yield condition is reached. 
FLAC is based upon a Lagrangian scheme, which is well suited for large deflections and has 
been used primarily for analysis and design in mine engineering and underground construction. 
The explicit time-marching solution of the full equations of motion, including inertial terms, 
permits the analysis of progressive failure and collapse. A detailed discussion on the general 
features and fields of the FLAC computer software applications is presented in the User's 
Manual (Itasca Consulting Group 2002). 
FLAC was used in mechanical analyses in this calculation. All input and output files for this 
software item used in this calculation have been archived on CD-ROMs and will be submitted to 
the Record Processing Center (RPC) as part of the records package for this calculation. The 
results are presented and described in Section 6. 
FLAC Version 4.0 was obtained from the Software Configuration Management (SCM) in 
accordance with the AP-SI.1Q procedure. FLAC is installed and run on stand-alone PCs with 
Windows 2000 operating systems. FLAC Version 4.0 was qualified for use in design in 
accordance with the AP-SI.1Q procedure. FLAC was appropriate for the applications used in 
this analysis, and used within the range of validation, as specified in the software qualification 
documentation. 
3.2 FLAC3D COMPUTER SOFTWARE 
The FLAC3D Version 2.1 (STN: 10502-2.1-00) is a three-dimensional explicit finite difference 
program for solving complex problems in geotechnical, civil, and mining engineering. FLAC3D 
simulates the behavior of three-dimensional structures built of soil, rock, or other materials that 
undergo plastic flow when a limiting yield condition is reached. The explicit, Lagrangian 
calculation scheme and the mixed-discretization zoning technique ensure that plastic collapse 
and flow are modeled very accurately. A detailed discussion on the general features and fields of 
the FLAC3D computer software applications is presented in the User's Manual (Itasca 
Consulting Group 2002). 
FLAC3D was used in mechanical analysis in this calculation. All input and output files for this 
software item used in this calculation have been archived on CD-ROMs and will be submitted to 

Title: Ground Control for Non-Emplacement Drifts for LA 20 
DI: 800-KMC-SSD0-00700-000-00A 
the RPC as part of the records package for this calculation. The results are presented and 
described in Section 6. 
FLAC3D Version 2.1 was obtained from the SCM in accordance with the AP-SI.1Q procedure. 
FLAC3D is installed and run on stand-alone PCs with Windows 2000 operating systems. 
FLAC3D Version 2.1 was qualified for use in design in accordance with the AP-SI.1Q 
procedure. FLAC was appropriate for the applications used in this calculation, and used within 
the range of validation, as specified in the software qualification documentation. 
3.3 3DEC COMPUTER CODE 
3DEC Version 2.01 (STN: 10025-2.01-00) is a three-dimensional computer code based on the 
distinct element method for discontinuum modeling. In 3DEC, the discontinuous medium is 
presented as an assemblage of discrete blocks. Individual blocks behave as either rigid or 
deformable material; deformable blocks are subdivided into a mesh of finite difference elements. 
3DEC is based on a Lagrangian calculation scheme that is well-suited to model the large 
deformations of blocks in a system. A detailed discussion on the general features and fields of 
3DEC computer software applications is presented in the User's Manual (Itasca Consulting 
Group 2002). 
3DEC was used in coupled thermomechanical and seismic analyses in this calculation. All input 
and output files for this software item used in this calculation have been archived on CD-ROMs 
and will be submitted to the RPC as part of the records package for this calculation. The results 
are presented and described in Section 6. 
3DEC Version 2.01 was obtained from the SCM in accordance with the AP-SI.1Q procedure. 
3DEC is installed and run on stand-alone PCs with Windows 2000 operating systems. It is noted 
that in this calculation all computer runs using 3DEC were conducted in batch mode. 3DEC 
Version 2.01 was qualified for use in design in accordance with the AP-SI.1Q procedure. 3DEC 
was appropriate for the applications used in this calculation, and used within the range of 
validation, as specified in the software qualification documentation. 
3.4 SPREADSHEET SOFTWARE 
Microsoft Excel 97 SR-2 spreadsheet software was used in displaying some of the FLAC, 
FLAC3D, and 3DEC results graphically. In this application, results from the FLAC, FLAC3D, 
or 3DEC analyses were used as inputs, and outputs were charts or figures. There are no formulas 
used in the spreadsheets. Microsoft Excel 97 SR-2 is an exempt software product in accordance 
with the AP-SI.1Q procedure. 

Title: Ground Control for Non-Emplacement Drifts for LA 21 
DI: 800-KMC-SSD0-00700-000-00A 
4. INPUTS 
This section presents the data and parameters, criteria, and codes and standards used in the 
calculation. 
4.1 DATA AND PARAMETERS 
4.1.1 Rock Mass Mechanical Properties 
4.1.1.1 Lithophysal Rock 
Rock mass properties for lithophysal rock are listed in Table 4-1. Note that all data are obtained 
from Table 5-8 of Input Parameters for Ground Support Design (BSC 2003a) except that values 
of bulk modulus and shear modulus are calculated based on values of modulus of elasticity and 
Poissons ratio in Table 4-1 and from equations 14 and 13, p. 111 of Jaeger and Cook 1979. The 
values of modulus of elasticity are expressed to two decimal points, which is also consistent with 
other values including decimal numbers. 
Table 4-1. Rock Mass Mechanical Properties for Lithophysal Rock 
Parameter Lithophysal Rock (Tptpul and Tptpll) 
Rock Mass Lithophysal Porosity Category 1 2 3 4 5 6 
Lithophysal Porosity (%) 25-30 20-25 15-20 10-15 <10 N/A 
Poissons Ratio 0.22 0.22 0.22 0.22 0.22 0.22 
Modulus of Elasticity (GPa)* 1.92 6.37 10.82 15.26 19.71 1.0 
Bulk Modulus (GPa) 1.14 3.79 6.44 9.08 11.73 0.60 
Shear Modulus (GPa) 0.79 2.61 4.43 6.25 8.08 0.41 
Cohesion (MPa) 2.07 3.11 4.14 5.18 6.21 1.24 
Friction Angle (degrees) 45 45 45 45 45 45 
Unconfined Compressive Strength (MPa) 10 15 20 25 30 6 
Source: BSC 2003a, Table 5-8. Note: * Values are calculated to second digit based on Eq. I.11.1 of BSC 2003a. 
4.1.1.2 Non-lithophysal Rock 
Rock mass properties for non-lithophysal rock (Tptpmn) are listed in Table 4-2. These values 
are determined based on the rock mass classification. Details on how these values are estimated 
are presented in the Subsurface Geotechnical Parameters Report (BSC 2003c, Section 8.5, Table 
8-41). 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
Table 4-2. Rock Mass Mechanical Properties for Non-lithophysal Rock 
Parameter Non-lithophysal Rock (Tptpmn) 
Rock Mass Quality Category 1 2 3 4 5 
Cumulative Frequency Distribution (%) 10 30 50 70 90 
Geologic Strength Index (GSI) 50.48 55.49 59.03 62.33 66.79 
Rock Mass Quality (Q or Qp) 2.05 3.59 5.31 7.67 12.58 
Elastic Modulus (GPa) 10.25 13.66 16.74 20.23 26.18 
Poissons Ratiob 0.19 0.19 0.19 0.19 0.19 
Bulk Modulus (GPa)a 5.51 7.34 9.00 10.88 14.08 
Shear Modulus (GPa)a 4.31 5.74 7.03 8.50 11.00 
Global Compressive Strength (MPa) 33.50 39.67 44.42 49.50 57.71 
Cohesion (MPa) 7.60 8.69 9.53 10.39 11.75 
Friction Angle (degrees) 40.15 42.29 43.64 44.92 46.66 
Tensile Strength (MPa) 0.08 0.12 0.16 0.21 0.32 
Source: BSC 2003c, Section 8.5, Table 8-41. 
Note: a Bulk modulus and shear modulus are calculated based on equation described in Section 4.1.1.1. b Based on 
Sec. 8.5.2.3 of BSC 2003c 
4.1.2 Rock Mass Density 
A rock mass saturated bulk density of 2,410 kg/m3 is used to estimate overburden and in situ 
stress state. This value is for the rock unit of Tptpln, and is the highest value of lithostratigraphic 
units (BSC 2001, Table 4-2). Therefore, use of this value is conservative for the purpose of this 
calculation. 
4.1.3 Dimensions and Properties for Fully Grouted Rock Bolts 
The dimensions and properties of fully grouted rock bolts are shown in Table 4-3. 
4.1.4 Dimension and Properties for Shotcrete 
The dimension and properties of shotcrete are shown in Table 4-4. 
4.1.5 Dimensions and Properties for Stainless Steel Super Swellex Bolts 
Stainless steel Super Swellex rock bolts are proposed for use in emplacement drifts. The 
dimensions and mechanical properties of stainless steel Super Swellex bolts are listed in Table 45. 
4.1.6 Seismic Velocity History 
Seismic velocity histories for the mean annual exceedance probability of 1 x 10-4 (10,000 years) 
at repository horizon and 5 x 10-4 (2,000 years) for the surface facilities area are shown in 
Figures 4-1a (DTN: MO0306SDSAVDTH.000) and 4-1b (DTN: MO0310SDSTMHIS.003), 

Title: Ground Control for Non-Emplacement Drifts for LA 23 
DI: 800-KMC-SSD0-00700-000-00A 
respectively. Details on how these seismic velocity histories are applied in numerical 
calculations are discussed in Section 6.5.2.3. 
4.1.7 Rock Mass Thermal Expansion Coefficients 
The coefficients of thermal expansion for rock mass at TSw2 thermal mechanical unit is shown 
in Table 4-6. 
Table 4-3. Dimensions and Properties of Fully Grouted Rock Bolts 
Parameter Value Source and Remark 
Diameter of Rock Bolt (m) 0.0254 Converted from a diameter of 1 inch (1 in  0.0254 m/in 
= 0.0254 m) 
Thickness of Grout Annulus (m) 0.00635 Converted from a thickness of 0.25 in (0.25 in  0.0254 
m/in = 0.00635 m) 
Length of Rock Bolts (m) 3 - 5 3 m in typical non-emplacement drifts and increase up to 
5 m in intersection area 
Spacing between bolts and rows 
(m) 1.25 Spacing may be changed to 1.5 m if necessary 
Perimeter of Rock Bolt (m) 0.08 Calculated: p = p D = 3.1415  0.0254 = 0.08 m 
Allowable Axial Force (kN)a 264 Based on the yield strength (force) of 264 kN (DSI 2002, 
Dywidag threadbar) 
Modulus of Elasticity of Steel 
(GPa) 200 AISC 1997, p. 1-117 (29,000 x1000/(145x 106)) 
Modulus of Elasticity of Grout 
(GPa) 14 Onofrei et al. 1993, Figure 33, p. 60 
Poissons Ratio of Grout 0.25 Set to be the same as concrete (see Merritt 1983, p. 6-8) 
Grout Unconfined Compressive 
Strength (MPa) 90 Onofrei et al. 1993, Figure 27b, p. 52 
Bond Stiffness of Grout (N/m/m) 8.68108 Calculated using Equation 3 of Ruest and Martin 2002 
Bond Strength of Grout 
(cohesion) (N/m) 
1.9105a 
3.0105b 
Based on recommendation by Hutchinson and 
Diederichs (1996, Figure 2.6.13). a for lith. Cat. 1 tuff, 
b for non lith. Cat. 1 tuff. See discussion in Sec. 6.5.5.1. 
Table 4-4. Dimension and Properties of Shotcrete 
Parameter Value Source 
Thickness of Shotcrete (m) 0.1 Converted from a thickness of 4 in (4 in  0.0254 m/in = 
0.102 m) 
Elastic Modulus E (GPa) 29 Based on mean value in Sec. 1.7 of ACI 506R-90(95) 
Poissons Ratio . 0.25 Assumed same as concrete, Merritt (1983, p. 6-8) 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
Table 4-5. Dimensions and Properties for Stainless Steel Super Swellex Rock Bolts 
Parameter Value Source 
Diameter of Rock Bolt (m) 0.054 Atlas Copco 2003, p. 10. 
Thickness of (m) 0.003 Atlas Copco 2003, p. 10. 
Youngs Modulus of Stainless 
Steel (GPa) 200 
The value is 193 GPa for 316 type stainless steel 
based on Table 21, p. 871 of ASM International 1990. 
It is rounded off to 200 GPa for conservatism. 
Limit Axial Force (kN) 260 See Assumption 5.5. 
Bond Stiffness (N/m/m) 3108 See Section 6.5.5.1. 
Bond Strength (N/m) 2.75105 See Section 6.5.5.1. 

Title: Ground Control for Non-Emplacement Drifts for LA25DI: 800-KMC-SSD0-00700-000-00A-60-40-20020406001020304050607080Time (Seconds)
Velocities (cm/s)
VH1H2Source: DTN: MO0306SDSAVDTH.000(a)
Source: DTN: MO0402SDSTMHIS.004(b)
Figure 4-1. Time Histories of Velocity Components of Seismic Motion for Mean Annual ExceedanceProbability of: (a) 1x10-4; (b) 5x10-4

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
Table 4-6. Coefficients of Thermal Expansion for TSw2 Unit 
Temperature Coefficient of thermal expansion 
C 1/C 
0-50 7.1410-6 
50-75 7.4710-6 
75-100 7.4610-6 
100-500 9.0710-6 
Source: DTN: SNL01B05059301.006 
4.1.8 Rock Mass Mechanical Properties of TCw Thermal Mechanical Unit 
Table 4-7 lists the rock mass mechanical properties for the TCw thermal mechanical unit. The 
mechanical properties are from Table 5-9 of BSC 2003a. 
Table 4-7. Rock Mass Mechanical Properties of the TCw Unit 
Thermal Mechanical Unit TCw 
Density (kg/m3) 2310 
Rock Mass Quality (RMQ) 1 5 
Modulus of Elasticity (GPa) 7.33 29.36 
Poissons Ratio 0.21 0.21 
Cohesion (MPa) 1.5 3.9 
Friction Angle (degrees) 54 57 
Tensile Strength (MPa) 0.97 2.35 
Source: BSC 2003a, Table 5-9 
4.2 DESIGN CRITERIA AND CONSTRAINTS 
4.2.1 Criteria 
The following criteria are applicable to the design of ground support system in non-emplacement 
drifts: 
4.2.1.1 The ground control system shall be designed to maintain adequate operating envelopes 
through permanent closure for access mains, exhaust air mains and miscellaneous 
openings (Minwalla 2003, Section 4.5.2.1). 
4.2.1.2 The ground control system shall accommodate geologic mapping of 100 percent of 
non-emplacement openings (Minwalla 2003, Section 4.5.2.1). 

Title: Ground Control for Non-Emplacement Drifts for LA 27 
DI: 800-KMC-SSD0-00700-000-00A 
4.2.1.3 The system shall be designed for the appropriate worst case combination of in situ, 
thermal, seismic, construction, and operational loads (Minwalla 2003, Section 4.5.2.1). 
4.2.1.4 The ground control system for non-emplacement openings shall be designed for safety 
factors compatible with maintenance plans (Minwalla 2003, Section 4.5.2.1). 
4.2.1.5 The ground control system shall use materials having acceptable long-term effects on 
waste isolation (Minwalla 2003, Section 4.5.2.2). 
4.2.1.6 The ground control system shall be designed to withstand a design basis earthquake 
(Minwalla 2003, Section 4.5.2.2). 
4.2.1.7 The ground control system shall be designed to prevent rock falls that could potentially 
result in personnel injury (Minwalla 2003, Section 4.5.2.3). 
4.2.1.8 The ground control system shall accommodate the maintenance of non-emplacement 
openings (Minwalla 2003, Section 4.5.2.6). 
4.2.2 Constraints 
The following design constraints are applicable to the ground support system in non-
emplacement drifts: 
4.2.2.1 Diameters of Emplacement Drift and Access/Exhaust Main 
The diameter of emplacement drift is 5.5 m and the diameter of access mains and exhaust mains 
is 7.62 m (BSC 2004, Table 1). 
4.2.2.2 Turnout Drift Dimension 
The turnout has a horseshoe shape in cross-section (8 m wide and 7 m high). The excavation 
springline is 3 m above the sill and the turnout crown section has a radius of 4 m (BSC 2004, 
Figure 3 and Table 1). 
4.2.2.3 Dimension and Configuration of Observation Drift and Test Alcove 
Both observation drift and test alcove have a horseshoe shape in cross-section (5 m wide and 5 m 
high). The observation drift is parallel to the emplacement drift in the plan view. Horizontal 
spacing, center-to-center, between the observation drift and the emplacement drift is 20 m. Along 
most of its length the vertical distance between the crown of the observation drift and the floor of 
the emplacement drift is 10 m. (Note that observation drift is below the emplacement drift). See 
Figures 6-15 and 6-15A for illustration of observation drift configuration. The test alcove is 
below the emplacement drift with a minimum of 10 m from crown to invert (BSC 2003e, Section 
8.5). 

Title: Ground Control for Non-Emplacement Drifts for LA 28 
DI: 800-KMC-SSD0-00700-000-00A 
4.2.2.4 TBM Launch Chamber Dimension and Configuration 
The TBM launch chamber is a horseshoe-shaped underground excavation. It is 11 m  11 m in 
cross-section and 20-m long. It will be excavated by conventional drill-and-blast method (BSC 
2003e, Section 8.3). 
4.2.2.5 North Portal Starter Tunnel Dimension and Configuration 
The North Portal Starter Tunnel was excavated nominally 9.8 (rounded-off to 10) m high by 9.9 
(rounded-off to 10) m wide in an arched shape. It was constructed by conventional drill-and-
blast method. (CRWMS M&O 1997, Section 7.1). The length of Starter Tunnel is about 60 m 
(BSC 2003e, Figure 3). 
4.3 CODES AND STANDARDS 
The codes and standards applicable to ground support system in non-emplacement drifts are as 
listed below. 
4.3.1 American Society for Testing and Materials (ASTM) 
ASTM A 36/ Standard Specification for Carbon Structural Steel 
A 36M-03a 
ASTM A 82-97a Standard Specification for Steel Wire, Plain, for Concrete 
Reinforcement 
ASTM A 242/ Standard Specification for High-Strength Low-Alloy Structural Steel 
A 242M-03a 
ASTM A 588/ Standard Specification for High-Strength Low-Alloy Structural Steel 
A 588M-03 with 50 ksi [345MPa] Minimum Yield Point to 4in. [100mm] Thick 
ASTM F 432-95 Standard Specification for Roof and Rock Bolts and Accessories 
(Reapproved 2001) 
4.3.2 American Institute of Steel Construction (AISC) 
AISC 1997 Manual of Steel Construction  Allowable Stress Design 
4.3.3 American Concrete Institute (ACI) 
ACI 506R-90 (95) Guide to Shotcrete 
ACI 506.2-95 Specification for Shotcrete 

Title: Ground Control for Non-Emplacement Drifts for LA 29 
DI: 800-KMC-SSD0-00700-000-00A 
5. ASSUMPTIONS 
The following assumptions were made in this calculation: 
5.1 INITIAL GROUND RELAXATION 
An initial ground relaxation value of 75 percent is assumed and used in the ground support 
analysis for the fully-grouted rock bolting in non-emplacement drifts. This ground relaxation 
value results in 25 percent of the pre-excavation in situ stress being imposed on the ground 
support system. This same ground relaxation value for rock bolts is also confirmed from ground 
control analysis for emplacement drifts (see Section 6.1.5 of BSC 2003f). This is considered 
adequate since the initial relaxation will likely be completed before the final support installed. 
Used in Section 6. 
An initial ground relaxation value of 100 percent is assumed and used in the ground support 
analysis for the final shotcrete lining in non-emplacement drifts. This value is considered to be 
adequate because the shotcrete lining will usually be installed after the drift excavation. Any 
rock deformation induced by the excavation will most likely be complete well before the 
application of the shotcreting. Used in Section 6. 
5.2 AVERAGE DEPTH OF REPOSITORY HOST HORIZON 
The average depth of repository host horizon is assumed to be 400 m measured from the center 
of an emplacement drift (BSC 2003a, Tables 5-2a to 5-2c). Depth of emplacement drifts varies 
from drift to drift. The depths near the centers of Panels 1, 3 East and 3 West are approximately 
296 m, 259 m, and 372 m, respectively. The depths of emplacement drifts range from 215 to 450 
m with the majority between 300 to 400 m. In a recent scoping analysis on emplacement drift 
stability, it was indicated that the results for the maximum value of 450 m would be similar to 
those for the bounding case of 400 m (BSC 2003g, Section 4.2.1). Since non-emplacement drifts 
are in the same horizon as that for emplacement drifts, it is therefore, considered adequate to 
make the same assumption for the purpose of this calculation. Used in Section 6. 
5.3 HORIZONTAL-TO-VERTICAL IN SITU STRESS RATIOS 
The major horizontal principal stress with a direction of N15oE is 62% of the vertical stress 
whereas the minor horizontal principal stress with a direction of N75oW is 36% of the vertical 
stress. This is according to the in situ stress measurement by hydraulic fracturing in a test hole 
located in the TSw2 unit (SNF37100195002.001). This assumption is used for 3-dimensional 
modeling in Section 6. 
The horizontal-to-vertical in situ stress ratios (Ko) is assumed to be 0.5 for two-dimensional 
modeling. This value is approximately the average of the minimum and maximum K0 values for 
the 3-dimensional modeling. 
This assumption is considered adequate for the purpose of this calculation. Used in Section 6. 

