Technical Basis Document No. 4: Mechanical Degradation and Seismic Effects Revision 1

June 2004 

1. INTRODUCTION 
This technical basis document provides a summary of the stability of repository excavations and 
potential mechanical degradation under the action of in situ, thermal, and seismic stresses during 
the preclosure and postclosure time periods. The document further identifies the interactions of 
the mechanical degradation of the emplacement drifts with the engineered barriers, in-drift 
environment, and water seepage into emplacement drifts. This document is one in a series of 
technical basis documents prepared for each component of the Yucca Mountain repository 
system relevant to predicting the likely postclosure performance of the repository. The 
relationship of mechanical degradation and seismic effects to the other components of the 
repository system is illustrated in Figure 1-1. 
The information presented in this document, and the associated references, is part of the ongoing 
development of the postclosure safety analysis that will be included in the license application. 
This information is also used to respond to open Key Technical Issue (KTI) agreements made 
between the U.S. Nuclear Regulatory Commission (NRC) and the U.S. Department of Energy 
(DOE). Placing the DOE responses to individual KTI and NRC Additional Information Needed 
(AIN) requests within the context of the overall mechanical degradation analyses, as it relates to 
1-1 June 2004 No. 4: Mechanical Degradation Revision 1 
postclosure safety analyses, allows for a more direct discussion of the relevance of the 
agreements. 
Appendices to this document are designed to allow for a transparent and direct response to each 
KTI agreement and AIN requests. Each appendix addresses one or more of the agreements. If 
agreements apply to similar aspects of the mechanical drift degradation issue, they were grouped 
in a single appendix. In some cases, appendices provide detailed discussions of data, analyses, 
or information related to the further conceptual understanding presented in this technical basis 
document. In other cases, the appendices provide information that is related to the technical 
basis document information but at a level of detail that relates more to the uncertainty in a 
particular data set or feature, event, or process that is less relevant to the overall technical basis. 
In these cases, the appendices reference the relevant section of the technical basis document to 
put the particular KTI agreement into context, but the technical basis document does not 
reference the appendices. 
This technical basis document provides a summary-level synthesis of many relevant aspects of 
the mechanical drift degradation and ground support design studies. This document describes 
the development of an understanding of the thermal-mechanical behavior of Yucca Mountain 
tuffs and the development of suitable models to represent this behavior when subjected to static 
and transient loading. The analyses described here are applied primarily to estimating the 
stability and rockfall potential of the excavations, and less to the design of ground support, which 
is primarily a preclosure design issue. However, the techniques described here are applicable to 
both the design and performance studies. 
This document presents a summary and synthesis of the detailed technical information presented 
in the analyses and model reports and other technical products that are used as the basis for the 
description of the mechanical degradation analysis and the incorporation of this work into the 
postclosure performance assessment. Several analyses, model reports, and other technical 
products support this summary: 
 Drift Degradation Analysis (BSC 2004a) 
 Subsurface Geotechnical Parameters Report (BSC 2003a) 
 Longevity of Emplacement Drift Ground Support Materials for LA (BSC 2003b) 
 Ground Control for Emplacement Drifts for the LA (BSC 2003c) 
 Scoping Analysis on Sensitivity and Uncertainty of Emplacement Drift Stability 
(BSC 2003d) 
 Ground Control for Non-Emplacement Drifts for LA (BSC 2004b) 
 Development of Earthquake Ground Motion Input for Preclosure Seismic Design and 
Postclosure Performance Assessment of a Geologic Repository at Yucca Mountain, NV 
(BSC 2004c) 
 Geology of the ECRB Cross DriftExploratory Studies Facility, Yucca Mountain 
Project, Yucca Mountain, Nevada (Mongano et al. 1999) 
June 2004 1-2 No. 4: Mechanical Degradation Revision 1 
 Resolution Strategy for Geomechanically-Related Repository Design and 
Thermal-Mechanical Effects (RDTME) (Board 2003). 