Title: Ground Control for Non-Emplacement Drifts for LA 30 
DI: 800-KMC-SSD0-00700-000-00A 
5.4 TENSILE STRENGTH OF LITHOPHYSAL ROCK 
Tensile strength is assumed to be equal to a half of cohesion for various categories of the 
lithophysal rock. For example, tensile strength of category 1 rock mass is equal to 1.04 MPa 
since cohesion for this category rock mass is 2.07 MPa as indicated in Table 4-1. This value is 
about 10% of the corresponding unconfined compressive strength value, and is considered 
reasonable for the lithophysal rock mass. Used in Section 6. 
5.5 LIMIT AXIAL FORCE FOR STAINLESS STEEL SUPER SWELLEX BOLTS 
The limit axial force for stainless steel Super Swellex bolts is assumed to be 260 kN in the 
calculation. Used in Section 6. 
The tensile strength of stainless steel 316 is 620 MPa (ASTM A 276-03, Table 2, p. 4). The limit 
axial force of 260 kN is calculated from 620106p/4(0.0542-0.0482)/1000=298 kN and using a 
reduction factor for conservatism. It is considered adequate for the purpose of this calculation. 
5.6 TENSILE STRENGTH IN THE 3DEC SIMULATIONS 
To avoid large, unrealistic deformation of zones around the boundary of the excavation that yield 
in tension due to numerical simulation, tensile strength in the 3DEC simulations was assumed to 
be infinitely large. This assumption is to facilitate numerical convergence without sacrificing the 
drift response. This is considered adequate for the purpose of this calculation. Used in Section 6. 
5.7 PROPERTIES OF INTERFACE BETWEEN SHOTCRETE AND ROCK 
The properties of interface between shotcrete and rock are assumed as follows: cohesion  1x109 
Pa, angle of friction  45o, bond stiffness  2x109 N/m/m, bond shear strength  2 x109 N/m, and 
tensile strength  1 x109 Pa. Used in Section 6. 
These property values are assumed in 3DEC modeling. Since shotcrete is assumed to be linearly 
elastic and the interface between shotcrete and rock is intimate and cannot break. It is 
considered adequate for the purpose of this calculation. 
5.8 GROUND SUPPORT TYPES 
The following ground support systems are assumed for non-emplacement drifts in this 
calculation. Used throughout. 
Non-emplacement openings except intersections between access mains and turnouts or between 
exhaust mains and emplacement drifts, ramps, and starter tunnel  fully grouted bolts with heavy 
duty welded wire fabric (WWF). 
Intersections between access mains and turnouts or between exhaust mains and emplacement 
drifts, ramps and starter tunnel - fully grouted bolts with fiber-reinforced shotcrete, and lattice 
girders as necessary for roof span control. It should be noted that stainless steel Super Swellex 
bolts are used at the small intersection area inside the emplacement drifts, which is consistent 
with the current ground support design in emplacement drifts. A very small amount of shotcrete 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
is used in this portion of emplacement drift for enhancing the opening stability. However, 
shotcrete might not be used if this small amount of cementitious material would cause significant 
impact on waste package performance. Figure 5-1 illustrates shotcreted area at the intersections 
with fresh air turnout (i.e., between access main and turnout) and exhaust air turnout (i.e., 
between exhaust main and emplacement drift). Note that shotcrete area will cover the crown and 
extend below the springline covering an arc of about 240o for circular openings and 2/3 of the 
total surface area in horseshoe-shaped openings. 
For the ground support at North Portal, fully grouted rock bolts with fiber-reinforced shotcrete 
will be installed around the portal frontal and lateral faces. 
The above ground support systems are developed based on the current ground support design 
concept, the majority of which are also similar to the ones assumed in Section 5.2 of Committed 
Ground Support Materials for LA Design (BSC 2004), and are considered appropriate for this 
calculation. 
Note: Plan view of underground layout is for illustration only. The correct layout is shown in Figure 6-1 
Source: BSC 2004, Figure 4 
Figure 5-1. Illustration of Shotcreted Areas at Intersections 

Title: Ground Control for Non-Emplacement Drifts for LA 32 
DI: 800-KMC-SSD0-00700-000-00A 
6. GROUND SUPPORT DESIGN 
6.1 INTRODUCTION 
This calculation provides the technical basis for the design of the ground control systems for 
non-emplacement drifts for License Application (LA). In the design of permanent underground 
facilities, the ground support needs are determined using several available methods, i.e., 
empirical methods, observational approach, and analytical methods (Bieniawski 1984, p. 35). 
The design and construction of an underground high-level nuclear waste repository facility 
introduce unique challenges that are not commonly experienced for other subsurface facilities. 
For example, the presence of high level nuclear waste and the resultant thermal loading 
conditions introduce a series of additional requirements to the overall design and construction of 
the facility in addition to the waste isolation requirements. In situ (excavation) loads, potential 
loads from repository operations, and loads due to seismic loading conditions during an 
earthquake must also be addressed in the design. 
The non-emplacement drift ground support design effort includes both empirical and analytical 
methods, coupled by observational approach (i.e., engineering experience). The empirical 
methods are primarily used for assessing the needs for ground support of non-emplacement drifts 
as well as for its selection. Design issues such as personnel safety, constructibility, and geologic 
mapping requirements should be factored into the design of the ground support system at this 
approach. However, due to the complexity of the problem involved with ground support design 
for nuclear waste repository, the repository ground control design effort will focus mainly on 
analytical methods by using computer programs to evaluate the stability of unsupported and 
supported openings. Applicable thermal and seismic loads will be considered in the design in 
addition to the in situ loading conditions. The best available experience of drift stability for the 
repository host rock can be obtained from observation of stability condition of openings in the 
Exploratory Studies Facility (ESF) and the Enhanced Characterization of the Repository Block 
(ECRB), which will also be discussed. 
The non-emplacement drifts include all other repository underground openings that will not be 
used for waste emplacement, such as mains, exhaust mains, ventilation shafts and raises, 
emplacement drift turnouts, and intersections between turnouts and mains, etc. For non-
emplacement drifts, both the initial and final ground support systems will be developed, but only 
the final ground support systems will be discussed in this calculation. Also, as indicated in 
Section 1, the ground support design in this calculation will include access mains, exhaust mains, 
turnouts, ramps, intersections between access mains and turnouts and between exhaust mains and 
emplacement drifts, observation drift, TBM launch chambers, and portals. 
It should be noted that the design for initial ground support installed during excavation is not 
included in this calculation. Initial ground support is placed in close proximity to the advancing 
tunnel face for worker safety and prevention of equipment from damage due to rock fall until the 
permanent support system is installed. Note that initial ground support may become the final 
ground support depending on its type and tunnel construction method. Usually, the initial 
ground support system consists of rock bolts and wire mesh but it may change depending on 
local ground condition and excavation method. In this calculation, no credit or account will be 

Title: Ground Control for Non-Emplacement Drifts for LA 33 
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given for the initial ground support in modeling the final ground support systems for non-
emplacement drifts. 
6.2 REPOSITORY HOST HORIZON AND GEOTECHNICAL CHARACTERIZATION 
In order to design the ground support system in non-emplacement drifts, it is important to 
understand the environment in which these drifts will be located. 
According to the Underground Layout Configuration calculation (BSC 2003e, Section 7.1.7), the 
repository host horizon (RHH) will be located in the lower part of the lithophysal zone of the 
densely welded devitrified lithophysal-rich tuff unit and the entire densely welded devitrified 
lithophysal-poor tuff unit of the Topopah Spring Tuff. The RHH contains four lithostratigraphic 
units, namely the upper lithophysal unit (Tptpul), the middle non-lithophysal unit (Tptpmn), the 
lower lithophysal unit (Tptpll), and the lower non-lithophysal unit (Tptpln). The underground 
layout and the geological units within each panel is shown in Figure 6-1 (BSC 2003e, Figure II2). 
Of the total emplacement areas, approximately 85% will lie within the Tptpll and Tptpul 
units combined and the rest 15% will be located in the Tptpmn and Tptpln units. In the 
lithophysal rock, 95% of the emplacement area lies in the Tptpll unit whereas in the nonlithophysal 
rock, 83% of the emplacement area lies in the Tptpmn unit (BSC 2003e, Table II-2). 
In other words, among the area for non-emplacement drifts, the primary tuff rock units for 
lithophysal and non-lithophysal rocks are Tptpll and Tptpmn units, respectively. 
Geological mapping was conducted in the ESF tunnel and the ECRB drift to characterize the 
rock units within the RHH. The data collected were analyzed using two empirical rock mass 
classification systems, the Geomechanics Rock Mass Rating (RMR) system (Bieniawski 1989) 
and the Rock Mass Quality (Q) system of Norwegian Geotechnical Institute (NGI) (Barton et al. 
1974). The geotechnical characteristics of the Tptpll and Tptpmn units are summarized as 
follows (Board 2003, Section 3.4): 
Tptpll Unit. The Tptpll unit has a horizontal RQD rating of 42 (poor), a Q rating of 7.9 (fair), 
and a RMR value of 57 (fair). Its content of lithophysae varies from 5 to 30 percent by volume, 
with a size ranging from 5 to 130 cm. The larger lithophysal cavities tend to be irregular or 
ellipsoidal features that exhibit prismatic fracturing. The unit has an average of two plus random 
joint sets; however no keyblock problems are expected. Typical lithophysae and fracturing in 
the Tptpll are shown in Figure 6-2. 
Tptpmn Unit. The Tptpmn unit has a mean horizontal RQD rating ranging from 60 to 62 (fair), 
and a RMR value of 60 (fair). It is characterized by less than 3 percent lithophysae by volume. 
This unit has an average of three to three plus random joint sets, with predominately two vertical 
joint sets and one horizontal joint set. The horizontal joint set, or vapor-phase partings, is the 
primary cause of potential formation of keyblocks. A typical fracture pattern in the Tptpmn unit 
is shown in Figure 6-3. 
6.3 GROUND SUPPORT FUNCTIONS 
Ground support for non-emplacement drifts has the following functional and performance 
requirements (see Section 4.2.1 and Board 2003, Section 7.3): 

Title: Ground Control for Non-Emplacement Drifts for LA 
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 Ensure stable conditions for operational worker safety 
 Prevent rock deformation to acceptable levels so that operational envelopes are 
maintained 
 Allow for geological mapping, performance confirmation activities (which may include 
remote observation and possible field testing), waste retrieval operations, and closure 
operations (which may include installation of permanent drip shields) 
 Prevent for rock falls that could potentially result in personnel injury 
 Use materials with acceptable long-term effects on waste isolation 
 Accommodate the planned or unplanned maintenance over the operational life of the 
repository. 

Title: Ground Control for Non-Emplacement Drifts for LA35DI: 800-KMC-SSD0-00700-000-00ASource: BSC 2003e, Figure II-2.
Figure 6-1. Underground Layout and Geological Units by PanelDBACNote: A, B, C D, and E inthe rectangular boxesshow typical modellocations.
EIntakeShaft #1

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
Source: Board 2003, Figure 10b 
Figure 6-2. Lithophysae and Fracturing in the Tptpll Unit 
Source: Board 2003, Figure 8 
Figure 6-3. Fractures in the Wall of the ECRB in the Tptpmn Unit 

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6.4 EMPIRICAL METHODS 
Empirical methods are usually applicable to designing tunnels in mining and civil engineering. 
The empirical approach relies on rock mass classification systems. Generally, these systems 
allow rock properties and geologic conditions shown in samples taken from boreholes, scanline, 
full-peripheral mapping, detailed test surveys, and certain outcrops at the planned site, to be 
compared with similar information compiled and categorized from existing underground 
facilities. Based on this comparison, support requirements or needs can be estimated. 
To date, many rock mass classification systems have been proposed, such as systems by Deere 
and Deere (1988), Wickham et al. (1972), Bieniawski (1989), Barton et al. (1974), and Barton 
(2002). Two common classification systems were recommended in the Drift Design 
Methodology and Preliminary Application for the Yucca Mountain Site Characterization Project 
(Hardy and Bauer 1991, p. 6-6) and in the Support of Underground Excavations in Hard Rock 
(Hoek et al. 2000, p. 44). These two classification systems are the RMR value developed by 
Bieniawski (1989), and the Q value developed by Barton (see Barton 2002). Both methods 
incorporate geological, geometric and design/engineering parameters in arriving at quantitative 
value of their rock mass quality (Hoek et al. 2000, p. 44). The latter system was used in 
designing the ESF, and the successful performance of that facility has lead to adoption of this 
system for analysis of ground control elements for SR (BSC 2001, p. 50). These two 
classification systems are briefly described in the following sections. A detailed description of 
them can be found in Bieniawski (1989) and Barton et al. (1974), respectively. 
6.4.1 Geomechanics Rock Mass Classification System 
The Rock Mass Rating (RMR) system, otherwise known as the Geomechanics Classification, 
was developed by Bieniawski during 1972  1973 (Bieniawski 1989, p. 51). The following six 
parameters are used to classify a rock mass using the RMR system (Bieniawski 1989, p. 52): 
1. Uniaxial compressive strength of rock material 
2. Rock quality designation (RQD) 
3. Spacing of discontinuities 
4. Condition of discontinuities 
5. Groundwater conditions 
6. Orientation of discontinuities 
These parameters not only are measurable in the field but can also be obtained from borings. 
Joints are the major factor in this classification system; four of the six parameters (RQD, joint 
spacing, joint conditions, and orientation of joints) are related to joint characteristics. Increments 
of rock mass rating corresponding to each parameter are summed to determine RMR. 
The RMR values for various rock units at the repository host horizon are generally available 
from data collected in the ESF and ECRB. In case these RMR values are not available, 
empirical correlation can be used to estimate RMR values based on known deformation modulus 
of rock mass. The empirical correlation used in this calculation is as follows (Barton 2002, p. 
193, Eq. 8): 

Title: Ground Control for Non-Emplacement Drifts for LA 
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RMR
=
log 40 E
m
+
10 (Eq. 6-1) 
where Em = deformation modulus of rock mass in GPa 
Once the RMR values are determined, the rock mass quality for each rock unit considered can be 
judged based on the guidelines provided by Bieniawski (1989, Tables 4.1 and 4.2). 
Recommendation for the excavation scheme and rock support needs can be made by following 
the guidelines presented in Table 6-1 (Bieniawski 1989, Table 4.4). Details on how to apply the 
RMR classification system to the preliminary design of ground support for rock tunnels such as 
those in repository non-emplacement drifts can be found in the Engineering Rock Mass 
Classifications (Bieniawski 1989, Section 4). 
6.4.2 NGI Rock Mass Classification System 
The NGI rock mass classification or the rock mass quality Q system, developed in earlier 1970s 
by Barton (2002, p. 187), is commonly used in the design of rock support for tunnels and large 
underground chambers. The six parameters in the NGI system include 1) RQD, 2) joint set 
number (Jn), 3) joint roughness number (Jr), 4) joint alteration number (Ja), 5) joint water 
reduction factor (Jw), and 6) stress reduction factor (SRF). 
These parameters chosen to describe the rock mass quality Q are combined in the following way 
(Barton et al. 1974, Eq. 1): 
.
...
...
..
. 
...
J
J
RQD
..
.
(Eq. 6-2)
..
. 
=
Q
w
r 
J 
J
SRF
.
.
.
n
a 
The three ratios in the equation - RQD/Jn, Jr/Ja, and Jw/SRF - represent block size, minimum 
inter-block shear strength, and active stress, respectively (Barton et al. 1974, p. 202). 
Similar to the RMR values, Q indices for various rock units at the repository host horizon are 
generally available from the ESF and ECRB. In case these Q indices are not available, empirical 
correlation can be used to estimate Q indices based on given rock mass modulus. The empirical 
correlation used in this calculation is as follows (Hoek et al. 2000, Eqs. 8.16 and 8.19): 
RMR 
44
-
J 
J 

' 
Q
=
Q
9 
(Eq. 6-3)
w 
w

= 
e
SRF 
SRF 
where 
.
...
...
...
.
J
RQD
Q
'
..
. 
=
r 
J 
J
.
.
n 
a 
The RMR value in Eq. 6-3 is estimated from the given deformation modulus of rock mass using 
Eq. 6-1. 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Table 6-1. Estimate of Ground Support Needs Based on RMR System 
Source: Bieniawski 1989, Table 4.4. 
The Q index is used with the Equivalent Dimension, defined as the largest of span, diameter, and 
height divided by the excavation support ratio (ESR). ESR is roughly analogous to the inverse 
of the factor of safety used in engineering design. The ESR reflects the degree of safety and 
ground support required for an excavation as determined by the purpose, presence of machinery, 
personnel, etc., to meet safety requirements. In essence, the safety factor of an opening can be 
increased by reducing the ESR value. The ESR values for various underground openings can be 
estimated based on Barton et al. (1974, Table 7). For example, the ESR for the access and 
exhaust mains and the turnouts, classified as access tunnels, is taken to be 1.3; the intersections 
are assigned a value of 1.0. 

Title: Ground Control for Non-Emplacement Drifts for LA 
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The Equivalent Dimension is plotted against Q on the design chart, Figure 6-4, to determine the 
required ground support need (Hoek et al. 2000, Figure 4.3). 
The NGI rock mass classification system provides guidance on bolt spacing, bolt length, and 
shotcrete thickness, based on the rock mass quality index (Q) and the opening dimensions. The 
length L of rockbolts can be estimated from the excavation width B and ESR (Hoek et al. 2000, 
Eq. 4-3): 
2 + 15.0 B
L = (Eq. 6-4)
ESR 
Source: Hoek et al. 2000, Figure 4.3. 
Figure 6-4. Estimated Ground Support Needs Based on Q Index 

Title: Ground Control for Non-Emplacement Drifts for LA 41 
DI: 800-KMC-SSD0-00700-000-00A 
6.4.3 Empirical Analysis of Ground Support Need 
Results of the empirical analysis based on the RMR and Q approaches for non-emplacement 
drifts in the non-lithophysal rock are presented in the following. Note that the empirical analysis 
is only considered for access/exhaust mains, intersections between access mains and 
emplacement drift turnouts, and intersections between exhaust mains and emplacement drifts, 
which comprise of majority of non-emplacement drifts; observation drift, TBM launch 
chambers, and portals are not included. The methods used are described in Sections 6.4.1 and 
6.4.2. 
For the non-emplacement drifts excavated in the lithophysal rock, however, use of the RMR or Q 
approach for the ground support design is non-conventional, and there are no sufficient data or 
field experiences available to support this application. This is primarily due to the fact that the 
lithophysal rock contains air-filled large cavities and is hard to be characterized using the RMR 
or Q index since a RQD value is defined for a rock with fractures not with voids. Therefore, 
these empirical methods are not used in this calculation for evaluating the requirements of 
ground support for non-emplacement drifts in the lithophysal rock. Selection of ground support 
methods for this rock type is based on experiences and observations from the construction of the 
ESF and the ECRB tunnels, and assessment from numerical analyses. 
Based on Geomechanics Rock Mass Classification approach, the ground support need for nonlithophysal 
rock in non-emplacement drifts can be estimated based on Table 6-1. First, RMR 
values are calculated from Em (elastic modulus) values shown in Table 4-2 based on Eq. 6-1. It 
needs to be noted that RMR values are not available, so the correlation between RMR and Em is 
used. Table 6-2 shows the RMR values (under third column) for rock mass categories 1 to 5. 
By comparing RMR values for various rock mass categories shown in Table 6-2 with those in 
Table 6-1, the following ground support needs are estimated: bolt lengths range from 3 m in the 
good rock (categories 4-5) to 4 m in the fair rock (categories 1-3). Bolt spacing ranges from 2.5 
m in the good rock to 1.5 - 2 m in the fair rock. In the good rock, 50-mm of shotcrete is 
recommended where required. In the fair rock, 50-100 mm of shotcrete in crown with 30-mm in 
the sides is recommended. It should be noted that Bieniawskis recommendations for ground 
support shown in Table 6-1 was based on a 10-m wide, horseshoe-shaped tunnel (Bieniawski 
1984, Table 6.11). Because this span is larger than those of the excavations outside the 
intersection, and because the support was designed for a civil application, the estimated ground 
support needs are considered very conservative for use at Yucca Mountain. 
For NGI Rock Mass Classification approach, Q values for various rock mass categories need to 
be calculated. The following steps are needed for estimating the Q values: 
 Calculate the RMR values (note that RMR values are already calculated, see the previous 
paragraph). 
 Determine the unconfined compressive strength (sc) of intact non-lithophysal rock 
(Tptpmn), which is about 207 MPa (BSC 2003a, Table 5-5). 

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 Estimate the major principal stress (s1) of rock adjacent to non-emplacement drifts, 
which is estimated to be about 20 to 56 MPa (see Section 6.5.3.1). 
 Calculate the ratio sc/s1 to be in the ranges of 4 to 10. 
 Assign SRF value according to Table 4.6 of the Support of Underground Excavations in 
Hard Rock (Hoek et al. 2000, p. 43). A SRF value ranging from 0.5 to 2 is considered 
appropriate. 
Based on Eq. 6-3 and Jw of 1, Q values are calculated for rock mass categories 1 to 5 and listed 
in Table 6-2 (under fifth column). 
The lengths of rock bolts calculated based on Eq. 6-4 for various roof spans and ESR values are 
shown in Table 6-3. Note that equivalent dimension, De, is obtained by dividing the span, B, by 
ESR (Hoek et al. 2000, p. 39). 
Table 6-2. RMR and Q Values for Non-lithophysal Rock 
Category Em (GPa) RMR SRF Q 
1 10.25 50 0.5  2.0 0.97  3.90 
2 13.66 55 0.5  2.0 1.70  6.79 
3 16.74 59 0.5  2.0 2.65  10.59 
4 20.23 62 0.5  2.0 3.69  14.78 
5 26.18 67 0.5  2.0 6.44  25.76 
Table 6-3. Equivalent Dimension De and Bolt Length for Various Openings 
Opening Types Span (m) ESR De (m) Bolt Length (m) 
Access and Exhaust Main 7.62 1.3 5.9 2.4 
Turnout 8.00 1.3 6.2 2.5 
Intersection (Typical) 13.5 1.0 13.5 4.0 
Intersection (largest) 22.0 1.0 22.0 5.3 
Figures 6-5 through 6-8 are plots of ground support needs for various rock mass categories for 
access/exhaust mains, turnouts, typical intersection, and largest intersection, respectively. These 
plots are derived based on De and Q values listed in Table 6-3. 
Tables 6-4 through 6-7 show the ground support needs for various non-emplacement drift 
excavations under rock mass categories 1 to 5 as determined from Figures 6-5 through 6-8. 
Based on these tables, the estimated ground support needs can be summarized as follows: 

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 For access and exhaust mains: bolt length - 2.4 m, bolt spacing - 1.7 m to 2.7 m, 
unreinforced shotcrete - 40 to 50 mm thick for category 1 rock whereas none to 50 mm 
thick for categories 2 to 4 rock, and no shotcrete for category 5 rock. 
 For turnouts: bolt length - 2.5 m, bolt spacing - 1.7 m to 2.7 m, unreinforced shotcrete 40 
to 50 mm thick for category 1 rock whereas none to 50 mm thick for categories 2 to 4 
rock, and no shotcrete for category 5 rock. 
 For intersections: bolt length  4.0 m to 5.3 m, bolt spacing - 1.7 m to 2.7 m, fiber-
reinforced shotcrete - 40 to 120 mm thick for categories 1 to 3 rock whereas unreinforced 
shotcrete none to 90 mm thick for categories 4 and 5 rock. 
As discussed in Section 6.2, typical rock conditions in non-lithophysal rock (Tptpmn) are close 
to fair and good (Geomechanics RMR approach, see Table 6-1) or support categories 3 and 4 
(NGI Q approach, see Figure 6-4). Therefore, by comparing the recommendations made by 
both RMR and Q approaches, use of pattern bolting with 3 m long, spaced at 1.25 to 1.5 m, in 
conjunction with 40 to 50 mm thick shotcrete is considered adequate for access and exhaust 
mains and turnouts. Note that since use of shotcrete is only limited to intersections in non-
emplacement drifts, use of shotcrete in access and exhaust mains and turnouts outside of 
intersection area is replaced by the heavy duty welded wire fabric (Assumption 5.8). For ground 
support needs at intersections, use of pattern bolting with bolt length of 5 m, spaced at 1.25 to 1.5 
m, and in conjunction with 100 mm thick fiber-reinforced shotcrete is considered adequate. 
Figure 6-5. Ground Support Needs for Rock Mass Categories 1 to 5 for Access and Exhaust Mains 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-6. Ground Support Needs for Rock Mass Categories 1 to 5 for Turnouts 
Figure 6-7. Ground Support Needs for Rock Mass Categories 1 to 5 for Typical Intersection 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
Figure 6-8. Ground Support Needs for Rock Mass Categories 1 to 5 for Largest Intersection 
Table 6-4. Ground Support Needs for Access and Exhaust Mains based on Q Values 
Category Bolt Spacing 
(m) 
Bolt Length 
(m) 
Shotcrete 
Thickness 
(mm) 
Other 
1 1.7  2.1 2.4 40  50 none 
2 1.8  2.2 2.4 none  50 none 
3 1.9  2.3 2.4 none - 40 none 
4 2.0  2.4 2.4 none  40 none 
5 2.2  2.7 2.4 none none 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Table 6-5. Ground Support Needs for Turnouts based on Q Values 
Category Bolt Spacing 
(m) 
Bolt Length 
(m) 
Shotcrete 
Thickness 
(mm) 
Other 
1 1.7  2.1 2.5 40  50 none 
2 1.8  2.3 2.5 none  50 none 
3 1.9  2.3 2.5 none  40 none 
4 2.0  2.4 2.5 none  40 none 
5 2.2  2.7 2.5 none none 
Table 6-6. Ground Support Needs for Typical Intersections based on Q Values 
Category Bolt Spacing 
(m) 
Bolt Length 
(m) 
Shotcrete 
Thickness 
(mm) 
Other 
1 1.7  2.1 4.0 50  90 Fibers 
2 1.8  2.3 4.0 40  90 Fibers 
3 1.9  2.3 4.0 40  90 none 
4 2.0  2.4 4.0 40  90 none 
5 2.2  2.7 4.0 none  50 none 
Table 6-7. Ground Support Needs for Large Intersections based on Q Values 
Category Bolt Spacing 
(m) 
Bolt Length 
(m) 
Shotcrete 
Thickness 
(mm) 
Other 
1 1.7  2.1 5.3 90  120 Fibers 
2 1.8  2.3 5.3 50  120 Fibers 
3 1.9  2.3 5.3 40  90 Fibers 
4 2.0  2.4 5.3 40  90 none 
5 2.2  2.7 5.3 40  50 none 

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6.5 ANALYTICAL METHODS 
The computer modeling techniques serve as an analytical basis for assessing the opening shape 
and determining the general stress distribution around the tunnel opening in underground design. 
In repository ground support design, numerical analyses will play an important role in 
determining and evaluating the effects of the seismic and thermal loading conditions. 
For the non-emplacement drift ground support analysis, a series of computer analyses are 
conducted and presented here to simulate the in situ, seismic, and/or thermal loading conditions. 
Note that in all numerical modeling cases the thermal loading is only applicable to exhaust mains 
and their intersections with emplacement drifts. An in situ thermal test has been conducting in 
the ESF to provide an experimental basis for evaluating the thermal-hydrological-
geomechanical-geochemical behavior at the site. The response of the host rock and ground 
support components to temperatures higher than the anticipated in the repository will provide 
knowledge about the response of the surrounding rock and the ground support components when 
subjected to high temperatures. 
In general, non-emplacement drift stability analyses include computer simulation of unsupported 
and supported openings. For unsupported openings, stability analysis includes the excavation of 
the non-emplacement drifts in a gravity-stressed rock medium, followed by seismic analysis. 
For stability analysis of exhaust mains and intersection between them and emplacement drifts, 
the introduction of elevated temperatures to simulate the waste emplacement process will follow 
in situ stress loading and seismic loading. For stability analyses of supported openings, the 
approach is the same as that for unsupported openings except that the final ground support 
system is incorporated into the modeling during the excavation process and, therefore, is 
subjected to in situ, thermal loading (only for intersections between exhaust mains and 
emplacement drifts, intersection between exhaust main and observation drift, and interburden 
area between access to intake shaft #1 and exhaust mains), and seismic loading. 
6.5.1 Modeling of Non-Emplacement Drifts 
6.5.1.1 Modeling of Access and Exhaust Mains and Turnouts 
For non-emplacement drifts such as access mains, exhaust mains, and turnouts, which have 
relatively large distance in longitudinal axis compared with its cross-sectional area, a two-
dimensional computer modeling is appropriate. Two-dimensional continuum approach will be 
used for the ground support design calculation for these openings. This is considered adequate 
and consistent with the conventional practice in mining or tunneling industry. 
In a continuum approach, the geologic features, such as fractures or lithophysae, in the rock mass 
are lumped into a thermomechanical constitutive model that represents the overall equivalent 
effect of these features. In a discontinuum approach, fractures or lithophysae are modeled 
explicitly as interfaces or cavities. The difference between these techniques is therefore the level 
of detail that is necessary in the model to adequately capture the deformation and failure 
mechanisms (Board 2003, Section 5.3.1). From a ground support design perspective, stability of 
non-emplacement drifts is judged by overall rock mass displacements and stresses. Two-
dimensional continuum approach that uses equivalent rock mass properties and constitutive 

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model may provide good tools for bounding analyses and also allow ease of parametric 
examination and model interpretation. Therefore, it is appropriate for use in ground support 
design related analyses. 
The FLAC computer code is employed in the two-dimensional analyses. In FLAC models, rock 
mass properties that reflect the effects of lithophysae and fractures on rock mass properties are 
used. These property values are presented in Tables 4-1 and 4-2 for the lithophysal and nonlithophysal 
rocks, respectively. The behavior of rock mass is judged using the Mohr-Coulomb 
yield criterion. Figure 6-9 illustrates the configuration of a FLAC model for access and exhaust 
mains. Figure 6-9A shows the FLAC model for turnout with rock bolts installed. To make the 
calculation more efficient, symmetry across the opening centerline is invoked, i.e., only half of 
the opening and the surrounding rock is modeled. The vertical dimension of the model is 100 m, 
and the horizontal is 50 m. Determination of the model dimension depends on requirements on 
accuracy and computational efforts. In general, the boundary effect is negligible if the model 
dimension is at least five times the size of an opening to be analyzed. 
6.5.1.2 Modeling of Intersections between Access Mains and Turnouts and between 
Exhaust Mains and Emplacement Drifts 
Because the geometry of the intersections is three-dimensional, the numerical analysis was 
conducted using the three-dimensional codes 3DEC and FLAC3D. Most of the analysis was 
carried out using 3DEC, which has the advantage of generating model geometry easily and 
quickly. Note that 3DEC is a discontinuum code, but in the simulations of intersections, the 
blocks were glued together to behave effectively as a continuum. Thus, the local stability of 
blocks created by joints around the excavation was not considered. It should also be noted that, 
because of the tetrahedral zones with constant strain used in 3DEC, the calculation of plastic 
deformation is sometimes inaccurate. The advantage of FLAC3D is that it performs more 
accurate calculations of plastic deformation, particularly in the case of softening plasticity. 
Because FLAC3D models for such complex geometries as those of the intersections are very 
time consuming to generate, only location C was analyzed with FLAC3D. The objective was to 
verify the 3DEC results. 
The geometries of two layouts A and C for intersections of access mains and turnouts as 
represented in the 3DEC model (showing the tunnels only, i.e., the surrounding rock mass is 
hidden) are shown in Figures 6-10 and 6-10A. The locations of the vertical section planes and 
profiles in which some of the results are presented are indicated in the figures. The geometry of 
the FLAC3D model for location C is shown in Figures 6-11 and 6-11A. The access main has a 
circular cross-section with a diameter of 7.62 m. The turnout has a horseshoe shape in the cross-
section of 8 m wide and 7 m high. The floors of the two tunnels are at the same elevation. The 
dimensions of the entire model for both locations are 200 m  200 m in plan and 100 m in 
height. The geometry of the entire 3DEC model for location A is shown in Figure 6-12. 
The geometries of the model for locations B and D for intersections of exhaust mains and the 
emplacement drifts are shown in Figures 6-13 and 6-13A. In both cases, the model size is 
200 m  81 m in plan and 100 m in height. The typical dimension for exhaust main is 7.62m 
except in Panel 1 of the repository, where it is 5.5 m. However, throughout the analyses it was 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-9. FLAC Model of Access and Exhaust Mains 
Figure 6-9A. FLAC Model of Turnouts with Rock Bolts 

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assumed that diameter of the exhaust main is 7.62 m, a conservative condition (resulting in 
larger spans) with regard to stability of the excavations. The symmetry condition, assumed on 
the vertical planes at mid-distance between the emplacement drifts, simplifies calculation of the 
thermal stresses. 
6.5.1.3 Modeling of Observation Drift 
Two geometrical dispositions of the observation drift were analyzed. A typical cross-section 
when the observation drift is parallel to the emplacement drift is analyzed using a two-
dimensional FLAC model. The geometry of the model and computational grid are shown in 
Figures 6-14 and 6-14A. The intersection of the observation drift with the exhaust main, 
accounting also for the interaction with the nearest emplacement drift, is analyzed in a three-
dimensional 3DEC model. The geometry of the observation drift in the intersection with exhaust 
main and the geometry of the 3DEC model are shown in Figures 6-15 and 6-15A. 
6.5.1.4 Modeling of TBM Launch Chamber 
The TBM launch chamber was analyzed using a two-dimensional FLAC model. Figures 6-16 
and 6-16A show the geometry and grid of the model in a cross-section normal to the chamber 
axis. 
6.5.1.5 Modeling of North Portal 
Figure 6-17 shows isometric and plan views of the 3DEC model of the North Portal. The model 
was constructed based on available topographical information and cross-sections of the portal 
region. Cross-sections C1-C1 and L-L, indicated in Figure 6-17, are shown in Figures 6-17Aa 
and 6-17Ab. 
6.5.1.6 Modeling of Interburden Pillar between Shaft Access and Exhaust Mains 
The location of the 3DEC model for interburden pillar between shaft access and exhaust mains is 
shown in location E in Figure 6-1. At this location, the access drift to #1 intake shaft overlies the 
two parallel exhaust mains. In plan view, the first drift intersects with the two exhaust mains. 
From three-dimensional viewpoint, these three drifts are not intersecting. The interburden 
between the first drift and the latter two drifts is about 10 m and the diameter of each of these 
drifts is 7.62 m. The model size is 200 m  81 m in plan and 100 m in height. Figures 6-17B 
and 6-17C show the configuration and geometry of the 3DEC model. Note that only 
unsupported drifts were modeled for this case. 
6.5.2 Loading and Boundary Conditions 
In designing the repository openings, stresses resulting from four sources must be considered: in 
situ (including excavation effects), construction and operation activities, thermal (nuclear waste), 
and seismic. In situ stresses are present before drift excavation and will be altered in the vicinity 
of openings during repository excavation. The stresses during construction, such as installation 
activities (e.g. jacking process) or stresses due to equipment movement such as TBM weight 
during mining, must be considered in the design of the ground support systems. The stresses due 

Title: Ground Control for Non-Emplacement Drifts for LA 
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to repository operations such as loads caused by gantry weight or waste package weight may also 
need to be considered in the design. Thermal stresses will occur after waste emplacement, and 
the timing and magnitude of the temperature-induced loads at any particular location in the 
repository are primarily dependent on the position relative to the stored waste packages. The 
magnitude of earthquake-induced stress and the duration of the earthquake event are a function 
of the intensity of the earthquake, the distance from the event to the repository, and the direction 
and size of the seismic wave relative to the opening. The applicability and magnitude of some of 
the design loads will vary depending on the type of ground support system. Some of the loads, 
such as thermal loads, will only apply to the final ground support system. In the following 
sections, a description for each design load type is presented. 
6.5.2.1 In Situ Loads 
The virgin stress field existing before excavation is the in situ or geostatic state of rock stress. 
Excavation of repository openings will disturb the surrounding in situ stress field. The stability 
of the opening will depend on the concentrations of excavation-induced stress and rock mass 
deformation behavior. For repository openings, computer simulation of the excavation process 
will be used to assess the stability. In situ stress estimates, opening dimensions, rock mass bulk 
density, rock mass elastic moduli, and rock mass strength parameters for the failure or yield 
criteria are required to perform the analyses. 
The in situ stress state at the repository has not been measured directly and will vary from 
location to location. For the initial state of stress, the vertical stress (s v) at some point is caused 
by the overburden weight and is given as 
n 
s- = .. gh (Eq. 6-5)
v ii 
i= 1 
where . i (kg/m3) is average bulk density of the ith layer of rock mass, hi (m) is thickness of the 
ith layer of rock mass above an opening, g (m/s2) is the gravitational acceleration, and n is the 
total number of overlaying layers of rock mass. 
Average initial horizontal stress (s h) at the same location is estimated as: 
s h = K0s (Eq. 6-6)
v 
where K0 = horizontal-to-vertical stress ratio, dimensionless. 
The vertical in-situ stress is gravitational, and it is the major principal stress. In this calculation, 
the average rock density of 2410 kg/m3 (Section 4.1.2) is assumed to estimate the overburden and 
in situ stress state for modeling all non-emplacement drifts except North Portal, in which a 
density of 2310 kg/m3 (see Table 4-7) is used. For 3DEC modeling, the major horizontal 
principal stress is 62% of the vertical stress whereas the minor horizontal principal stress is 36% 
of the vertical stress (Section 5.3). This is according to the in situ stress measurement by 
hydraulic fracturing in a test hole located in the TSw2 unit (SNF37100195002.001). For FLAC 
modeling, the horizontal-to-vertical stress ratio is 0.5 (Assumption 5.3). The overburden weight 
is applied as a stress boundary condition at the top of the model. 

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In evaluation of the stresses induced by excavation in the ground support components for the 
non-emplacement drifts, an initial ground relaxation value of 75 percent is used. This results in 
25 percent of the pre-excavation in situ stress being imposed on the ground support system 
(Section 5.1). This value is considered to be conservative to account for the effect of face 
advance and stand-off between the face and the installed rock bolts. Any rock deformation 
induced by the excavation will most likely be complete well before the installation of the final 
ground support. Also, as discussed in Assumption 5.1, an initial ground relaxation value of 100 
percent is assumed and used in the ground support analysis for the final shotcrete lining in non-
emplacement drifts. In other words, there will be no pre-excavation in situ stress being imposed 
on the shotcrete system. 

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Access Main 
Turnout 
Note: the distance between sections 1 and 2 is 6.2 m.
Figure 6-10. 3DEC Model Configuration of Intersection at Location A

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Access Main 
Turnout 
Note: the distance between sections 1 and 2 is 14.8 m.
Figure 6-10A. 3DEC Model Configuration of Intersection at Location C

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Figure 6-11. FLAC3D Model Configuration of Intersection at Location C 
Figure 6-11A. Perspective View of Intersection at Location C in FLAC3D Model 

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Figure 6-12. 3DEC Model Geometry and Dimension for Intersection at Location A 

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Drift 69o 
Exhaust Main 
Emplacement 
Figure 6-13. Geometry of 3DEC Model for Intersection at Location B 

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45o 
Exhaust Main 
Emplacement Drift 
Figure 6-13A. Geometry of 3DEC Model for Intersection at Location D 

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Figure 6-14. Geometry of Two-dimensional Model of Observation Drift and Emplacement Drift 
Figure 6-14A. Grid of Two-dimensional Model of Observation Drift and Emplacement Drift 

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Figure 6-15. Geometry of Intersection of Observation Drift with Exhaust Main 

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Figure 6-15A. Geometry of 3DEC Model for Intersection between Observation Drift and Exhaust Main 

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Figure 6-16. Geometry of Two-dimensional Model of TBM Launch Chamber 
Figure 6-16A. Grid of Two-dimensional Model of TBM Launch Chamber 

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Figure 6-17. a) Isometric and b) Plan Views of North Portal Model 

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Figure 6-17A. a) Frontal Cross-sectional (C1-C1) and b) Longitudinal Cross-sectional (L-L) Views of North 
Portal Model 

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69o 
10 m 
23 m 
Exhaust Main 
Access Drift 
Figure 6-17B. Configuration of 3DEC Model for Interburden Pillar between Access Drift to Intake Shaft #1 
and Exhaust Mains 

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Figure 6-17C. Geometry of 3DEC Model for Interburden Pillar between Access Drift to Intake Shaft #1 
and Exhaust Mains 

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6.5.2.2 Operation Loads 
Operational loads, such as waste package weight and invert material weight, are not considered 
in this calculation due to the preliminary nature of the design. Exclusion of these loads was 
believed to result in an overestimate of the inward rock displacements (particularly below the 
springline) since these loads are expected to offset some of the displacements caused by 
excavation and heating (note that heating is only considered for those areas in the proximity of 
emplacement drifts and exhaust mains or directly connected with them). 
6.5.2.3 Seismic Loads 
Ground motions associated with earthquakes are required to be considered in the design of the 
repository underground openings. The critical ground motions for subsurface design are ground 
velocity and acceleration. 
In contrast to surface structures, underground structures such as tunnels and their lining or 
reinforcement systems are constrained by the surrounding medium and do not move 
independently of the surrounding rock. In reality, the underground structures display 
significantly greater degrees of redundancy due to the confinement from the ground compared to 
surface structures, which are generally unsupported above their foundation. Therefore, for 
underground openings, the surrounding rock acts as a support during a seismic event. Case 
history studies of underground openings which have been subjected to seismic activities are 
generally used in practice to provide a basis for ground support design. Forty-one (out of 
seventy-one) cases of observed damage to rock tunnels from earthquake movements were 
compared to calculate peak surface motions to determine damage thresholds (Dowding 1979). 
The tunnels were built between late 1800s and the 1960s, and, thus, represent a wide variety of 
construction methods. It is shown that peak surface accelerations which cause heavy damage to 
surface structures, cause only minor damage to tunnels. For accelerations up to 0.19g and 
velocities up to 20 cm/s, no damage was experienced even for unsupported openings. Minor 
damage (new cracks and minimal rock fall) was observed for accelerations up to 0.5g and 
velocities up to about 90 cm/s. 
Other than case history studies, there are no empirical or closed-form solutions available to 
assess seismic effects on underground openings. Limited progress has been made in seismic 
design methodology for underground tunnels, possibly because of favorable performance data. 
The lack of applicable codes in the past has led to widely varied measures of precaution taken by 
different engineers. Moreover, the development of pertinent computer codes, as well as vast 
improvements in computational capacities of hardware, provides effective tools for seismic 
design of underground openings. For ground support analysis for LA, fully dynamic analyses 
will be performed to simulate seismic effects and assess opening stability. 
Dynamic loading was applied at the bottom of the model, propagating vertically upwards. 
Although the dynamic loading was specified as velocity time histories, it was numerically 
applied at the bottom model boundary as stress boundary condition by using formulas developed 
for plane waves in elasto-dynamics with the direct relation between velocity and stress given as 
follows (Itasca Consulting Group 2002, Manuals/3DEC/Optional Features/Section 2: Dynamic 
Analysis, Section 2.6): 

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s= 2.Cv 
s
y pv 
xy = 2.Cvh1 (Eq. 6-7) 
s 
s 
zy = 2.Cvh2
s 
where . is material density; Cp and Cs are P and S wave velocity, respectively; and .v and .h1and 
.h2 are vertical and horizontal velocity components. The factor 2 in Equation 6-7 is due to quiet 
boundaries. 
Seismic load used corresponds to a mean annual exceedance probability of 1x10-4 (10,000 years) 
and 5x10-4 (2,000 years) as mentioned in Section 4.1.5. Note that seismic load for a mean annual 
exceedance probability of 1x10-4 is used in all models except for North Portal, in which seismic 
load for a mean annual exceedance probability of 5x10-4 is used (the 1x10-4 ground motion for 
North Portal is not available). Seismic load is considered in both 3DEC and FLAC3D models by 
applying seismically-induced stresses to the lower boundary of a model. Both P- and S-waves 
are applied simultaneously. In order to reduce computer computational time, only portion of the 
velocity time histories that cover 5 to 95% of energy bracket is used. Note that below and 
beyond this bracket dynamic amplitudes are significantly low. The duration corresponding to 
this portion of velocity time histories for 10,000 and 2,000 year return periods are from 9.78 to 
58.79 seconds (DTN: MO0306SDSAVDTH.000) and 4.84 to 32.55 seconds (DTN: 
MO0402SDSTMHIS.004), respectively. 
6.5.2.4 Thermal Loads 
Thermally induced stresses are generated by thermal expansion of the rock mass due to the 
thermal energy released from the stored nuclear waste. Thermal stresses at any location depend 
on the proximity and timing of waste emplacement, the waste heat generation, the age of the 
waste, packaging and emplacement configuration, and the thermomechanical properties of the 
rock mass. 
Heat transfer process due to heat decay of the waste packages in the emplacement drift is a 
complicated thermal process. During the preclosure period, the heat transfer process is dictated 
by the thermal radiation from the waste package to the drift wall, convection due to the 
preclosure ventilation, and the thermal conduction of heat into the rock mass around the drift. 
The convective heat transfer occurs due to the temperature difference between the airflow and 
surfaces of the waste package and the drift wall. The temperature difference was originated by 
the thermal radiation from the waste package to drift wall. 
It should be noted that thermal loading is only considered in evaluating the stability analysis for 
intersections between exhaust mains and emplacement drifts, observation drift, intersection 
between observation drift and exhaust main, and interburden pillar between access to intake shaft 
#1 and exhaust mains. 

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The exhaust mains are located near (intersecting) the emplacement drifts. After waste 
emplacement, the exhaust mains and the intersections between the exhaust mains and the 
emplacement drifts will be subjected to the thermal stresses. The temperature change and thermal 
stresses will be different around the intersection at location B (in the middle of the repository) 
and at location D (at the edge of the repository). The evolution of the temperature field around 
the intersections (both locations B and D) was calculated by the NUFT code. Temperature 
changes at the edge of the repository (location D) are overpredicted by NUFTs two-dimensional 
results (because emplacement drift is assumed to be infinitely long in 2-D model whereas there is 
no emplacement drift on the other side of exhaust main at location D), resulting in conservative 
estimate of stress increase (and more unfavorable stability conditions). The temperature fields 
(at 1, 2, 3, 5, 10, 20, 30 and 50 years) from the NUFT code (BSC 2003g, Section 6.4.1) were 
imported into 3DEC in order to calculate the thermal stress around the excavations, according to 
the following formula (Itasca Consulting Group 2002, Manuals/3DEC/Optional Features/Section 
1: Thermal Option, Section 1.2.3): 
a (
.s ij =d ij 3KT ). T (Eq. 6-8) 
where .sij is the stress change, dij is the Kronecker delta, K is the bulk modulus of the rock 
() is the coefficient of the linear thermal expansion, a function of temperature (as listed mass, aT 
in Table 4-5), and .T is the temperature change. 
3DEC was used to model the effects of the thermal stress on stability and deformation of the 
intersections between emplacement drifts and exhaust mains. The thermal expansion of the 
ground support was not accounted for. In general, this approximation results in conservative 
estimate of the load in the rock bolts for the preclosure. For the purpose of this calculation, it is 
considered adequate. 
6.5.2.5 Boundary Conditions 
The boundary conditions for 3DEC computer analysis are presented in Table 6-8. At the in situ 
loading stage and the later thermal loading period, fixed velocity at lateral and bottom 
boundaries were used to ensure boundary effect does not affect the stress distribution around the 
opening. The prescribed temperatures generated from NUFT at various time periods were 
incorporated into 3DEC model to evaluate the thermomechanical behavior. For the dynamic 
analysis, non-reflecting boundary is used for both the top and bottom of the model, whereas free-
field boundary is imposed at the lateral perimeter of the model. The free-field boundaries ensure 
that plane waves propagating upward suffer no distortion at the boundary. At the dynamic 
loading stage, the seismic load is considered in the 3DEC models by applying seismically-
induced stresses (both normal and shear tractions) to the lower boundary of a model. FLAC and 
FLAC3D are used for static analysis in this calculation. Their boundary conditions are the same 
as those for 3DEC model. 

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Table 6-8. Boundary Conditions for 3DEC Analysis 
Boundary Static Analysis Stage Dynamic Analysis Stage 
Lateral Fixed at the direction 
normal to the face Free-Field boundary 
Bottom Fixed at the vertical 
direction Non-reflecting boundary 
Top Applied pressure at the 
vertical direction Non-reflecting boundary 
Drift Wall Free Non-reflecting boundary 
6.5.3 Stability of Unsupported Non-Emplacement Drifts 
This section presents the results of assessment of stability of unsupported non-emplacement 
drifts. The assessment is based on numerical analysis using the FLAC, 3DEC, and FLAC3D 
computer codes. The analysis evaluates displacement and stress in the vicinity of an 
unsupported non-emplacement drift, factor of safety, ground reaction curves (GRC), and 
temperature increases in rock following waste emplacement. 
6.5.3.1 In Situ Stress Loading Condition 
The effect of the rock mass quality variability within the lithophysal and non-lithophysal units at 
the repository level was accounted for by considering different rock mass categories. There are 
six categories in the lithophysal unit and five categories in the non-lithophysal units. The rock 
mass properties for different categories listed in Tables 4-1 and 4-2 for the lithophysal and nonlithophysal 
rock mass, respectively, are determined by rock mass characterization based on 
laboratory and in-situ testing. In this calculation, the computer simulations for all non-
emplacement openings except for North Portal were mainly carried out for the most conservative 
conditions, category 1 rock (the poorest quality rock mass), both in lithophysal and nonlithophysal 
units. Stability analysis for category 5 rock (the strongest quality rock mass) was 
also included for access and exhaust mains. Category 6 rock is considered representative of 
extremely poor quality lithophysal rock mass in the disturbed rock zone around excavations. 
Therefore, it was not considered in the simulations because it would result in excessively 
conservative, unrealistic predictions of deformation and damage of the rock mass. A Mohr-
Coulomb, elastic-plastic constitutive model was assumed for all simulations. It should be noted 
that in computer simulations for North Portal the properties considered for the rock are those of 
the poorest quality rock mass (i.e., category 1) in the TCw unit (see Table 4-7). 
6.5.3.1.1 Access and Exhaust Mains and Turnouts 
Table 6-9 summarizes the results of unsupported access and exhaust mains and turnouts for drift 
displacement and rock stresses induced by excavation. For access and exhaust mains, the 
horizontal displacement varies from 6.4 mm for lithophysal category 1 rock to 0.3 mm for non

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lithophysal category 5 rock. The vertical displacement is estimated to range from 32.8 mm for 
lithophysal category 1 rock to 2.3 mm for non-lithophysal category 5 rock. For turnouts, the 
horizontal displacement for lithophysal category 1 rock is 9.0 mm and 1.3 mm for nonlithophysal 
category 1 rock. The vertical displacement for lithophysal category 1 rock is 36.2 
mm and 6.5 mm for non-lithophysal category 1 rock. Note that the sign convention in FLAC is 
that positive displacement signifies displacement in positive X or Y direction. Thus, the above 
mentioned vertical and horizontal displacement are moving in downward and inward directions, 
respectively. Also, all the maximum vertical displacements occur at the crown whereas the 
maximum horizontal displacements at springline. Figure 6-18 shows the displacements around 
access mains for lithophysal category 1 rock. The largest displacement of 32.8 mm occurs at the 
crown. The results show the effect of rock mass quality on the drift deformation. 
The maximum tangential stress for access and exhaust mains varies from about 20 MPa in 
compression occurring near the skin of springline to about 5 MPa in compression occurring near 
the skin of crown. The maximum tangential stress for turnouts varies from 18 MPa in 
compression for non-lithophysal category 1 rock occurring near the skin of springline to about 5 
MPa in compression occurring near the skin of crown. The stresses at the crown are much 
smaller than those at the springline. The minor principal (radial) stresses are about 1 MPa or 
lower, which are much lower than those of the maximum tangential stresses. 
Figures 6-19 and 6-20 show the contours of safety factor (SF) against Mohr-Coulomb shear 
failure around the unsupported access/exhaust mains for category 1 rock in lithophysal and nonlithophysal 
units, respectively. The SF in this case is defined as the ratio of shear strength to the 
shear stress on the potential failure plane. For lithophysal rock, a very narrow zone adjacent to 
wall (about 0.2 m into wall) is shown with potential yield. At distance of 1.5 m into wall SF 
increases to 2. For non-lithophysal rock, almost no yield is observed near the wall and SF 
increases to 2 with a very short distance (about 0.2 m) into wall. Figures 6-21 and 6-22 show the 
contours of safety factors around the unsupported turnouts for category 1 rock in lithophysal and 
non-lithophysal units, respectively. The distribution of SF contours is similar to those of access 
mains except the extent is a little larger, which is mainly due to the stress concentration at 90o 
angle intersection between sidewall and invert. There is no indication of instability of the rock 
surrounding the access/exhaust mains and turnouts. 
6.5.3.1.2 Intersections 
6.5.3.1.2.1 Intersection between Access Main and Turnout 
Stability of the intersections in the lithophysal and non-lithophysal rock mass for in-situ stress 
conditions, assuming no ground support, is analyzed for two layouts at locations A and C. The 
intersection at location C represents a typical layout; the intersection at location A represents the 
critical conditions with respect to the maximum span of the roof and stress concentrations in the 
pillars. 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Table 6-9. Unsupported Access and Exhaust Mains and Turnouts - Results for In Situ Stress Loading 
Condition from FLAC 
Items 
Access/Exhaust Mains Access/Exhaust Mains Turnouts 
L. Cat. 1 NL. Cat. 1 L. Cat. 5 NL. Cat. 5 L. Cat. 1 NL. Cat. 1 
Horizontal Displacement (mm) -6.4 -0.8 -0.5 -0.3 -9.0 -1.3 
Vertical Displacement (mm) -32.8 -5.9 -3.1 -2.3 -36.2 -6.5 
Major Principal (Tangential) Stress 
near Crown (MPa) -5.3 -5.3 -5.3 -5.3 -5.3 -5.3 
Minor Principal (Radial) Stress 
near Crown (MPa) -0.5 -0.4 -0.4 -0.4 -0.2 -0.2 
Major Principal (Tangential) Stress 
near Springline (MPa) -19.0 -21.5 -21.5 -21.5 -11.0 -18.0 
Minor Principal (Radial) Stress at 
Springline (MPa) -1.0 -1.2 -1.2 -1.2 -0.2 -0.3 
Note: a) L  lithophysal, NL  non-lithophysal, b) The vertical displacement occurs at crown whereas horizontal 
displacement at springline. 
To avoid large, unphysical deformation of zones around the boundary of the excavation that 
yield in tension, tensile strength in the 3DEC simulations was assumed to be infinitely large. 
Simulation of the intersection at location C in the lithophysal rock mass category 1 was 
conducted with both 3DEC and FLAC3D. In the FLAC3D simulation the rock mass was 
assumed to have no tensile strength. The results of 3DEC (with infinite rock mass tensile 
strength) and FLAC3D (with no tensile strength) are compared to demonstrate that the 
assumption used in the 3DEC model has little effect on the results (see Figure 6-37 and 
discussion in Section 6.5.3.1.2.1.1). 
The results for the intersection at the location A (hereinafter called intersection A) and at the 
location C (hereinafter called intersection C) are shown in Figures 6-23 through 6-56. For both 
units (lithophysal and non-lithophysal) and both intersection locations (A and C), the figures 
include the plots of: 
1. Displacement fields in two vertical sections 1 and 2 (see section location in Figures 6-11 
and 6-12); 
2. Displacement profile along the vertical line though the point A (see Point A location in 
Figures 6-11 and 6-12); 
3. Stress fields in the same vertical sections as for the displacement fields; 
4. Average pillar stress as a function of the distance from the tip (Point A). The stresses 
were averaged over the entire pillar width and height for a 2-m long segments measured 
radially from the pillar tip; and 
5. Potential yield zones in the horizontal and the vertical sections. 

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Figure 6-18. Displacements at Access and Exhaust Mains for Lithophysal Cat. 1 Rock 
6.5.3.1.2.1.1 Rock Displacements 
Figures 6-23 and 6-24 show displacement fields in vertical section 1 at intersection A for 
category 1 rock of lithophysal and non-lithophysal rock, respectively. For the same intersection, 
Figures 6-25 and 6-26 show displacement fields in vertical section 2 for category 1 rock of 
lithophysal and non-lithophysal rock, respectively. Note that vertical section 1 cuts across two 
turnouts and one access main as well as two very narrow pillars. Vertical section 2 cuts through 
the intersecting access main and turnout, one narrow pillar, and another turnout. The important 
features of these two sections are: section 1 cuts through very narrow pillar whereas section 2 
cuts through an opening with large roof span. The results show that the maximum displacement 
in lithophysal rock occurs at the intersection area with large roof span with a value of 7.21 cm. 
The maximum displacement in non-lithophysal rock at the same location is 1.12 cm. Figure 627 
shows the vertical displacements along the vertical line through point A in section 2. The 
profiles of the vertical displacement along the vertical line through point A show a gradual 
increase in displacements approaching the back of the intersection. There is no indication of 
large displacement gradients, which is a consequence of distressing and loosening of potentially 
unstable ground. As shown in this figure, the displacements for lithophysal category 5 rock are 
smaller than those of non-lithophysal category 1 rock. The major reason is that the elastic 
modulus of the former is larger than that of the latter. Relatively large displacements for the 
lithophysal category 1 rock mass result from the very low elastic modulus assigned to the 
lithophysal category 1 rock mass. These observations are confirmed with plots of the rock mass 

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regions that deformed plastically, as shown in Figures 6-28 and 6-29. Plastic deformation 
extends to approximately 1 m from the drift walls in the lithophysal, category 1 rock mass. There 
is practically no plastic deformation around the drifts in the non-lithophysal rock mass. 
Similar results are predicted for intersection C as shown in Figures 6-30 through 6-36. However, 
since each access main only intersects with one turnout and the roof span at intersection C is 
smaller, the corresponding opening displacements are smaller compared with those of 
intersection A. For instance, the maximum displacements in lithophysal and non-lithophysal 
rock occur at section 2 (i.e., with large roof span) and are 6.4 and 1 cm, respectively. The 
maximum displacements in lithophysal and non-lithophysal rock occur at section 1 are 4.8 and 
0.8 cm, respectively. Figure 6-37a shows the contour plots of vertical displacements at section 1 
for lithophysal category 1 rock based on FLAC3D results. The maximum displacements at the 
crown area range from 4.0 to 5.2 cm and are comparable to that of 3DEC results. (Note that 
Figure 6-37b shows the contour plots of stress field at section 1 for lithophysal category 1 rock 
based on FLAC3D results. The range of stress fields and distribution in this figure is comparable 
to that of Figure 6-48.) The results justify the use of 3DEC and the assumption of infinite tensile 
strength in the analyses of deformation and stability of the intersections (also see discussion in 
Section 6.5.3.1.2.1). 
6.5.3.1.2.1.2 Stresses in Rock Adjacent to Openings 
Figures 6-38 and 6-39 show stress fields in vertical section 1 at intersection A for category 1 
rock of lithophysal and non-lithophysal rock, respectively. For the same intersection, Figures 640 
and 6-41 show stress fields in vertical section 2 for category 1 rock of lithophysal and nonlithophysal 
rock, respectively. The results indicate that the maximum compressive stress occurs 
near the sidewall or in the pillar between adjacent openings and slight destressing above crown 
and below invert. The highest compressive stress of 56 MPa occurs near the sidewall of the 
opening at intersection area with large roof span in non-lithophysal category 1 rock. Note that 
this stress value is larger than the unconfined compressive strength of the rock mass (see Table 
4-2). This result does not mean that rock will fail at that location since the rock mass in situ is 
under confined condition. Yet, it does indicate that there is a potential for rock spalling near the 
sidewall, which is also revealed in Figures 6-42 and 6-43. In these two figures, the contours of 
safety factors ranging from 1 to 1.6 are observed along the sidewalls of the openings, especially 
more evident for lithophysal rock. Figure 6-44 shows the stress field with indication of tensile 
stresses in vertical section 2 for lithophysal category 1 rock. A very narrow region of tensile 
stress is observed in the crown and the invert area. There are no indications of large-scale 
instability of the roof even in category 1, lithophysal rock mass. It is possible that a limited 
volume of the rock mass becomes unstable due to local structures unaccounted for in this 
analysis (continuum approach) if the intersections are unsupported. The designed ground support 
in the intersection will be needed to prevent any such rockfall. 
Stability of the pillar between the access main and the turnout can be assessed based on a plot of 
average pillar stress (shown in Figure 6-45) and plots of regions of plastic deformation (i.e., 
potential yield zone) shown in the horizontal section in Figures 6-46 and 6-47 (note that the 
similar plots in vertical section are shown in Figures 6-28 and 6-29). The stress concentrations 
will cause damage and fracturing of the pillar. The extent of the damage will depend on the 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
quality of the rock mass. The calculations predict pillar damage to extend 8 m from the tip in the 
lithophysal rock mass, category 1. In good quality lithophysal rock mass and non-lithophysal 
rock mass the pillar will be damaged to a distance of approximately 3 m from the tip. However, 
in reality, the pillars will be rounded to a distance from the tip where the major damage is 
observed. It is not needed (and probably impossible) to create the pillar tip. If rounding of the 
pillars is not performed during excavation, it will take place naturally by broken ground falling 
off. The pillars will eventually attain the stable geometry. It seems, from the numerical results, 
that the rounding length should not be larger than 5 m in the poorest quality lithophysal rock 
mass. If the additional ground support is used the pillar rounding length could be reduced even 
more. 
Similar results are predicted for intersection C as shown in Figures 6-48 through 6-56. However, 
since each access main only intersects with one turnout and the roof span at intersection area is 
smaller, the corresponding rock stresses are smaller compared with those of intersection A. 
There are no indications of large-scale instability of the roof even in category 1, lithophysal rock 
mass. It is possible that a limited volume of the rock mass becomes unstable due to local 
structures unaccounted for in this analysis (continuum approach) if the intersections are 
unsupported. The designed ground support in the intersection will be needed to prevent any such 
rockfall. The pillar stability for intersection C should be achieved as long as the pillar tip is 
going to be rounded off as discussed above for intersection A. 
6.5.3.1.2.1.3 Discussion 
The 3DEC models predict similar deformation and stability conditions for intersections at 
locations A and C. The rock mass deformation around intersections will be predominantly 
elastic. The plastic deformation is expected to extend approximately 1 m from the drift walls into 
the rock mass, in the worst case, for the poorest quality rock mass. A relatively large 
displacement of 7 cm predicted for the lithophysal, category 1 rock mass (see Figure 6-25), is a 
consequence of the very low Youngs modulus used. There are no indications of loosening and 
instability of the rock mass in the crown for any of the analyzed cases. The roof in the 
intersections appears to be stable, even for the largest spans. 
The tips of the pillars between the access main and the turnouts will be damaged due to stress 
concentrations. The extent of this damage from the pillar tip will depend on the rock mass 
quality. Rounding of the pillars and additional ground support will resolve the problem of pillar 
instability. 
6.5.3.1.2.2 Intersection between Exhaust Main and Emplacement Drift 
Two intersection layouts, B and D, were analyzed as being representative of the extreme 
conditions existing in the exhaust main intersections from the perspective of excavation stability. 
The diameter of the exhaust main is 7.62 m  except in Panel 1 of the repository, where it is 
5.5 m. Throughout the analyses presented in this calculation, it was assumed that diameter of the 
exhaust main is 7.62 m, a conservative condition (resulting in larger spans) with regard to 
stability of the excavations. 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
The results of the stability analysis of unsupported intersections at locations B and D are shown 
in Figures 6-57 through 6-66. The analysis was carried out for categories 1 and 5 in both 
lithophysal (location B) and non-lithophysal (location D) rock masses, representing the 
variability of rock mass quality at the repository horizon. The results are presented here for 
category 1 only. The rock mass around the intersections in category 5 deforms elastically at both 
locations. The modeling sequence followed the expected construction sequence: the exhaust 
mains were excavated prior to excavation of the emplacement drifts. The presented results are 
for the final stage of the excavation. 
The maximum vertical displacement in the roof is 4.3 cm in the poorest-quality lithophysal rock 
mass (Figure 6-57). The maximum displacements in the non-lithophysal rock mass are predicted 
to be less than 1 cm (Figure 6-58 for all categories. The intersections cause stress concentrations 
in excess of 25 MPa (see Figures 6-59 and 6-60) in the walls of the exhaust main. However, 
stress tensors colored by the factor-of-safety with respect to a Mohr-Coulomb yield condition 
(shown in Figures 6-61 and 6-62), and the extent of the rock mass that undergoes plastic 
deformation in the vertical and horizontal sections (shown in Figures 6-63 through 6-66) indicate 
that deformation of the rock mass around the intersections will be predominantly elastic. There is 
a narrow (less than a meter) region of the rock mass around the exhaust main at location B for 
category 1 that yields (Figure 6-63). However, the effect of the interaction between the exhaust 
main and the emplacement drift on the extent of rock mass yielding is negligible. The sharp 
angle between the exhaust main and the emplacement drift creates a pillar (Figure 6-66) between 
these two excavations. At the most, the rock mass yields 1 m from the tip of the pillar, indicating 
that rounding the pillar tip would resolve any potential problems. 
6.5.3.1.3 Observation Drift 
The model results for the typical cross-section perpendicular to the observation drift axis are 
shown in Figures 6-66A to 6-66D. The maximum displacements predicted for the observation 
drift are similar to those for the emplacement drift (for the same rock-mass quality), of the order 
of 0.03 m, because the sizes of both tunnels are similar. The plot of the contours of the major 
principal stress (Figure 6-66B) clearly shows that, for a given distance, two excavations (the 
emplacement drift and the observation drift) do not interact with each other. The predicted extent 
of the damage of the rock mass (Figure 6-66C) surrounding the observation drift and the 
emplacement drift is similar. There is a little more damage in the walls and the floor of the 
observation drift because of mechanically less-favorable shape (i.e., flat walls and corners). 
However, the factor-of-safety with respect to Mohr-Coulomb shear failure (Figure 6-66D) 
increases very quickly as a function of distance from the drift boundary; the factor-of-safety is 
larger than 2 only a couple of meters from the drift walls. 
The results of modeling the intersection between the observation drift and the exhaust main are 
presented in Figures 6-66E to 6-66J. The maximum displacements (Figures 6-66E and 6-66F) are 
comparable to predictions of the two-dimensional model because of the larger span of the 
exhaust main (7.62 m). However, for the analyzed conditions, there are no indications of stability 
problems in the intersection. The regions of the rock mass that undergo plastic deformation are 
confined to the vicinity of the excavations (Figures 6-66G and 6-66H). The extent of damage 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
does not increase in the intersection, and it does not seem that the excavations interact with each 
other. The pillar between the observation drift and the exhaust main appears (in the vertical 
section 2) to be almost completely elastic. Factor-of-safety with respect to Mohr-Coulomb shear 
failure is larger than 2 (Figures 6-66I and 6-66J) except in close vicinity of the excavations. 
6.5.3.1.4 TBM Launch Chamber 
The larger span of the TBM launch chamber (11 m) compared to other excavation considered 
(i.e., emplacement drifts, access mains) results in larger displacements due to excavation and 
relaxation of the in situ stresses. The maximum displacement is 0.055 m (Figure 6-66K). The 
stress state around the chamber (Figure 6-66L) is similar to the stress state around the 
observation drift (Figure 6-66B), because two excavations have the same shape. Consequently, 
the size of the region of the rock mass undergoing inelastic deformation (Figure 6-66M), when 
scaled with characteristic dimension of the excavation, is the same for the launch chamber as for 
the observation drift. Because the span (or height) of the launch chamber is approximately two 
times of the span of the observation drift, the depth of the damaged rock in the wall of the launch 
chamber is about 2 m, compared to the depth of about 1 m predicted in the wall of the 
observation drift. Outside the yielded region, the factor-of-safety to shear failure increases 
rapidly as a function of distance from the drift wall (Figure 6-66N), indicating overall stability of 
the excavation. 
6.5.3.1.5 North Portal 
The model results under static in situ conditions after both the slope and the starter tunnel are 
excavated are shown in Figures 6-66O to 6-66Q. The stress state is completely elastic. Small 
overburden and slope heights result in stresses (Figure 6-66P) that are small compared to 
cohesion, even in the case of the poorest-quality rock mass, category 1, which was considered in 
the calculation. The factor-of-safety with respect to Mohr-Coulomb shear failure is quite large 
(approximately 5) throughout the entire model (see Figure 6-66O). The displacement field shown 
in Figure 6-66Q is due to the excavation of the starter tunnel only. The magnitude of the 
displacements is very small, of the order of 0.002 m or less. It appears from the results of the 
numerical modeling that the North Portal and the starter tunnel would be stable for the 
considered mechanical properties of the rock mass even if no ground support was used. 
6.5.3.1.6 Interburden Pillar between Shaft Access and Exhaust Mains 
The results of the interburden pillar stability analysis are shown in Figures 6-66R through 6-66Y. 
The results are for the final stage of the excavation. The maximum vertical displacement in the 
exhaust mains (crown and floor) is approximately 3.0 cm. In the access drift, the maximum 
vertical displacement of 4.3 cm (Figures 6-66R and 6-66S) occurs in the crown of drift. 
The excavation of drifts causes maximum stress concentrations of approximately 25 and 23 MPa 
(see Figures 6-66T and 6-66U) in the walls of the exhaust mains and access drift, respectively. 
However, safety factors with respect to the Mohr-Coulomb yield condition averaging 1 to 3 are 
shown for the rock immediately surrounding the drifts (shown in Figures 6-66V and 6-66W). In 
addition, a minor extent of the rock mass yielding is shown in Figures 6-66X and 6-66Y. 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Therefore, the results indicate that deformation of the rock mass around the drifts and in the 
pillar between the drifts will be predominantly elastic. 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-19. Contours of Safety Factor around Unsupported Access/Exhaust Main for Lith. Cat. 1 Rock 
Figure 6-20. Contours of Safety Factor around Unsupported Access/Exhaust Main for N. Lith. Cat. 1 
Rock 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-21. Contours of Safety Factor around Unsupported Turnout for Lith. Cat. 1 Rock 
Figure 6-22. Contours of Safety Factor around Unsupported Turnout for N. Lith. Cat. 1 Rock 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-23. Intersection A: Displacement Field in Vertical Section 1 for Lith. Cat. 1 Rock 
Figure 6-24. Intersection A: Displacement Field in Vertical Section 1 for N. Lith. Cat. 1 Rock 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-25. Intersection A: Displacement Field in Vertical Section 2 for Lith. Cat. 1 Rock 
Figure 6-26. Intersection A: Displacement Field in Vertical Section 2 for N. Lith. Cat. 1 Rock 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-27. Intersection A: Vertical Displacement along Vertical Line through Point A 
Figure 6-28. Intersection A: Potential Yield Zone in Vertical Section 2 for Lith. Cat. 1 Rock 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-29. Intersection A: Potential Yield Zone in Vertical Section 2 for N. Lith. Cat. 1 Rock 
Figure 6-30. Intersection C: Displacement Field in Vertical Section 1 for Lith. Cat. 1 Rock 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-31. Intersection C: Displacement Field in Vertical Section 1 for N. Lith. Cat. 1 Rock 
Figure 6-32. Intersection C: Displacement Field in Vertical Section 2 for Lith. Cat. 1 Rock 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-33. Intersection C: Displacement Field in Vertical Section 2 for N. Lith. Cat. 1 Rock 
Figure 6-34. Intersection C: Vertical Displacement along Vertical Line through Point A 

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Figure 6-35. Intersection C: Potential Yield Zone in Vertical Section 1 for Lith. Cat. 1 Rock 
Figure 6-36. Intersection C: Potential Yield Zone in Vertical Section 1 for N. Lith. Cat. 1 Rock 

Title: Ground Control for Non-Emplacement Drifts for LA 
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(a) Vertical Displacements
(b) Stress Fields 
Figure 6-37. Intersection C: Contours of Vertical Displacements and Stress Fields in Vertical Section 1 for 
Lith. Cat. 1 Rock Using FLAC3D 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-38. Intersection A: Stress Field in Vertical Section 1 for Lith. Cat. 1 Rock 
Figure 6-39. Intersection A: Stress Field in Vertical Section 1 for N. Lith. Cat. 1 Rock 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-40. Intersection A: Stress Field in Vertical Section 2 for Lith. Cat. 1 Rock 
Figure 6-41. Intersection A: Stress Field in Vertical Section 2 for N. Lith. Cat. 1 Rock 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-42. Intersection A: Factor of Safety in Vertical Section 2 for Lith. Cat. 1 Rock 
Figure 6-43. Intersection A: Factor of Safety in Vertical Section 2 for N. Lith. Cat. 1 Rock 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-44. Intersection A: Compressive and Tensile Stresses in Vertical Section 2 for Lith. Cat. 1 Rock 
Figure 6-45. Intersection A: Average Pillar Stress vs. Distance from the Pillar Tip 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-46. Intersection A: Potential Yield Zone in Horizontal Section for Lith. Cat. 1 Rock 
Figure 6-47. Intersection A: Potential Yield Zone in Horizontal Section for N. Lith. Cat. 1 Rock 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-48. Intersection C: Stress Field in Vertical Section 1 for Lith. Cat. 1 Rock 
Figure 6-49. Intersection C: Stress Field in Vertical Section 1 for N. Lith. Cat. 1 Rock 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-50. Intersection C: Stress Field in Vertical Section 2 for Lith. Cat. 1 Rock 
Figure 6-51. Intersection C: Stress Field in Vertical Section 2 for N. Lith. Cat. 1 Rock 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-52. Intersection C: Factor of Safety in Vertical Section 2 for Lith. Cat. 1 Rock 
Figure 6-53. Intersection C: Factor of Safety in Vertical Section 2 for N. Lith. Cat. 1 Rock 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-54. Intersection C: Average Pillar Stress vs. Distance from Pillar Tip 
Figure 6-55. Intersection C: Potential Yield Zone in Horizontal Section for Lith. Cat. 1 Rock 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-56. Intersection C: Potential Yield Zone in Horizontal Section for N. Lith. Cat. 1 Rock 
Figure 6-57. Intersection B: Displacement Field in Vertical Section 1 for Lith. Cat. 1 Rock 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
Figure 6-58. Intersection D: Displacement Field in Vertical Section 1 for N. Lith. Cat. 1 Rock 
Figure 6-59. Intersection B: Stress Field in Vertical Section 1 for Lith. Cat. 1 Rock 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-60. Intersection D: Stress Field in Vertical Section 1 for N. Lith. Cat. 1 Rock 
Figure 6-61. Intersection B: Factor of Safety in Vertical Section 1 for Lith. Cat. 1 Rock 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-62. Intersection D: Factor of Safety in Vertical Section 1 for N. Lith. Cat. 1 Rock 
Figure 6-63. Intersection B: Potential Yield Zone in Vertical Section 1 for Lith. Cat. 1 Rock 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-64. Intersection D: Potential Yield Zone in Vertical Section 1 for N. Lith. Cat. 1 Rock 
Figure 6-65. Intersection B: Potential Yield Zone in Horizontal Section for Lith. Cat. 1 Rock 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-66. Intersection D: Potential Yield Zone in Horizontal Section for N. Lith. Cat. 1 Rock 
Figure 6-66A. Displacements due to Excavations of Observation Drift and Emplacement Drift 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-66B. Contours of Major Principal Stress after Excavation of Observation Drift and Emplacement 
Drift 
Figure 6-66C. Potential Yield Zone after Excavation of Observation Drift and Emplacement Drift 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-66D. Factor of Safety after Excavation of Observation Drift and Emplacement Drift 
Figure 6-66E. Displacement Field in Vertical Section 1 at Intersection between Observation Drift and 
Exhaust Main 

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Figure 6-66F. Displacement Field in Vertical Section 2 at Intersection between Observation Drift and 
Exhaust Main 
Figure 6-66G. Potential Yield Zone in Intersection between Observation Drift and Exhaust Main in 
Section 1 

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Figure 6-66H. Potential Yield Zone in Intersection between Observation Drift and Exhaust Main in 
Section 2 
Figure 6-66I. Factor of Safety in Intersection between Observation Drift and Exhaust Main in Section 1 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-66J. Factor of Safety in Intersection between Observation Drift and Exhaust Main in Section 2
Figure 6-66K. Displacements due to Excavation of TBM Launch Chamber 

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Figure 6-66L. Contours of Major Principal Stress after Excavation of TBM Launch Chamber 
Figure 6-66M. Potential Yield Zone after Excavation of TBM Launch Chamber 

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Figure 6-66N. Factor of Safety after Excavation of TBM Launch Chamber 

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Figure 6-66O. Factor of Safety after Excavation of North Portal Starter Tunnel: a) Horizontal Cross-
section at Elevation 1128 m, and b) Longitudinal Cross-section L-L 

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Figure 6-66P. Major Principal Stresses after Excavation of North Portal Starter Tunnel for Cross-sections: 
a) C1-C1 and b) C2-C2 

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Figure 6-66Q. Displacement Field after Excavation of North Portal Starter Tunnel for Cross-sections: 
a) C1-C1 and b) C2-C2 

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Figure 6-66R. Displacement Field in Vertical Section 1 for Lith. Cat. 1 Rock at Interburden Area 
Figure 6-66S. Displacement Field in Vertical Section 2 for Lith. Cat. 1 Rock at Interburden Area 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-66T. Stress Field in Vertical Section 1 for Lith. Cat. 1 Rock at Interburden Area 
Figure 6-66U. Stress Field in Vertical Section 2 for Lith. Cat. 1 Rock at Interburden Area 

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Figure 6-66V. Factor of Safety in Vertical Section 1 for Lith. Cat. 1 Rock at Interburden Area 
Figure 6-66W. Factor of Safety in Vertical Section 2 for Lith. Cat. 1 Rock at Interburden Area 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-66X. Potential Yield Zone in Vertical Section 1 for Lith. Cat. 1 Rock at Interburden Area 
Figure 6-66Y. Potential Yield Zone in Vertical Section 2 for Lith. Cat. 1 Rock at Interburden Area 

Title: Ground Control for Non-Emplacement Drifts for LA 
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6.5.3.2 Thermal Loading Condition 
It should be noted the thermal loading condition is only applied to intersections of exhaust mains 
with emplacement drifts, observation drift, intersection of observation drift with exhaust main, 
and interburden pillar between access to intake shaft #1 and exhaust mains. The discussion in 
this section is limited to these areas. 
6.5.3.2.1 Intersections of Exhaust Mains and Emplacement Drifts 
The temperature fields in intersection B, as calculated in NUFT and imported into 3DEC, after 3 
and 20 years of heating are shown in Figures 6-67 and 6-68. The temperature histories at points 
1, 2 and 3 (indicated in Figures 6-13 and 6-14) are shown in Figures 6-69 and 6-70 for locations 
B and D, respectively. Note that the temperatures shown in these two figures are reasonable 
except temperatures at point 2 in Figure 6-69. The reason that temperatures at this point are 
higher than those at point 3 is because of the two-dimensional temperature field used in the 
calculation. In the three-dimensional model, the emplacement drift is not continuous, instead, it 
intersects and ends at the exhaust main. Consequently, some of the points on the wall of the 
exhaust main which are located within the projected outline of the emplacement drift have higher 
temperatures than those of the points on the emplacement drift. However, the impact of this 
result is insignificant. As indicated in Figure 6-71, thermally induced variation of major 
principal stress in the preclosure period is relatively small, i.e., about few percent with respect to 
the initial stress state before heating began. 
The thermal stress changes, shown (at points 1, 2 and 3 at locations B and D) in Figures 6-71 
through 6-74, are much larger in better quality and stiffer rock masses. The increase in the 
maximum major principal stress is about 15 MPa in category 5 of the lithophysal and nonlithophysal 
rock masses. However, the stress change does not change factor-of-safety with 
respect to the Mohr-Coulomb yield condition throughout the rock mass (Figures 6-75 through 678) 
nor cause significant additional yielding of the rock mass (comparing Figures 6-79 and 6-80 
with Figures 6-63 and 6-64). The extent of plastic deformation for both locations, and for the 
different rock-mass categories remains practically unchanged after 50 years of heating. The 
heating causes movement of the entire model vertically upward. The change in displacement in 
the vertical direction of the excavation in the intersection (at point 1) due to temperature change 
is not significant (Figures 6-81 and 6-82). For locations B and D, it appears that displacement in 
the vertical direction at point 1 is reduced. 
6.5.3.2.2 Observation Drift 
The temperature field after 50 years of heating around the observation drift is shown in Figure 682A. 
Temperature histories during 50 years of heating at three points around the observation 
drift are shown in Figure 6-82B. The observation drift does not significantly affect heat transport 
or the predicted temperature fields. This simplification is conservative, resulting in 
overestimation of the temperatures around the observation drift. 
It is predicted that the maximum temperature increase around the observation drift during the 50yr 
period after waste emplacement is approximately 13oC. It does not appear that this 
temperature change (and the corresponding stress change shown in Figure 6-82C) causes any 

Title: Ground Control for Non-Emplacement Drifts for LA 
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additional damage around the observation drift (Figures 6-82D and 6-82E). Induced 
displacements of the points on the drift boundary are of the order of millimeters (Figure 6-82F). 
Figures 6-82G to 6-82J indicate that thermal loads do not affect the stability conditions of the 
rock mass around the intersection between the observation drift and the exhaust main. 
6.5.3.2.3 Interburden Pillar between Shaft Access and Exhaust Mains 
After waste emplacement, the exhaust mains and the pillar between the exhaust mains and the 
access drifts drift to the ventilation shaft will be subjected to the thermal stresses. The evolution 
of the temperature field around the pillar was calculated by the NUFT code, conservatively 
assuming that the emplacement drift is located below the access drift to the ventilation shaft. (In 
fact, the access drift to the ventilation shaft is located in the plan half way between two 
emplacement drifts.) The temperature fields (at 1, 2, 3, 5, 10, 20, 30 and 50 years) from the 
NUFT code were imported into 3DEC in order to calculate the thermal stress around the 
excavations, according to Equation 6-8. 
The temperature fields, as calculated in NUFT and imported into 3DEC, after 3 and 20 years of 
heating are shown in Figures 6-82K and 6-82L, respectively. As shown in these two figures, the 
temperatures at the middle of pillar are higher than those near the exhaust mains. This is mainly 
due to the conservative assumption made in the previous paragraph. The similar temperature 
distribution is also seen in temperature histories at points 1, 2, 3 and 4 (indicated in Figure 617C) 
shown in Figure 6-82M. 
The thermal stress changes are shown (at points 1, 2, 3 and 4) in Figure 6-82N. It is noted that 
the major principal stresses at points 1 and 2 (i.e., at middle of pillar) are higher than those at 
points 3 and 4 (i.e., at crown of drifts). The reason of this is due to the higher temperatures 
assumed in the pillar, which is already described in the previous paragraph. The reason that the 
stress at point 4 (i.e., at crown of access drift) is higher than that at point 3 (i.e., at crown of 
exhaust main) is due to the orientation of the major horizontal stress used in the 3-dimensional 
model. The model x-axis is aligned with the direction of the two exhaust mains. The orientation 
of the horizontal principal stresses is adjusted due to rotation of the model axes relative to the 
global coordinate system. The major horizontal principal stress strikes at N15E (see Assumption 
5.3). Due to rotation of the model horizontal axes by 18 counterclockwise, the major horizontal 
principal stress strikes in the model at N33E. However, it should be noted that the stress 
change due to the thermal load does not change the level of factor-of-safety with respect to the 
Mohr-Coulomb yield condition due to in situ stress (comparing Figure 6-66V with Figures 682O 
and 6-82P) nor cause significant additional yielding of the rock mass (comparing Figure 666X 
and Figure 6-82Q). 

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Figure 6-67. Intersection B: Temperature Field in Vertical Section 1 after 3 Years of Heating 
Figure 6-68. Intersection B: Temperature Field in Vertical Section 1 after 20 Years of Heating 

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Figure 6-69. Intersection B: Temperature History at Points 1, 2 and 3
Figure 6-70. Intersection D: Temperature History at Points 1, 2 and 3
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Figure 6-71. Intersection B: Time History of Major Principal Stresses for Lith. Cat. 1 Rock 
Figure 6-72. Intersection B: Time History of Major Principal Stresses for Lith. Cat. 5 Rock 

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Figure 6-73. Intersection D: Time History of Major Principal Stresses for N. Lith. Cat. 1 Rock 
Figure 6-74. Intersection D: Time History of Major Principal Stresses for N. Lith. Cat. 5 Rock 

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Figure 6-75. Intersection B: Factor of Safety in Vertical Section 1 for Lith. Cat. 1 Rock after 1 Year of 
Heating 
Figure 6-76. Intersection B: Factor of Safety in Vertical Section 1 for Lith. Cat. 1 Rock after 50 Years of 
Heating 

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Figure 6-77. Intersection D: Factor of Safety in Vertical Section 1 for N. Lith. Cat. 1 Rock after 1 Year of 
Heating 
Figure 6-78. Intersection D: Factor of Safety in Vertical Section 1 for N. Lith. Cat. 1 Rock after 50 Years of 
Heating 

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Figure 6-79. Intersection B: Potential Yield Zone in Vertical Section 1 for Lith. Cat. 1 Rock after 50 Years 
of Heating 
Figure 6-80. Intersection D: Potential Yield Zone in Vertical Section 1 for N. Lith. Cat. 1 Rock after 50 
Years of Heating 

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Figure 6-81. Intersection B: Time History of Vertical Displacement at Point 1 for Lith. Rock 
Figure 6-82. Intersection D: Time History of Vertical Closure at Point 1 for N. Lith. Rock 

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Figure 6-82A. Temperature Contours around Observation Drift after 50 Years of Heating 
Figure 6-82B. Temperature Histories at Three Points around Observation Drift during 50 Years of Heating 

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Figure 6-82C. Major Principal Stresses at Three Points around Observation Drift during 50 Years of 
Heating 
Figure 6-82D. Potential Yield Zone around Observation Drift after 50 Years of Heating 

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Figure 6-82E. Factor of Safety around Observation Drift after 50 Years of Heating 
Figure 6-82F. Displacement Histories at Three points around Observation Drift during 50 Years of 
Heating 

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Figure 6-82G. Potential Yield Zone at Intersection between Observation Drift and Exhaust Main in 
Vertical Section 1 after 50 Years of Heating 
Figure 6-82H. Potential Yield Zone at Intersection between Observation Drift and Exhaust Main in 
Vertical Section 2 after 50 Years of Heating 

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Figure 6-82I. Factor of Safety at Intersection between Observation Drift and Exhaust Main in Vertical 
Section 1 after 50 Years of Heating 
Figure 6-82J. Factor of Safety at Intersection between Observation Drift and Exhaust Main in Vertical 
Section 2 after 50 Years of Heating 

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Figure 6-82K. Temperature Field in Vertical Section 1 after 3 Years of Heating at Interburden Area 
Figure 6-82L. Temperature Field in Vertical Section 1 after 20 Years of Heating at Interburden Area 

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Figure 6-82M. Temperature Histories at Various Points at Interburden Area 
Figure 6-82N. Histories of Major Principal Stresses for Lith. Cat. 1 Rock at Interburden Area 

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Figure 6-82O. Safety Factor in Vertical Section 1 for Lith. Cat. 1 Rock after 1 Year of Heating at 
Interburden Area 
Figure 6-82P. Safety Factor in Vertical Section 1 for Lith. Cat. 1 Rock after 50 Years of Heating at 
Interburden Area 

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Figure 6-82Q. Potential Yield Zone in Vertical Section 1 for Lith. Cat. 1 after 50 Years of Heating 
6.5.3.3 Seismic Loading Condition 
6.5.3.3.1 Intersections between Access Mains and Turnouts 
The non-emplacement excavations including the access main, the turnout and their intersections 
will be subject to seismic loading during the preclosure period. The stability condition for these 
underground openings was analyzed for ground motion with a 10-4 probability of annual 
recurrence (i.e., earthquake with 10,000 year return period) (Section 4.1.5). The dynamic 
analysis was carried out for the intersection at location A (the most critical regarding the size of 
the span), assuming the intersection to be located in category 1 rock mass (poorest quality), of 
both the lithophysal and non-lithophysal units. 
The results of dynamic analysis are shown in Figures 6-83 through 6-93. Clearly the level of 
ground shaking due to an earthquake with 10,000 year return period is not expected to cause 
significant damage to access mains, turnouts, and intersections between access mains and 
turnouts. The continuum analysis indicates that the extent of inelastic deformation in the 
surrounding rock mass and, particularly, in the pillar between the two drifts, does not increase 
with respect to the volume of the rock mass yielded under static conditions (compare Figures 683 
through 6-86 with Figures 6-28, 6-29, 6-46, and 6-47). During the passage of the seismic 
waves, the stresses in the model oscillate. The maximum transient stress change caused by an 
earthquake with 10,000 year return period is approximately 1.1 MP and 2.1 MPa for the 
lithophysal and non-lithophysal units, respectively. Note that the stress changes are calculated 
based on Eq. 6-7, without multiplier 2, assuming the velocity amplitude to be 0.5 m/s. However, 
at the end of the simulation, when oscillations die out, the stresses return to the state that existed 

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prior to the shaking. Figures 6-87 and 6-88 indicate that there is no permanent stress change, or 
stress redistribution, indicating loosening or instability of a portion of the rock mass. The stress 
fields at the end of simulation in the lithophysal rock mass are colored by: a) the factor-of-safety 
with respect to the Mohr-Coulomb yield condition; and b) indication of tensile and compressive 
stresses, as shown in Figures 6-89 and 6-90, respectively. 
Inspecting these figures and comparing them with static results, it can be concluded that an 
earthquake with 10,000 year return period does not: a) cause significant, permanent change in the 
factor-of-safety with respect to the Mohr-Coulomb yield condition throughout the rock mass; or 
b) increase the volume of the rock mass subject to tensile stresses. The average stresses in the 
pillar between the drifts are almost unchanged as shown in Figure 6-91. All three components of 
the velocity histories were monitored during the simulations at four points in the roof and the 
walls of the intersection. The histories at point A in the roof, shown in Figure 6-92, are almost 
the same (offset for traveling time from the boundary to the history point) as applied velocities at 
the bottom boundary. The accumulation of permanent displacement due to inelastic deformation 
cannot be detected. Such an observation is confirmed by the plot of an increment of the residual 
displacement at the end of the dynamic simulation as shown in Figure 6-93. The maximum 
displacement increments are negligible. 
Seismic load due to an earthquake with 10,000 year return period does not result in a significant 
increase in damage of the rock mass around the drifts, nor an increase in the permanent 
displacements. The dynamic analysis of global stability of the intersections was carried out using 
the continuum models. 
Figure 6-83. Intersection A: Potential Yield Zone in Vertical Section 2 for Lith. Cat. 1 Rock under In Situ 
and Seismic Loads 

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Figure 6-84. Intersection A: Potential Yield Zone in Vertical Section 2 for N. Lith. Cat. 1 Rock under In Situ 
and Seismic Loads 
Figure 6-85. Intersection A: Potential Yield Zone in Horizontal Section for Lith. Cat. 1 Rock under In Situ 
and Seismic Loads 

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Figure 6-86. Intersection A: Potential Yield Zone in Horizontal Section for N. Lith. Cat. 1 Rock under In 
Situ and Seismic Loads 
Figure 6-87. Intersection A: Stress Field in Vertical Section 2 for Lith. Cat. 1 Rock under In Situ and 
Seismic Loads 

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Figure 6-88. Intersection A: Stress Field in Vertical Section 2 for N. Lith. Cat. 1 Rock under In Situ and 
Seismic Loads 
Figure 6-89. Intersection A: Factor of Safety in Vertical Section 2 for Lith. Cat. 1 Rock under In Situ and 
Seismic Loads 

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Figure 6-90. Intersection A: Compressive and Tensile Stresses in Vertical Section 2 for Lith. Cat. 1 Rock 
under In Situ and Seismic Loads 
Figure 6-91. Intersection A: Average Pillar Stress vs. Distance from Pillar Tip under In Situ and Seismic 
Loading 

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Figure 6-92. Intersection A: Velocity Histories in x, y and z Direction near Point A for Lith. Cat. 1 Rock 
Figure 6-93. Intersection A: Vertical Displacements along Vertical Line through Point A under Static and 
Seismic Loads 

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6.5.3.3.2 Intersections between Exhaust Mains and Emplacement Drifts 
The results of dynamic simulation of the intersections between exhaust mains and emplacement 
drifts for seismic ground motion with 10,000 year return period are shown in Figures 6-94 
through 6-101. Seismic loading was applied on the model at the initial state, after excavation of 
the drifts but before heating began. 
The results are presented for category 1 in both lithophysal (location B) and non-lithophysal 
(location D) rock masses. The dynamic load causes insignificant additional yielding, which was 
caused by stress oscillations. After completion of the dynamic simulation, the models for all 
simulated cases are stable. There is no indication of permanent stress redistribution or destressing 
of portions of rock mass indicating grounds prone to rockfall. Velocity histories were recorded 
during the simulations at a number of points throughout the model, particularly at the wall and 
back of the intersection. The histories near point 1 (located at the intersection of longitudinal 
axes of exhaust main and emplacement drift) in the intersections at locations B and D are shown 
in Figures 6-94 and 6-95, respectively. These histories are almost identical to the velocity 
histories applied at the base of the model due to the vertically propagating seismic wave. Thus, 
the model deforms elastically under seismic load. There is no indication of accumulation of 
irreversible, plastic deformation. 
Figure 6-94. Intersection B: Velocity Histories in X, Y and Z Directions near Point 1 for Lith. Cat. 1 Rock 

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Figure 6-95. Intersection D: Velocity Histories in X, Y and Z Directions near Point 1 for N. Lith. Cat. 1 
Rock 
Figure 6-96. Intersection B: Potential Yield Zone in Vertical Section 1 for Lith. Cat. 1 Rock under In Situ 
and Seismic Loads 

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Figure 6-97. Intersection D: Potential Yield Zone in Vertical Section 1 for N. Lith. Cat. 1 Rock under In 
Situ and Seismic Loads 
Figure 6-98. Intersection B: Potential Yield Zone in Horizontal Section for Lith. Cat. 1 Rock under In Situ 
and Seismic Loads 

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Figure 6-99. Intersection D: Potential Yield Zone in Horizontal Section for N. Lith. Cat. 1 Rock under In 
Situ and Seismic Loads 
Figure 6-100. Intersection B: Factor of Safety in Vertical Section 1 for Lith. Cat. 1 Rock under In Situ and 
Seismic Loads 

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Figure 6-101. Intersection D: Factor of Safety in Vertical Section 1 for N. Lith. Cat. 1 Rock under In Situ 
and Seismic Loads 
6.5.3.3.3 Observation Drift 
The observation drift was analyzed for seismic loading due to 10,000-yr ground motion in both 
typical cross-section and at the intersection between the observation drift and the exhaust main. 
The results are shown in Figures 6-101A to 6-101H. Seismic load causes insignificant increase in 
damage around the observation drift (compare Figures 6-101A, 6-101E and 6-101F with 
Figures 6-66C, 6-66G and 6-66H), but there is no indication of large-scale instability. A 
preclosure level of ground motion could cause some rockfall resulting from shaking down of 
loose blocks, but designed ground support will be sufficient to prevent most of such rockfall. The 
velocity histories in the tunnel crown (shown in Figure 6-101C) are almost identical to the 
velocity histories of the incoming seismic wave. There is no amplification of the velocities due to 
interaction of the wave with the excavations. The velocities indicate predominantly elastic 
response. The histories of the principal stress magnitudes (Figure 6-101D) are recorded in the 
crown and the walls of the observation drift. Variation of the major principal stress in the crown 
of the observation drift is smaller than the variation of the major principal stress in the walls. The 
major principal stress in the crown of the tunnel is horizontal and is shielded by the tunnel from 
the seismic wave propagating vertically upward. Fluctuations of the major principal stresses in 
the drift walls, which are vertical, are direct consequence of the passage of the P-wave (see 
comparison between the history of the vertical component of the velocity in Figure 6-101C with 
histories of the major principal stresses in the walls shown in Figure 6-101D). However, a trend 
of stress decrease at the points on the walls of the observation drift, within the region that is 
yielding plastically, can be observed in Figure 6-101D: the major principal stress decreases from 
approximately 12 MPa to 10 MPa. Seismic load causes some permanent stress redistribution 
that results in a small increase in the size of the plastic region. 

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Figure 6-101A. Potential Yield Zone around Observation Drift after Seismic Shaking 
Figure 6-101B. Factor of Safety around Observation Drift after Seismic Shaking 

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Figure 6-101C. Velocity Histories of Point at Crown of Observation Drift during Seismic Shaking 
Figure 6-101D. Stress Histories of Points at Observation Drift during Seismic Shaking 

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Figure 6-101E. Potential Yield Zone in Intersection between Observation Drift and Exhaust Main in 
Vertical Section 1 after Seismic Shaking 
Figure 6-101F. Potential Yield Zone in Intersection between Observation Drift and Exhaust Main in 
Vertical Section 2 after Seismic Shaking 

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Figure 6-101G. Factor of Safety in Intersection between Observation Drift and Exhaust Main in Vertical 
Section 1 after Seismic Shaking 
Figure 6-101H. Factor of Safety in Intersection between Observation Drift and Exhaust Main in Vertical 
Section 2 after Seismic Shaking 

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6.5.3.3.4 TBM Launch Chamber 
The results of the seismic analysis of the TBM launch chamber, summarized in Figures 6-101I to 
6-101L, are very similar to the results obtained for the observation drift. Preclosure seismic 
shaking (10,000-year earthquake) causes relatively small additional damage of the rock mass 
(Figure 6-101I), but there is no indication of global instability of the launch chamber during or 
after the shaking. As expected, based on similarity of the problems, the velocity and stress 
histories at the points around the excavations are almost identical for the TBM launch chamber 
(see Figures 6-101K and 6-101L) and the observation drift (see Figures 6-101C and 6-101D). 
6.5.3.3.5 North Portal 
The results of dynamic analysis of stability of the North Portal for the seismic ground motion are 
shown in Figures 6-101M to 6-101R. The velocity histories recorded at the ground surface 
(Figure 6-101M) are almost identical to the provided velocity histories of the design earthquake 
(Figure 4-1b). This comparison is a proof that boundary conditions are implemented correctly in 
the model. The velocity histories recorded at the crown of the starter tunnel (Figure 6-101N) 
indicate no additional amplification of velocity at the tunnel boundary and elastic model 
response to the seismic shaking. 
Shaking due to the 2,000-year ground motion does not cause any plasticity in the model. Stress 
tensor fields colored by the major principal stress magnitude and factor-of-safety with respect to 
Mohr-Coulomb shear failure are shown (Figures 6-101O to 6-101R) at two states during 
dynamic simulation: after 19.84 seconds, and at the end of dynamic simulation. Ground shaking 
does not cause residual changes in the stress field, which is consistent with the observation that 
model behaves elastically. The factor of safety with respect to Mohr-Coulomb shear failure 
remained large in the entire model throughout the dynamic simulation. 
6.5.3.3.6 Interburden Pillar between Shaft Access and Exhaust Mains 
The results of dynamic simulation of the pillar for preclosure seismic ground motion of 10,000-
year return period are shown in Figures 6-101S through 6-101W. Seismic loading was applied to 
the model at the initial state, after excavation of the drifts, before heating began. The results of 
the thermomechanical analysis indicate that if the seismic load were applied to the model at any 
stage during 50 years of the preclosure period, the results of the seismic analysis would not be 
different. 
The dynamic load causes insignificant additional yielding caused by stress oscillations 
(comparing Figures 6-101T and 6-101U with 6-66X and 6-66Y). The contours of safety factor 
after seismic loading are almost the same as those due to in-situ stress loading (comparing 
Figures 6-101V and 6-101W with 6-66V and 6-66W). After completion of the dynamic 
simulation, the models for all simulated cases are stable. There is no indication of permanent 
stress redistribution or destressing of portions of rock mass indicating areas prone to rockfall. 
Velocity histories were recorded during the simulation at a number of points throughout the 
model. The histories near the center of the model are shown in Figure 6-101S. These histories are 
almost identical to the velocity histories applied at the base of the model due to the vertically 

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propagating seismic wave (see Figure 4-1a). Thus, the model deforms elastically under seismic 
load. There is no indication of accumulation of irreversible, plastic deformation. 
Based on the results of this section and those from Sections 6.5.3.1.6 and 6.5.3.2.3, it is indicated 
that the effect of the excavation of an overlying 7.62-m-diameter access drift and two underlying 
7.62-m-diameter exhaust mains on the stability of 10-m high interburden pillar is minimum or 
insignificant. 
Figure 6-101I. Potential Yield Zone around TBM Launch Chamber after Seismic Shaking 

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Figure 6-101J. Factor of Safety around TBM Launch Chamber after Seismic Shaking 
Figure 6-101K. Velocity Histories of Point at Crown of TBM Launch Chamber after Seismic Shaking 

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Figure 6-101L. Stress Histories at Three Points around TBM Launch Chamber after Seismic Shaking 
Figure 6-101M. Velocity Histories in X, Y, and Z Directions at Ground Surface 

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Figure 6-101N. Velocity Histories in X, Y, and Z Directions at Crown of Starter Tunnel 

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Figure 6-101O. Factor of Safety after 19.84 Seconds of Ground Shaking: a) Horizontal Cross-section at 
Elevation 1128 m; and b) Longitudinal Cross-section L-L 

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Figure 6-101P. Major Principal Stresses after 19.84 Seconds of Ground Shaking for Cross-sections: 
a) C1-C1 and b) C2-C2 

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Figure 6-101Q. Factor of Safety at End of Ground Shaking: a) Horizontal Cross-section at Elevation 1128
m; and b) Longitudinal Cross-section L-L

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Figure 6-101R. Major Principal Stresses at End of Ground Shaking for Cross-sections: a) C1-C1 and b) 
C2-C2 

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Figure 6-101S. Velocity Histories in X, Y, and Z Directions near Point 1 for Lith. Cat. 1 Rock at 
Interburden Area 
Figure 6-101T. Potential Yield Zone in Vertical Section 1 for Lith. Cat. 1 Rock under In Situ and Seismic 
Loading at Interburden Area 

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Figure 6-101U. Potential Yield Zone in Vertical Section 2 for Lith. Cat. 1 Rock under In Situ and Seismic 
Loading at Interburden Area 
Figure 6-101V. Safety Factor in Vertical Section 1 for Lith. Cat. 1 Rock under In Situ and Seismic 
Loading at Interburden Area 

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Figure 6-101W. Safety Factor in Vertical Section 2 for Lith. Cat. 1 Rock under In Situ and Seismic 
Loading at Interburden Area 
6.5.3.4 Ground Reaction Curves 
The Ground Reaction Curve (GRC) is often used for the analysis of ground support requirements 
or to examine the interaction between ground support components and rock. The visualization of 
GRC can serve as a check of the stability of unsupported non-emplacement drifts, and to 
quantify the potential load that may be induced in ground support components due to further 
ground convergence after ground support component installation. In this calculation, the GRC 
concepts are described in conjunction with presenting the results of the analysis of the access and 
exhaust main in category 1 lithophysal rock. This scenario was chosen for demonstration as it 
represents the worst-case rock mass scenario. The displacements developed around the 
excavation and resulting loads in the ground support are significantly greater than any other rock 
mass conditions for these drifts at Yucca Mountain. 
To develop the GRC, forces that are statically equivalent to the initial stresses inside the drift are 
applied at each of the gridpoints defining the outline of the excavation. The forces are reduced 
gradually, and the resulting crown and wall displacements are monitored. 
Figures 6-102a, 6-102b, 6-103a and 6-103b show the ground reaction curves for unsupported 
access and exhaust mains and turnouts for roof and wall, respectively, by FLAC. 
Figures 6-102 and 6-103 indicate that, in all cases, the excavation comes to equilibrium, i.e., 
GRC reach finite displacement for zero confinement pressure, indicating the drifts are 
anticipated to be stable. With exception of deformation of wall for the excavations in the 

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lithophysal category 1 rock, the deformation in all other cases is completely elastic. The 
horseshoe-shaped turnout in lithophysal category 1 rock shows the largest displacements, i.e., 
about 37 mm for roof. The circular exhaust main with category 5 non-lithophysal rock shows 
the smallest displacement, i.e., about 0.3 mm for wall. Non-lithophysal rock shows smaller 
displacement than that of lithophysal rock. Among lithophysal or non-lithophysal rock, category 
1 rock shows larger displacement than that of category 5 rock. Clearly, the circular access and 
exhaust mains have more stable geometries, resulting in smaller displacements. Also, stronger 
rock such as non-lithophysal rock with category 5 quality shows smaller deformation compared 
with that of weaker rock such as lithophysal rock with category 1 quality. The major reason is 
due to the larger elastic modulus with the former compared with the latter (see Tables 4-2 and 41). 
It should be noted that equilibrium in the context of modeling with FLAC implies that no active 
failure planes have developed in the model. However, FLAC does not model the impact the 
displacements have on the dilation of intersecting joint planes that might cause wedges to loosen 
and fall under gravity into the excavation. Therefore, the equilibrium in the FLAC model 
without ground support does not imply that ground support is not necessary. 

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10
8
Access and Exhaust Main - Litho. C1 
Roof - Unsupported, Circular Excavation 
Access and Exhaust Main - Non Litho. C1 
Exhaust Main - Litho. C5 
Exhaust Main - Non Litho. C5 
P ressu re (M P a) 
6 
4 
2 
0 
0 10203040 
Displacement (mm) 
(a) 
5 
4 
3 
2 
1 
0 
WallAccess and Exhaust Main - Litho. C1 
- Unsupported, Circular Excavation 
Access and Exhaust Main - Non Litho. C1 
Exhaust Main - Litho. C5 
Exhaust Main - Non Litho. C5 
P r es su r e (M P a) 
02468 
Displacement (mm) 
(b) 
Figure 6-102. Ground Reaction Curves for Unsupported Access and Exhaust Mains: (a) Roof; (b) Wall. 

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Pressure (MPa)P ressure (M Pa)
10 
8 
6 
4 
2 
0 
0 10203040 
Roof - Unsupported, Horseshoe Excavation 
Turnout - Litho. C1 
Turnout - Non Litho. C1 
Displacement (mm) 
(a) 
12
10
8
6
4
2
0
Wall - Unsupported, Horseshoe Excavation 
Turnout - Litho. C1 
Turnout - Non Litho. C1 
0 24 6810 
Displacement (mm) 
(b) 
Figure 6-103. Ground Reaction Curves for Unsupported Turnouts: (a) Roof; (b) Wall. 

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6.5.4 Ground Support Systems for Non-Emplacement Drifts 
6.5.4.1 Candidate Ground Support System 
The ground support system proposed to be used in non-emplacement drifts is described as 
follows (Assumption 5.8): 
 For non-emplacement openings except intersections between access mains and 
turnouts and between exhaust mains and emplacement drifts, ramps, and starter 
tunnel: fully grouted rock bolts with heavy duty welded wire fabric (WWF). 
 For intersections between access mains and turnouts and between exhaust mains and 
emplacement drifts, ramps, and starter tunnel: fully grouted bolts with fiber-
reinforced shotcrete, and lattice girders as necessary for roof span control. 
Figures 6-104 and 6-105 show the sketch of ground support system in access and exhaust mains 
and turnouts, respectively. Figure 5-1 shows the typical shotcrete area at intersections. 
The fully grouted rock bolts with typical length of 3 m, spaced at 1.25 m with heavy duty WWF 
are designed to be used for ground support at typical non-emplacement openings, which include 
access mains, exhaust mains, observation drift and test alcove, TBM launch chambers, and North 
Portal starter tunnel. In the intersection areas, in order to enhance the opening stability with 
large roof span, increased bolt length of about 5 m with 0.10 m thick fiber-reinforced shotcrete 
will be installed, and supplemented with lattice girder as necessary depending on rock mass 
quality and the associated roof span. It is noted that the same ground support system for the 
intersections will be applied to ramps except that the bolt length is 3 m long, even though ramps 
have the same shape and dimension as those of access and exhaust mains. The major reason is 
that ramps provide access either for men, materials, waste emplacement, or ventilation air. They 
begin at the interface with the portal and ends at interface with the access mains. Due to their 
important functions, especially the waste package transportation for North Ramp, it is desirable 
to supplement fully grouted bolts by shotcrete to enhance the ground support function. It should 
be noted that the designed ground support system for ramps may be only for North Ramp and 
may need to be changed for North Construction Ramp and South Ramp as fully grouted rock 
bolts with heavy duty WWF, with fiber-reinforced shotcrete installed on as-needed basis, 
depending on the construction schedule. 
For the ground support at North Portal, fully grouted rock bolts with fiber-reinforced shotcrete 
will be installed around the portal frontal and lateral faces. Rock bolts of 3 m long with 1.5 m 
spacing will be installed at the lateral faces of the portal whereas rock bolts of 5 m long with 1.5 
m spacing will be applied at the frontal face. The portal face will be fibercreted to a thickness of 
0.1 m. Similar ground support design is expected for North Construction Portal and South Portal 
if their topographical and ground conditions, geometry, and construction method are similar. 
6.5.4.2 Candidate Ground Support Materials 
In order to make a proper selection of suitable ground support materials, it is important to know 
the service life in which these materials will be functional and understand under what kind of 
environment they will be subjected to during the service life. 

Title: Ground Control for Non-Emplacement Drifts for LA 168 
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For steel ground support components to be used in non-emplacement excavations, carbon steel 
including high-strength low-alloy steel is considered adequate as long as strength requirements 
are met. It should be noted that non-emplacement drifts are, in general, accessible during the 
preclosure period, i.e., proper maintenance of ground support system can be achieved. Steel 
components that are manufactured based on the following standard specifications included in the 
American Society for Testing and Materials (ASTM) are expected to perform satisfactorily in the 
non-emplacement drift environment: 
A 36 Standard Specification for Carbon Structural Steel,
A 242 Standard Specification for High-Strength Low-Alloy Structural Steel
A 588 Standard Specification for High-Strength Low-Alloy Structural Steel with 50 ksi
[345MPa] Minimum Yield Point to 4in. [100mm] Thick 
F 432 Standard Specification for Roof and Rock Bolts and Accessories 
A 82 Standard Specification for Steel Wire, Plain, for Concrete Reinforcement 
For cementitious materials to be used for grouted rock bolts, shotcrete or concrete, it is suggested 
to use a low pH grout and shotcrete/concrete mix, which will be developed with lesser amounts 
of portland cement and silica fume or to potentially use other types of grouts and cements which 
are of non-portland cement types. For information on material and properties and construction 
of shotcrete, the following documents are recommended: 
ACI 506R Guide to Shotcrete 
ACI 506.2 Specification for Shotcrete 
The typical dimension and material mechanical properties for fully grouted rock bolts and 
shotcrete are listed in Tables 4-3 and 4-4, respectively. 
6.5.5 Stability of Supported Non-Emplacement Drifts 
Based on the results presented in Section 6.5.3, static (in-situ stress), thermal (where applicable), 
and dynamic loads are not going to cause any major instability of the surrounding rock mass. 
The excavations appear to be stable without any ground support. However, a numerical model 
based on continuum mechanics does not account for the effects of joints and deformation along 
the joint due to stress relaxation on the formation of loose blocks, and eventual local rockfall. 
The access main, the turnout and particularly the intersections will be supported. 
The assessment of stability of supported non-emplacement drifts is presented in this section. The 
bolts and shotcrete were included in the models of the non-emplacement drifts, but the heavy-
duty wire mesh was not. The dimensions and mechanical properties of the rock bolts and the 
shotcrete are listed in Tables 4-3 and 4-4. The calculated loads in the ground support are the 
result of global deformation of the rock mass, subsequent to installation of the ground support. 
However, the purpose of the ground support is not to prevent or restrain the global deformation 
of the excavations, which are stable without any ground support, but to prevent local rock fall. 
The analysis of the ground support in the non-emplacement drifts except North Portal was 
carried out mainly for the category 1 lithophysal rock mass, which is the extreme condition of 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
rock mass quality at the repository level. The analysis of the ground support for North Portal was 
carried out for the category 1 rock in TCw unit. 
6.5.5.1 Mechanical Properties of Fully Grouted Rock Bolts and Swellex Bolts 
It is important to use appropriate material properties for modeling fully grouted bolts. Two key 
input parameters in the numerical approach are the bond stiffness (Kbond) and the bond shear 
strength (Sbond), which reflect the interaction between bolt and rock and control the bolt behavior. 
The bond stiffness of 8.68 x 108 N/m/m was estimated based on empirical correlation shown in 
Table 4-3. It is noted that the bond strength is a function of rock modulus; thus, the selection of 
properties should consider the host rock properties. Based on recommendation by Hutchinson 
and Diederichs (1996, Figure 2.6.13), the bond strengths for lithophysal category 1 rock (with 
Youngs modulus of 1.92 GPa) and non-lithophysal category 1 rock (with Youngs modulus of 
10.25 GPa) are 190 and 300 kN/m, respectively, for grout with water/cement ratio of 0.35. 
For the Swellex bolts to be used in emplacement drifts, the appropriate values for bond stiffness 
and bond strength are estimated to be 3108 N/m/m and 2.75105 N/m, respectively (see detailed 
discussion in Section 6.4.1 of BSC 2003f). 
Figure 6-104. Ground Support System in Access or Exhaust Mains 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
Note: same ground support for lithophysal rock 
Figure 6-105. Ground Support System in Turnouts 
6.5.5.2 In Situ Stress Loading Condition 
6.5.5.2.1 Access and Exhaust Mains and Turnouts 
The planned ground support for the access mains, exhaust mains and turnouts is analyzed using 
the FLAC code. 
Figure 6-106 shows the displacements around supported access mains for lithophysal category 1 
rock. Comparing this figure with Figure 6-18, the displacements are nearly identical in both 
figures, with the maximum displacement about 3.28 and 3.27 cm for unsupported and supported 
drifts, respectively. Figure 6-107 shows the axial load along fully grouted rock bolts around 
access mains for lithophysal category 1 rock. Note that the negative value in bolt load signifies 
tensile force in FLAC. The predicted load in the center bolt is 2 x 37.5 kN, or 70 kN/ meter of 
excavation. To account for the out-of-plane spacing by multiplying by 1.25; thus, the predicted 
bolt load for a center cable is 87.5 kN, which is much smaller than the yield strength of fully 
grouted bolt, i.e., 264 kN (see Table 4-3). Figures 6-108 and 6-109 show the safety factors of 
supported openings for access and exhaust mains in lithophysal and non-lithophysal rock, 
respectively. Figures 6-110 and 6-111 show the safety factors of supported openings for turnouts 
in lithophysal and non-lithophysal rock, respectively. 
Comparing Figures 6-108 through 6-111 with Figures 6-19 through 6-22, it appears that the 
safety factor contours for supported openings are almost identical to those of unsupported cases, 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
i.e., the effect of adding ground support to the modeled openings has minimal impact. This 
demonstrates an important point regarding the role of fully grouted rock bolt system in deep 
underground excavations. Ground support such as rock bolts should not be expected to arrest 
movement of the rock or reduce the rock mass stresses. Rather, it is there to support the loose 
blocks that result from the slip and dilation of joint planes intersecting the excavation surface. 
6.5.5.2.2 Intersections between Access Mains and Turnouts 
The analysis of the ground support in the access main intersections was carried out using 3DEC 
for the category 1 lithophysal rock mass, which is the extreme condition of rock mass quality at 
the repository level. 
The elements of ground support are simulated in the numerical model, following the expected 
sequence of their installation during tunnel construction. The bolts will be installed at a certain 
distance behind the excavation face. From the three-dimensional model, which simulated the 
face advance, it is assumed that 75% of stress relaxation is completed before the bolts are 
installed (if the face advance is not explicitly considered). The bolts are loaded due to 
displacements caused by the remaining 25% of stress relaxation. Although the model of the 
intersection is three-dimensional, simulation of the face advance and sequential installation of 
the bolts behind the face would be a very time-consuming process. 
The following are modeling steps performed during the analysis of the ground support in the 
intersection to simulate the consequence sequence. 
1. 75% of stress relaxation in the access main. 
2. Support of the access main with the grouted rock bolts and complete excavation of the 
access main. 
3. 75% of stress relaxation in the first turnout. 
4. Support of the first turnout with the grouted rock bolts and complete excavation of the 
first turnout. 
5. 75% of stress relaxation in the second turnout. 
6. Support of the second turnout with grouted rock bolts and complete excavation of the 
second turnout. 
7. Installation of the shotcrete in the intersection. 
According to this construction sequence, it is clear that shotcrete in the intersection is not going 
to be loaded due to stress relaxation, which is caused by the excavation of the drifts under in-situ 
stresses. The shotcrete will be loaded by deformation due to time-dependent strength degradation 
or seismic loading subsequent to installation of the shotcrete. In order to examine the effect of 
sequential turnout excavation on the forces in the rock bolts already installed in the access main, 
the excavation of the second turnout was simulated following the expected sequence of face 
advance and bolt installation behind the advancing face. 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
The bolt forces in the view along the access main, due to 25% of stress relaxation after 
excavation of the access main, are shown in Figure 6-112. Locally, in the bolts installed just 
above the springline, the forces reach 143 kN. However, the bolt forces are predominantly lower 
than 100 kN. The bolts in the back are loaded to almost 100 kN, while the bolts in the walls (with 
the exception of those above the springline) are loaded to less than 50 kN. This gives a safety 
factor of about 2.6 (264/100 = 2.64). In other words, the bolts have a reserve of at least 150 kN 
(i.e., 264  100 =164 > 150) to take the load by loose, locally unstable blocks. 
The effect of excavation of the first turnout on the bolt forces in the access main is illustrated in 
Figures 6-113 and 6-114. As expected, excavation of the turnout has the largest effect on the 
bolts already installed above the intersection. The force in the bolts circled in Figure 6-113 
increases from approximately 100 kN (shown in Figure 6-112 after excavation of the access 
main) to less than 200 kN. In the limited number of bolts that undergo the maximum force 
increase, the reserve available for additional load due to loose blocks is 50 kN (i.e., 264  200 
=64 > 50). A plot of the bolt forces in the plan view (Figure 6-114) indicates that the bolts that 
are additionally loaded due to excavation of the turnouts are located in the area above the 
intersection. The effect of the excavation of the turnout on the bolts in the access main rapidly 
decays as a function of distance in the plan view. That effect is illustrated in the plot (Figure 6115) 
of the bolt forces after both intersections are excavated. The region of interested bolt forces 
above the first intersection does not increase, nor do the forces in the bolts within that region 
increase due to excavation of the second turnout. Displacement vector fields shown in Figure 6116 
are in the vertical cross-section normal to the access main, through the center of the first 
intersection, for the model states before and after excavation of the second turnout. Any increase 
in displacements in this cross-section due to excavation of the second turnout is relatively small 
(shown at Figure 6-116b). Consequently, the excavation of the second turnout has little effect on 
the load in the bolts in the first intersection. 
6.5.5.2.3 Intersections between Exhaust Mains and Emplacement Drifts 
The ground support in the exhaust main is the same as the ground support used in the access 
main. The emplacement drifts will be supported by 3-m long Super Swellex rock bolts, at 1.25 m 
spacing, and Bernold-type perforated sheets (BSC 2003f, Section 8). The mechanical properties 
of the Super Swellex bolts used in the simulation of the ground support in the intersections 
between the exhaust main and the emplacement drifts are listed in Table 4-5. Bernold-type sheets 
were not included in the model. The ground support in the intersection (i.e., about 15-m length 
along the exhaust main and 5 m from the intersection into the emplacement drift) includes 5-m 
long rock bolts and 0.10-m thick steel-fiber reinforced shotcrete. The ground support analysis 
was carried out only for the intersection at location B, and using lithophysal rock mass category 
1, as the most unfavorable conditions for the load in the ground support. 
The loads in the bolts in the exhaust main after excavation of the exhaust main, looking along the 
exhaust main and in plan view, are shown in Figures 6-117 and 6-118, respectively. With the 
exception of a few bolts affected by the local conditions, the maximum bolt force is around 
100 kN. The forces in the bolts after excavation of both exhaust mains and the emplacement 
drifts are shown in Figures 6-119 through 6-121. The maximum force increases to approximately 
120 kN in the regions above the intersection. Those loads are well within the limit of the elastic 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
deformation of the bolts, which have yield strength of 264 kN, which gives a safety factor of 2.2. 
Note that the sign convention in 3DEC is that tensile force is positive. 
6.5.5.2.4 Observation Drift 
The ground support for the observation drift and the intersection between the observation drift 
and the exhaust main is described in Section 6.5.4.1. The ground support will be used even 
though the numerical models show that the observation drift will be stable without any ground 
support (see Section 6.5.3). The used models are based on continuum theory and do not consider 
stability of the blocks created by joints in the walls and the crown of the drift. The used 
numerical models provide assessment of the global stability of the excavation. The designed 
ground support will prevent rockfall due to local instability. 
The stability of supported observation drift was analyzed numerically. The objective was to 
assess performance of the ground support during the preclosure period for the expected loading 
conditions. Although the ground support is not designed to prevent (or reduce) deformation of 
the drift due to stress relaxation or thermal or seismic loading, it is important to demonstrate that 
such loads are not going to damage the ground support and make it ineffective in preventing 
local rockfall. The heavy-duty wire mesh was not included in the models. The observation drift 
is planned to be excavated using the drill-and-blast method. Consequently, it is assumed that 
ground relaxation due to excavation of the observation drift will be completed before the ground 
support inside the observation drift is installed. 
Two different excavation sequences that have an effect on rockbolt loads were considered with 
regard to the sequence of drift excavation: in one case, the observation drift was excavated first; 
in the other, the emplacement drift was excavated first. The analysis has shown that the sequence 
of excavation has minor effect on the rockbolt forces. However, because the rockbolt forces are 
larger if the emplacement drift is excavated after excavation and support of the observation drift, 
the results in this section are presented for that case only. 
The bolt forces after excavation of the emplacement drift are shown in Figure 6-121A. Note that 
the bolt forces in a two-dimensional model are presented per unit thickness of the model. The 
actual bolt forces can be obtained by multiplying values in the plot with spacing of the rows, 
which, in this case, is 1.25 m. These forces are insignificant considering the bolt yield strength 
of 264 kN. 
6.5.5.2.5 TBM Launch Chamber 
The designed ground support in the launch chamber includes 3-m long grouted rockbolts and 
heavy-duty wire mesh. Only the grouted rockbolts are included in the numerical model. The 
analysis was carried out for the seismic load induced by the preclosure earthquake only. It is 
assumed that the rockbolts do not take any load due to rock mass deformation caused by the 
relaxation of in situ stresses, as the launch chamber will be excavated using the drill-and-blast 
method. 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
6.5.5.2.6 North Portal 
Although it is demonstrated that the North Portal and the starter tunnel are stable under static and 
dynamic loading conditions, both excavations will be supported. The objective of the ground 
support is to prevent local instability (of blocks created by joints not accounted for in the models) 
and to slow deterioration of rock mass properties with time. 
For the North Portal modeling, shotcrete and rockbolt reinforcement have been considered (see 
Figure 6-105A). Three different regions, denoted as a) Lateral surfaces, b) Frontal surface and 
c) Tunnel surface, can be distinguished in Figure 6-105A. 
The ground support is installed after complete relaxation of the in situ stresses due to excavation 
of the North Portal and the starter tunnel. Consequently, the ground support does not take any 
load under static conditions. 
Figure 6-105A. Isometric View of Shotcrete and Rockbolts in North Portal Model 
6.5.5.3 Thermal Loading Condition 
The thermal loading condition for supported non-emplacement drifts is only applied to 
intersections between exhaust mains and emplacement drifts, observation drift, and intersection 
of observation drift with exhaust main. 

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DI: 800-KMC-SSD0-00700-000-00A 
6.5.5.3.1 Intersections between Exhaust Mains and Emplacement Drifts 
As illustrated in Figures 6-122 through 6-125, the forces in the bolts increase during the 
preclosure period due to heating. The predicted maximum force increases are between about 
20 kN and 35 kN (Figure 6-122). The predicted maximum forces in the rock bolts are between 
180 kN and 190 kN (i.e., at a safety factor of about 1.4) during the preclosure period, which are 
less than the yield strength. 
6.5.5.3.2 Intersection between Observation Drift and Exhaust Main 
Thermal load causes an increase in bolt forces, as indicated in Figure 6-125A, which shows the 
bolt forces after 50 years of heating, and in Figure 6-125B, which shows histories of bolt forces 
along three bolts: one in the crown, and two in the walls. Note that the number within the 
parenthesis in the legend indicates bolt number shown in Figure 6-125A. However, after 
50 years of heating, the loads are well within the range of elastic deformation of the rockbolts. 
Figure 6-125C shows the extent of inelastic deformation predicted around the observation drift. 
The designed bolt length of 3 m appears to be sufficient to provide anchoring into the elastic 
(undamaged) rock. 
6.5.5.4 Seismic Loading Condition 
6.5.5.4.1 Intersections between Access Mains and Turnouts 
The intersection between Access Mains and Turnouts supported with rock bolts and shotcrete is 
simulated for dynamic loading due to earthquake with 10,000 year return period. The shotcrete is 
represented in the model with triangular shell elements, which can take axial and bending forces. 
The shell elements are generated in such a way to coincide with triangular faces of zones 
exposed on the walls of the excavations. The geometry of the shotcrete support is shown in the 
plan view in Figure 6-126 (together with the bolts), and in the view along the access main in 
Figure 6-127. Prior to seismic loading, the shotcrete carries no load; the forces in the bolts are as 
described in Section 6.5.5.2.2. 
The unsupported intersection is stable after earthquake with 10,000 year return period even in 
category 1 lithophysal rock mass, the poorest quality rock mass. The increase in the permanent 
displacement is insignificant. Consequently, the residual load increase in the ground support after 
the seismic event predicted by this model is small. However, transient variations of the load 
(stresses) in the ground support could be a concern. Figures 6-128, 6-129 and 6-130 show bolt 
forces before shaking, after 40 seconds of shaking and at the end of shaking, respectively. The 
maximum transient bolt force increase during shaking to more than 200 kN (Figure 6-129) after 
40 seconds of ground shaking when one horizontal and the vertical component of velocity are 
approximately 0.4 m/s (as shown in Figure 6-92). Even the maximum transient load in the bolts 
is less than their yield strength. Knowing that steel is quite ductile (it can undergo substantial 
strain after yield before it breaks), it is not expected that earthquake with 10,000 year return 
period will cause significant damage of the rockbolts in the intersections, even in the poorest 
quality rock mass (i.e., category 1, lithophysal rock mass). Note in Figure 6-130 that two bolts, 
pulled out during the dynamic simulation, shown inside the access main, are yielding (indicated 
maximum force of 264 kN). These bolts are consequence of a local effect. All other bolts in the 

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DI: 800-KMC-SSD0-00700-000-00A 
model behave properly, with the maximum force less than 200 kN, i.e., a safety factor greater 
than 1.3. 
The shell elements used for simulation of the shotcrete are linearly elastic. The forces in the 
elements are checked during the simulation and they indicate that earthquake with 10,000 year 
return period will cause localized damage in the shotcrete, consisting mostly of tensile fractures. 
Figures 6-131 and 6-132 show the contours of the major principal membrane stress in the 
shotcrete after 40 seconds and at the end of ground shaking, respectively. Note that compression 
is positive in these two figures. It shows that there is indication of tensile stresses in the 
significant portion of the shotcrete during ground shaking. However, at the later stage of ground 
motion, i.e., at the end of dynamic simulation, the tensile forces do not exist in the shotcrete, 
except in local areas. It should be noted that shotcrete liner is not part of the ground support as a 
structural element, intended to carry the force. Its function is mainly to prevent weathering of the 
rock and hold small pieces of the rock from falling out. So, even if shotcrete fails (cracks) it does 
not affect its functionality. 
6.5.5.4.2 Intersections between Exhaust Mains and Emplacement Drifts 
For stability analysis of supported openings in intersections between exhaust mains and 
emplacement drifts, the maximum bolt forces, after 40 seconds of dynamic loading caused by 
ground motion with 10,000 year return period (applied after drift excavation but before heating 
began), are shown in Figure 6-133. The maximum increase in bolt forces compared to the static 
condition is approximately 55 kN in the region above the intersection, resulting in a maximum 
bolt force of about 185 kN (comparing Figure 6-133 to Figure 6-117) with a safety factor of 
about 1.4. 
6.5.5.4.3 Observation Drift 
The bolt forces in observation drift after seismic shaking is shown in Figure 6-134. The histories 
of bolt forces along three bolts in observation drift during seismic shaking is shown in Figure 6135. 
The extent of the potential yield zone around the observation drift after seismic shaking is 
shown in Figure 6-136. The maximum bolt force after seismic shaking is 73.1 kN 
(1.25m  58.47 kN/m), much less than the bolt yield strength of 264 kN. 
The ground support in the intersection between observation drift and exhaust main was analyzed 
using a three-dimensional model. The predicted bolt forces after 50 years of heating and after 
seismic ground shaking are shown in different views in Figures 6-137 to 6-141. The forces in the 
rockbolts installed in the exhaust main at the intersection with the observation drift are 
approximately 150 kN after thermal loading and during seismic loading. These loads are of the 
same order, but less than those predicted by the forces in the rockbolts at the intersection 
between the exhaust main and the emplacement drift (see Section 6.5.5.4.2). In the same time, 
the forces in the rockbolts in the observation drift in the intersection are quite small, less than 
50 kN. 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
6.5.5.4.4 TBM Launch Chamber 
The forces in the rockbolts at the end of dynamic simulation, shown in Figure 6-142, are small 
relative to the yield strength of the rockbolts. Note that the values indicated in the figure need to 
be multiplied by the 1.25-m spacing of the bolt rows to obtain the forces in the bolts. Figure 6143 
indicates that the forces in the bolts installed in the wall (in the region of the largest damage 
of the rock mass) have a residual component in addition to transient oscillation after seismic 
loading. Comparison of the bolt lengths and the size of the region of damaged rock suggests that 
the designed bolt length is sufficient, because the anchoring length into the elastic rock mass is 
approximately at least 1 m (see Figure 6-144). 
6.5.5.4.5 North Portal 
The forces and stresses in the ground support at North Portal and starter tunnel induced by the 
seismic loading at two stages during simulation are shown in Figures 6-145 and 6-146. The 
forces in the bolts are less than 10 kN. The yield strength of the fully grouted bolts is 264 kN. 
The compressive stresses in the shotcrete are very small, i.e., less than about 1.7 MPa (see 
Figures 6-147 and 6-148). However, the considered ground motion could cause localized tensile 
failure in the shotcrete. The failure will occur in the form of fractures that will not affect 
functionality of the shotcrete either on the slopes or in the starter tunnel. 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
Figure 6-106. Displacements around Supported Access Main for Lith. Cat. 1 Rock 
Figure 6-107. Axial Force along Bolts at Access Main for Lith. Cat. 1 Rock 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-108. Contours of Safety Factor around Supported Access/Exhaust Main for Lith. Cat. 1 Rock 
Figure 6-109. Contours of Safety Factor around Supported Access/Exhaust Main for N. Lith. Cat. 1 Rock 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-110. Contours of Safety Factor around Supported Turnout for Lith. Cat. 1 Rock 
Figure 6-111. Contours of Safety Factor around Supported Turnout for N. Lith. Cat. 1 Rock 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
Figure 6-112. Axial Forces in Bolts along the Access Main after Excavation 
Figure 6-113. Axial Forces in Bolts along the Access Main after Excavation of First Turnout 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
Figure 6-114. Plan View: Axial Forces in Bolts along the Access Main after Excavation of First Turnout 
Figure 6-115. Plan View: Axial Forces in Bolts along the Access Main after Excavation of both Turnouts 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
(a) 
(b) 
Figure 6-116. Displacement in Vertical Section 2 through First Intersection: a) before, b) after Excavation 
of Second Turnout 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
Figure 6-117. Axial Forces in Bolts along Exhaust Main after Excavation 
Figure 6-118. Plan View of Axial Forces in Bolts along Exhaust Main after Excavation 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
Figure 6-119. Axial Forces in Bolts along Exhaust Main after Excavation of Exhaust Main and 
Emplacement Drift 
Figure 6-120. Axial Forces in Bolts along Emplacement Drift after Excavation of Exhaust Main and 
Emplacement Drift 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
Figure 6-121. Plan View of Axial Forces in Bolts along Exhaust Main after Excavation of Exhaust Main 
and Emplacement Drift 
Figure 6-121A. Bolt Forces in Observation Drift after Excavation of Emplacement Drift 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
Figure 6-122. Time History of Bolt Forces at Exhaust Main and Emplacement Drift during Preclosure 
Figure 6-123. Intersection B: Forces in Bolts along Exhaust Main after 5 Years of Heating 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
Figure 6-124. Intersection B: Forces in Bolts along Exhaust Main after 50 Years of Heating 
Figure 6-125. Plan view of Forces in Bolts along Exhaust Main after 50 Years of Heating 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
Figure 6-125A. Bolt Forces in Observation Drift after 50 Years of Heating 
Figure 6-125B. Bolt Forces in Three Points of Observation Drift after 50 Years of Heating 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
Figure 6-125C. Potential Yield Zone around Observation Drift after 50 Years of Heating 
Figure 6-126. Plan View of Rock Bolts and Shotcrete at Intersection of Access Main and Turnout 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
Figure 6-127. Perspective View along Access Main showing Geometry of Shotcrete 
Figure 6-128. Axial Forces in Bolts along Access Main before Ground Shaking 

Title: Ground Control for Non-Emplacement Drifts for LA 
DI: 800-KMC-SSD0-00700-000-00A 
Figure 6-129. Axial Forces in Bolts along Access Main after 40 seconds of Ground Shaking 
Figure 6-130. Axial Forces in Bolts along Access Main at End of Ground Shaking 

Title: Ground Control for Non-Emplacement Drifts for LA 
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Figure 6-131. Contours of Major Principal Stresses in Shotcrete after 40 Seconds of Ground Shaking 
Figure 6-132. Contours of Major Principal Stresses in Shotcrete at End of Ground Shaking 

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Figure 6-133. Forces in Bolts along Exhaust Main after 40 Seconds of Ground Shaking 
Figure 6-134. Bolt forces in Observation Drift after Seismic Shaking 

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Figure 6-135. Histories of Forces in Three Bolts in Observation Drift during Seismic Shaking 
Figure 6-136. Potential Yield Zone around Observation Drift after Seismic Shaking 

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Figure 6-137. Plan View of Bolt Forces in Intersection between Observation Drift and Exhaust Main after 
50 Years of Heating 
Figure 6-138. Bolt Forces in Intersection between Observation Drift and Exhaust Main after 50 Years of 
Heating  View along Exhaust Main 

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Figure 6-139. Bolt Forces in Intersection between Observation Drift and Exhaust Main after 50 Years of 
Heating  View along Observation Drift 
Figure 6-140. Plan View of Bolt Forces in Intersection between Observation Drift and Exhaust Main after 
40 Seconds of Seismic Shaking 

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Figure 6-141. Plan View of Bolt Forces in Intersection between Observation Drift and Exhaust Main at 
End of Seismic Shaking 
Figure 6-142. Bolt Forces in TBM Launch Chamber after Seismic Shaking 

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Figure 6-143. Histories of Forces in Three Bolts in TBM Launch Chamber during Seismic Shaking 
Figure 6-144. Potential Yield Zone around TBM Launch Chamber after Seismic Shaking 

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Figure 6-145. Bolt Forces in North Portal after 19.84 Seconds of Ground Shaking 
Figure 6-146. Bolt Forces in North Portal at End of Ground Shaking 

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Figure 6-147. Contours of a) Major and b) Minor Principal Membrane Stresses in Shotcrete after 19.84 
Seconds of Ground Shaking 

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Figure 6-148. Contours of a) Major and b) Minor Principal Membrane Stresses in Shotcrete at End of 
Ground Shaking 

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6.6 UNCERTAINTY DISCUSSION 
There are uncertainties associated with this calculation. These uncertainties are discussed 
primarily in the following areas: empirical approach, numerical modeling methods, and input 
data. 
6.6.1 Uncertainties with Empirical Approach 
As discussed in Section 6.4, there are two empirical approaches adopted in this calculation for 
estimating the ground support need for openings at non-lithophysal rock in non-emplacement 
drifts, i.e., RMR and Q approaches. 
For RMR approach, the ground support recommendation shown in Table 6-1 was based on a 10m 
wide, horseshoe-shaped tunnel and mainly for conventional excavation method, i.e., not TBM 
operation. Since most tunnels or openings related to non-emplacement drifts are circular and 
will be excavated by TBM, therefore, there exist some uncertainties associated with the 
recommended ground support system based on RMR approach. In particular, the opening sizes 
for access and exhaust mains and turnouts are smaller than 10 m, which was based for Table 6-1. 
Therefore, although the recommended ground support system based on RMR approach provides 
good design guidance, it is considered very conservative. 
For Q approach, the most important value for estimating the ground support need is the Q value. 
However, in order to derive the appropriate Q value there are a number of steps to estimate 
various parameters as shown in Section 6.4.3, in which uncertainties are involved with each step. 
Moreover, in applying Q value to Figure 6-4, the corresponding bolt spacing and shotcrete 
thickness are estimated from ranges of values covered by various intervals. Again, uncertainties 
are associated with the selection of appropriate bolt spacing and shotcrete thickness. 
Since uncertainties are associated with both empirical approaches, their use will provide us the 
design guidance. The final selection ground support system will depend on numerical analysis 
as well as the experience obtained from construction of ESF and ECRB. 
6.6.2 Uncertainties with Numerical Modeling Method 
In this calculation, two-dimensional FLAC is used to simulate any long tunnels compared with 
its dimension. 3DEC is used to simulate the intersection area, which is appropriate for three-
dimensional layout configuration. FLAC3D is also run for very limited cases for comparison 
purposes. 
For numerical analysis of stability of underground openings, there are two major modeling 
methods, i.e., continuous and discontinuum. In a continuum approach, the geologic features in 
the rock mass are lumped into the constitutive model that represents the overall equivalent 
effect of these features. In a discontinuum approach, fractures or lithophysae are modeled 
explicitly as interfaces or cavities. Note that FLAC and FLAC3D are continuous approaches 
whereas 3DEC is a discontinuum one. However, for 3DEC approach used in this calculation, the 
joints between blocks are glued together to behave as a continuous one. 

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As mentioned before, all intersection areas were modeled with 3DEC except that FLAC3D run 
for very limited cases. The analysis using 3DEC has the advantage of generating model 
geometry easily and quickly. However, because of the tetrahedral zones with constant strain used 
in 3DEC, the calculation of plastic deformation is sometimes inaccurate. The advantage of 
FLAC3D is that it performs more accurate calculations of plastic deformation, particularly in the 
case of softening plasticity. However, FLAC3D models for such complex geometries as those of 
the intersections are very time consuming to generate. Although there are some uncertainties 
associated with these approaches, the difference between these techniques is therefore the level 
of detail that is necessary in the model to adequately capture the deformation and failure 
mechanisms. As shown in Section 6.5.3.1.2.1, the maximum displacements calculated based on 
FLAC3D are comparable to that of 3DEC results. From a ground support design perspective, 
stability of non-emplacement drifts is judged by overall rock mass displacements and stresses. 
Therefore, the adoption of the approaches used in the calculation is considered adequate. 
6.6.3 Uncertainties with Input Data 
The uncertainties associated with input data for the numerical analyses are mainly involved in 
the following areas: material properties, overburden depth, and opening sizes. 
The effect of the rock mass quality variability within the lithophysal and non-lithophysal units at 
the repository level was accounted for by considering different rock mass categories. As shown 
in Tables 4-1 and 4-2, there are six and five rock mass categories for lithophysal and nonlithophysal 
rock, respectively. Category 6 lithophysal rock is representative of extremely poor 
quality lithophysal rock mass in the disturbed rock zone around excavations, which was not 
considered in the simulations because it would result in excessively conservative, unrealistic 
predictions of deformation and damage of the rock mass. The computer simulations in this 
calculation were mainly carried out for the most conservative conditions, category 1 rock (the 
poorest quality rock mass), both in lithophysal and non-lithophysal units. Stability analysis for 
category 5 rock (the strongest quality rock mass) was also included mainly for intersection 
between exhaust main and emplacement drift since thermally induced stress is more in rock with 
higher Youngs modulus. Therefore, the ground support design based on the poorest quality is 
conservative and should also cover the good rock condition. 
The average depth of repository host horizon is assumed to be 400 m for all the model 
calculations. Although depth of non-emplacement drifts varies from drift to drift. The depths of 
emplacement drifts range from 215 to 450 m with the majority between 300 to 400 m. As 
discussed in Assumption 5.2 that the emplacement drift stability results for the maximum value 
of 450 m would be similar to those for the bounding case of 400 m. The overburden depth of the 
majority of non-emplacement drifts are less than 400 m, use of a depth of 400 m for calculating 
in situ stress at the non-emplacement drift horizon is considered adequate for the purpose of this 
calculation. 
There are several configurations in intersections between access mains and turnouts and between 
exhaust mains and emplacement drifts, with roof spans at openings varying from about 10 m to 
greater than 20 m. The intersection A at intersections between access mains and turnouts 
represents the critical conditions with respect to the maximum span of the roof and stress 
concentrations in the pillars whereas intersection C represents a typical layout. The intersections 

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B at intersections between exhaust mains and emplacement drifts represent intersection locations 
in the middle of the repository with large roof span whereas D represents locations at the edge of 
the repository. Temperature changes at the edge of the repository (location D) are overpredicted 
by NUFTs two-dimensional results, resulting in conservative estimate of stress increase and 
more unfavorable stability conditions. Also, the diameter of exhaust mains are assumed as 7.62 
m throughout the analysis, even though it is 5.5 m in Panel 1 of the repository. With large roof 
span, the larger stresses will be resulted, which will give conservative estimate on ground 
support system. 
Based on the above discussion, the model layout in terms of locations and opening sizes should 
result in conservative results regarding the on stability of non-emplacement drifts. 

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7. SUMMARY AND CONCLUSIONS 
7.1 SUMMARY 
This analysis for ground support design at non-emplacement drifts demonstrates that a 
satisfactory ground control system can be developed for the Yucca Mountain repository. The 
repository ground support design was based on both empirical approaches and analytical 
methods using acquired computer codes, and focused on the final support systems. 
Empirical approaches using RMR value and Q index were used to estimate the ground support 
need for openings in non-lithophysal rock. For analytical methods, both continuum and 
discontinuum modeling approaches were employed in the analyses of the rock mass behavior for 
both lithophysal and non-lithophysal rock and in the evaluation of the stability of the openings. 
Stability of the unsupported and supported non-emplacement drifts was evaluated in this 
analysis. The stability analysis was conducted for access mains, ramps, exhaust mains, turnouts, 
and intersections between main drifts with turnouts and emplacement drifts, observation drift, 
TBM launch chambers, and North Portal. The stability analysis of unsupported drifts at the 
interburden area between access drift to intake shaft #1 and two underlying exhaust mains was 
also performed. 
The excavation effects (i.e., state of the stress change due to excavation), thermal effects (i.e., 
due to heat output from waste packages), and seismic effects (i.e., from potential earthquake 
events) were evaluated, and stress controlled modes of failure were examined for representing 
rock mass categories of 1 and 5. The focus was on the category 1 rock because this rock mass is 
generally weaker than other categories of rock mass. 
No credit or account was given for the initial or temporary ground support in modeling the final 
ground support systems for non-emplacement drifts in this analysis. 
It should be noted that the results presented in this calculation are based on the adopted methods, 
input data and assumptions. If any input information is changed, the output information may 
vary. Also, as the design of ground support systems for LA progresses, the evaluation of ground 
control for non-emplacement drifts will be updated as necessary. 
7.2 STABILITY OF UNSUPPORTED NON-EMPLACEMENT DRIFTS 
For unsupported access/exhaust mains under in situ stress loading, a very narrow zone adjacent 
to wall (about 0.2 m into wall) is shown with potential yield for lithophysal rock. At distance of 
1.5 m into wall SF increases to 2. For non-lithophysal rock, almost no yield is observed near the 
wall and SF increases to 2 with a very short distance (about 0.2 m) into wall. There is no 
indication of instability of the rock surrounding the unsupported access/exhaust mains and 
turnouts. 
For unsupported intersection between access mains and turnouts, the 3DEC models predict 
similar deformation and stability conditions for intersections with large and small roof spans. 
The rock mass deformation around intersections will be predominantly elastic. The damage is 

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expected to extend approximately 1 m from the drift walls into the rock mass, in the worst case, 
for the poorest quality rock mass. A relatively large displacement of 7 cm predicted for the 
lithophysal rock mass, category 1, is a consequence of the very low Youngs modulus used. 
There are no indications of loosening and instability of the rock mass in the crown for any of the 
analyzed cases. The roof in the intersections appears to be stable, even for the largest spans. 
The tips of the pillars between the access main and the turnouts will be damaged due to stress 
concentrations. The extent of this damage from the pillar tip will depend on the rock mass 
quality. Rounding of the pillars and additional ground support will resolve the problem of pillar 
instability. 
Similar stability condition for intersections and pillars between exhaust mains and emplacement 
drifts as those at intersections at access mains are also obtained based on the computer results. 
The thermal stress changes are much larger in better quality and stiffer rock masses. The increase 
in the maximum major principal stress is about 15 MPa in category 5 of the lithophysal and nonlithophysal 
rock masses. However, the stress change does not change factor-of-safety with 
respect to the Mohr-Coulomb yield condition throughout the rock mass nor cause significant 
additional yielding of the rock mass. The extent of plastic deformation for locations B and D, and 
for the different rock-mass categories remains practically unchanged after 50 years of heating. 
The results of seismic analysis can be concluded that the seismic events with a mean annual 
exceedance probability of 1x10-4 does not cause significant, permanent change in the factor-of-
safety with respect to the Mohr-Coulomb yield condition or increase the volume of the rock mass 
subject to tensile stresses. The average stresses in the pillar between the drifts are almost 
unchanged. 
For unsupported observation drift, excavations of the emplacement drift and the observation drift 
do not interact with each other. The predicted extent of the damage of the rock mass surrounding 
the observation drift and the emplacement drift is similar. There is a little more damage in the 
walls and the floor of the observation drift because of mechanically less-favorable shape. 
However, the factor-of-safety with respect to Mohr-Coulomb shear failure increases very quickly 
as a function of distance from the drift boundary; the factor-of-safety is larger than 2 only a 
couple of meters from the drift walls. 
It does not appear that the temperature change causes any additional damage around the 
observation drift. It also indicates that thermal loads do not affect the stability conditions of the 
rock mass around the intersection between the observation drift and the exhaust main. 
The observation drift was analyzed for seismic loading due to 10,000-yr ground motion in both 
typical cross-section and at the intersection between the observation drift and the exhaust main. 
The results indicate that seismic load causes insignificant increase in damage around the 
observation drift, but there is no indication of large-scale instability. 
For unsupported TBM launch chamber, due to the span (or height) of the launch chamber is 
approximately two times of the span of the observation drift, the depth of the damaged rock in 
the wall of the launch chamber is about 2 m, compared to the depth of about 1 m predicted in the 

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wall of the observation drift. Outside the yielded region, the factor-of-safety to shear failure 
increases rapidly as a function of distance from the drift wall, indicating overall stability of the 
excavation. 
The results of the seismic analysis of the TBM launch chamber are very similar to the results 
obtained for the observation drift. Preclosure seismic shaking causes relatively small additional 
damage of the rock mass, but there is no indication of global instability of the launch chamber 
during or after the shaking. 
The stability analysis for interburden pillar between access drift to intake shaft #1 and two 
underlying exhaust mains subjected to in-situ stress, thermal, and seismic loading has been 
performed. The results indicate that the effect of the excavation of an overlying 7.62-m-diameter 
access drift and two underlying 7.62-m-diameter exhaust mains on the stability of 10-m high 
interburden pillar is minimum or insignificant. 
For unsupported North Portal including starter tunnel, the model under static in situ conditions 
after both the slope and the starter tunnel are excavated indicate that the stress state is completely 
elastic. Small overburden and slope heights result in stresses that are small compared to 
cohesion, even in the case of the poorest-quality rock mass, category 1, which was considered in 
the calculation. The factor-of-safety with respect to Mohr-Coulomb shear failure is quite large 
(approximately 5) throughout the entire model. The magnitude of the displacements due to the 
excavation of the starter tunnel is very small, of the order of 0.002 m or less. It appears from the 
results of the numerical modeling that the North Portal and the starter tunnel would be stable for 
the considered mechanical properties of the rock mass even if no ground support was used. 
7.3 GROUND SUPPORT SYSTEM FOR NON-EMPLACEMENT DRIFTS 
The ground support system designed for non-emplacement drifts is described as follows: 
 For non-emplacement openings except intersections between access mains and 
turnouts and between exhaust mains and emplacement drifts, ramps, and starter 
tunnel: fully grouted rock bolts with heavy duty welded wire fabric (WWF). 
 For intersections between access mains and turnouts and between exhaust mains and 
emplacement drifts, ramps, and starter tunnel: fully grouted bolts with fiber-
reinforced shotcrete, and lattice girders as necessary for roof span control. 
The fully grouted rock bolts with typical length of 3 m, spaced at 1.25 m with heavy duty WWF 
are designed to be used for ground support at typical access mains, exhaust mains, turnouts, 
observation drift and test alcove. For ground support system in intersection areas, bolting pattern 
with the same bolt spacing as that for access mains with increased bolt length of about 5 m and 
0.10 m thick fiber-reinforced shotcrete will be installed, and supplemented with lattice girder as 
necessary, depending on rock mass quality and control of roof span. 
It is noted that the same ground support system for the intersections will be applied to starter 
tunnel and ramps except that the bolt length is 3 m long. The designed ground support system for 
ramps may be only for North Ramp and may need to be changed for North Construction Ramp 

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and South Ramp as fully grouted rock bolts with heavy duty WWF, with fiber-reinforced 
shotcrete installed on as-needed basis, depending on the construction schedule. 
For the ground support at North Portal, 3 m long fully grouted rock bolts on 1.5 m spacing will 
be installed around the portal lateral faces whereas 5 m long bolts with the same spacing will be 
installed at the frontal face. The portal face will be fibercreted to a thickness of 0.1 m. Similar 
ground support design is expected for North Construction Portal and South Portal if their 
topographical and ground conditions, geometry, and construction method are similar. 
Candidate ground support components to be used in non-emplacement excavations will be 
carbon steels including high-strength low-alloy steel, which are considered adequate as long as 
they are manufactured based on the proper ASTM standard specifications and the strength 
requirements are met. For cementitious materials to be used for grouted rock bolts, shotcrete or 
concrete, a low pH grout and shotcrete/concrete mix or use of other types of grouts and cements 
made of non-portland cement types is proposed. 
7.4 STABILITY OF SUPPORTED NON-EMPLACEMENT DRIFTS 
For stability condition of supported non-emplacement openings, the computer models do not 
indicate formation of a failure mechanism, or the accumulation of residual displacement, which 
is a result of plastic deformation of the rock mass during shaking. However, the results do not 
imply that rockfall of any kind will not occur during an earthquake with 10,000 year return 
period. The limitation of this model is that it is formulated based on continuum mechanics. 
Consequently, it is not possible to simulate the formation of new or reopening of existing 
fractures, which could form loose blocks, resulting in a rockfall. The analysis implies that, if 
there is no ground support, the rockfall will be fairly limited and confined to the drift boundary. 
The designed ground support would prevent any potential rockfall. 
To a large extent the load in the rock bolts is determined by the percentage of total relaxation 
completed at the moment of bolt installation, and by the stiffness and shear strength of the grout. 
Assuming 75% of stress relaxation due to a stand-off between the advancing face and installed 
rock bolts, the maximum forces in the bolts in the access main before excavation of the turnouts 
is predicted to be 143 kN. However, the majority of the bolts are loaded to less than 100 kN. 
Excavation of the turnouts results in an increase in load to less than 200 kN in the bolts above the 
turnouts. The bolts will remain elastic due to the loads generated by the convergence of the drifts 
for in-situ stresses. The forces in the rock bolt generated by the convergence of the drifts will act 
as an active force, as if the bolts are pre-tensioned. However, the bolts will have a reserve 
between 50 kN and 100 kN for the additional loading. 
Seismic load has insignificant effect on the rock bolts. It appears that the shotcrete would be 
damaged during an earthquake with 10,000 year return period, but that damage (mostly tensile 
cracks) would be localized. 
The stability condition of supported exhaust mains and their intersections are similar to those of 
access mains and the adjacent intersections. The maximum load in the bolts in the exhaust main 
after excavation of the exhaust main is around 100 kN, with exception of a few bolts affected by 
the local conditions. The maximum force in the bolts increases to approximately 120 kN in the 

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regions above the intersection after excavation of both exhaust mains and the emplacement 
drifts. Those loads are well within the limit of the elastic deformation of the bolts, which have 
yield strength of 264 kN. 
The forces in the bolts increase during the preclosure period due to heating. However, even using 
the conservative analysis, the predicted maximum force in the rock bolts is about 185 kN during 
the preclosure period, which is less than the yield strength. The maximum increase in bolt forces 
due to seismic loading compared to the static condition is approximately 55 kN in the region 
above the intersection, with the bolt load less than the yield strength. Note that the dynamic 
analysis of global stability of the intersections was carried out using the continuum models. In 
the simulations of intersections, the blocks were glued together to behave effectively as a 
continuum. Thus, the local stability of blocks created by joints around the excavation was not 
considered. However, such blocks will be of limited size and number and will be effectively 
supported by the designed ground support. 
For supported observation drift under static loading, the forces in bolts are insignificant 
considering the bolt yield strength of 264 kN. Thermal load causes an increase in bolt forces, 
however, the loads are well within the range of elastic deformation of the rockbolts after 50 years 
of heating. The designed bolt length of 3 m appears to be sufficient to provide anchoring into the 
elastic (undamaged) rock. The maximum bolt force after seismic shaking is much less than the 
bolt yield strength of 264 kN. 
For supported TBM launch chambers, the rockbolts do not take any load due to rock mass 
deformation caused by the relaxation of in situ stresses, as the launch chamber will be excavated 
using the drill-and-blast method. The forces in the rockbolts at the end of dynamic simulation are 
small relative to the yield strength of the rockbolts. Comparison of the bolt lengths and the size 
of the region of damaged rock suggests that the designed bolt length is sufficient, because the 
anchoring length into the elastic rock mass is approximately at least 1 m. 
For supported North Portal, the ground support is installed after complete relaxation of the in situ 
stresses due to excavation of the North Portal and the starter tunnel. Consequently, the ground 
support does not take any load under static conditions. The forces and stresses in the ground 
support at North Portal and starter tunnel induced by the seismic loading at two stages during 
simulation are analyzed. The forces in the bolts are less than 10 kN, which is much smaller than 
the yield strength of the fully grouted bolts, i.e., 264 kN. The compressive stresses in the 
shotcrete are very small. However, the considered ground motion could cause localized tensile 
failure in the shotcrete. The failure will occur in the form of fractures that will not affect 
functionality of the shotcrete either on the slopes at the portal surface or in the starter tunnel. 

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Title: Ground Control for Non-Emplacement Drifts for LA 
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Title: Ground Control for Non-Emplacement Drifts for LA 213 
DI: 800-KMC-SSD0-00700-000-00A 
Minwalla, H.J. 2003. Project Design Criteria Document. 000-3DR-MGR0-00100-000-001. Las 
Vegas, Nevada: Bechtel SAIC Company. ACC: ENG.20030402.0001. 
Onofrei, M.; Gray, M.N.; Pusch, R.; Borgesson, L.; Karnland, O.; Shenton, B.; and Walker, B. 
1993. Sealing Properties of Cement-Based Grout Materials Used in the Rock Sealing Project. 
AECL-10815. Pinawa, Manitoba, Canada: Atomic Energy of Canada Limited, Whiteshell 
Laboratories. TIC: 210964. 
Ruest, M. and Martin, L. 2002. FLAC Simulation of Split-Pipe Tests on an Instrumented Cable 
Bolt. CIM Vancouver 2002, Paper No. 287. Montreal, Canada: Canadian Institute of Mining, 
Metallurgy and Petroleum. TIC: 255291. 
Wickham, G.E.; Tiedemann, H.R.; and Skinner, E.H. 1972. "Support Determinations Based on 
Geologic Predictions." Proceedings: North American Rapid Excavation and Tunneling 
Conference, Chicago, Illinois, June 5-7, 1972. 43-64. New York, New York: American Institute 
of Mining, Metallurgical, and Petroleum Engineers. TIC: 226274. 
8.2 CODES, STANDARDS, REGULATIONS, AND PROCEDURES 
ACI (American Concrete Institute) 506R-90(95). 1995. Guide to Shotcrete. Farmington Hills, 
Michigan: American Concrete Institute. TIC: 238721. 
ACI 506.2-95. 1995. Specification for Shotcrete. Detroit, Michigan: American Concrete Institute. 
TIC: 226412. 
AISC (American Institute of Steel Construction) 1997. Manual of Steel Construction, Allowable 
Stress Design. 9th Edition, 2nd Revision, 2nd Impression. Chicago, Illinois: American Institute 
of Steel Construction. TIC: 240772. 
AP-3.12Q, Rev. 2, ICN 1. Design Calculations and Analyses. Washington, D.C.: U.S. 
Department of Energy, Office of Civilian Radioactive Waste Management. ACC: 
DOC.20030827.0013. 
ASTM (American Society for Testing and Materials) ASTM A 36/A 36M-03a. 2003. Standard 
Specification for Carbon Structural Steel. West Conshohocken, Pennsylvania: American Society 
for Testing and Materials. TIC: 254846. 
ASTM A 82-97a. 1998. Standard Specification for Steel Wire, Plain, for Concrete 
Reinforcement. West Conshohocken, Pennsylvania: American Society for Testing and Materials. 
TIC: 246400. 
ASTM A 242/A 242M-03a. 2003. Standard Specification for High-Strength Low-Alloy 
Structural Steel. West Conshohocken, Pennsylvania: American Society for Testing and 
Materials. TIC: 254844. 

Title: Ground Control for Non-Emplacement Drifts for LA 214 
DI: 800-KMC-SSD0-00700-000-00A 
ASTM A 276-03. 2003. Standard Specification for Stainless Steel Bars and Shapes. West 
Conshohocken, Pennsylvania: American Society for Testing and Materials. TIC: 254842. 
ASTM A 588/A 588M-03. 2003. Standard Specification for High-Strength Low-Alloy Structural 
Steel with 50 ksi [345MPa] Minimum Yield Point to 4in. [100mm] Thick. West 
Conshohocken, Pennsylvania: American Society for Testing and Materials. TIC: 254843. 
ASTM F 432-95 (Reapproved 2001). 1995. Standard Specification for Roof and Rock Bolts and 
Accessories. West Conshohocken, Pennsylvania: American Society for Testing and Materials. 
TIC: 253986. 
BSC 2002. Software Code: 3DEC. V2.01. PC WINDOWS 2000/NT 4.0. 10025-2.01-00. 
BSC 2002. Software Code: FLAC. V4.0. PC WINDOWS 2000/NT 4.0. 10167-4.0-00. 
BSC 2002. Software Code: FLAC3D. V2.1. PC WINDOWS 2000/NT 4.0. 10502-2.1-00. 
8.3 SOURCE DTNS 
MO0306SDSAVDTH.000. Seismic Design Spectra and Acceleration, Velocity, and 
Displacement Time Histories for the Emplacement Level at 10-4 Annual Exceedance Frequency. 
Submittal date: 06/26/2003. 
MO0402SDSTMHIS.004. Seismic Design Spectra and Time Histories for the Surface Facilities 
Area (Points D/E) for 5E-4 Annual Exceedance Frequency. Submittal date: 02/09/2004. 
SNL01B05059301.006. Laboratory Thermal Expansion Data for Boreholes UE25 NRG-4, NRG5; 
USW NRG-6 and NRG-7/7A. Submittal date: 02/07/1996. 
SNF37100195002.001. Hydraulic Fracturing Stress Measurements in Test Hole: ESF-AOD-
HDFR1, Thermal Test Facility, Exploratory Studies Facility at Yucca Mountain. Submittal date: 
12/18/1996.