The basic approach used in this document is to review the relevant portions of these reports to 
provide a comprehensive overview of the data and analyses that describe the process of 
mechanical degradation and to provide a context for the attached KTI resolution documents. The 
details of the data and analyses can be found in the above reports. 
1.1 OBJECTIVE AND SCOPE 
The objectives of this technical basis document are to: 
 Provide an overview of the anticipated mechanical degradation processes in the 
emplacement drifts and the boundary conditions this degradation provides to the 
engineered barrier system and drift seepage 
 Present an overview of the KTI agreements, how they relate to this process, and the 
strategy employed to resolve these issues 
 Present a summary of the relevant geomechanical rock mass properties database 
developed for lithophysal and nonlithophysal rocks 
 Describe the numerical modeling and analysis techniques used for ground support and 
drift degradation analysis 
 Present the analysis of postclosure mechanical degradation under the action of in situ, 
thermal, and seismic stresses, as well as time-dependent strength changes. 
The purpose of the mechanical degradation analyses is to provide an estimate of the temporal 
evolution of the stability of the emplacement drifts under the action of in situ, thermal, and 
seismic loading, as well as time-dependent changes in rock mass strength. In particular, the 
model(s) provide an estimate of: 
 The yield of the rock mass around the emplacement drifts, and the associated rockfall 
once the ground support function is lost due to postclosure corrosion 
 The temporal change in emplacement drift shape and size 
 The temporal evolution of rockfall, including the rockfall particle size distribution and 
total amount 
 The temporal evolution of rockfall loading (both dynamic and quasi-static1) on the drip 
shield. 
1 Quasi-static refers here to the development of loading on the drip shield, which slowly evolves over time but is 
more-or-less in static equilibrium. 
June 2004 1-3 No. 4: Mechanical Degradation Revision 1 
Due to the distinctly different mechanical characteristics of the major rock units that comprise 
the repository host horizon (lithophysal and nonlithophysal rocks), the above estimates will vary 
with location within the repository. 
For the postclosure period, 10 CFR 63.114 requires that DOE conduct a performance assessment 
that considers only events that have at least one chance in 10,000 of occurring over 10,000 years. 
Seismic events that have this probability of occurrence have been analyzed and are discussed in 
this technical basis document. In addition, 10 CFR 63.102(j) provides that, for the postclosure 
period, the event classes analyzed in the performance assessment should consist of all possible 
specific initiating events that are caused by a common natural process (e.g., the event class for 
seismicity includes the range of credible earthquakes for the Yucca Mountain site). Additional 
sensitivity studies are currently being conducted to ascertain the range of credible ground 
motions, and those studies are not discussed in this document. The additional studies do not 
affect the technical bases or conclusions described in this document. The results of those 
sensitivity studies will be included in the license application, as appropriate. 
1.2 SUMMARY OF CURRENT UNDERSTANDING 
1.2.1 General Description of the Repository Layout and Waste Emplacement 
A general description of the design and layout of the repository can be found in Underground 
Layout Configuration (BSC 2003e). The repository is located at approximately 300 m below 
ground surface within the Topopah Spring Tuffa densely welded tuff unit comprised of a 
number of subunits that dip approximately 15 from west to east. These subunits can be divided 
into two broad, mechanical categories: nonlithophysal and lithophysal2 welded tuffs. The basic 
matrix material of these two subunits is similar in mineralogical, textural, and mechanical 
properties. However, due to varying cooling histories and as a result of position within the 
formation, they are structurally and thermal-mechanically significantly different in character. 
The nonlithophysal rocks are hard, mechanically strong, fine-grained and fractured volcanic 
rocks whose mechanical behavior is strongly controlled by the geometry and surface 
characteristics of its fracturing. The lithophysal rock is composed of the same strong, hard 
matrix material, but has porosity in the form of lithophysal cavities ranging from about 10% to 
30% by volume. The presence of these cavities results in significantly different deformability 
and strength of the rock mass. The current repository layout places approximately 85% of the 
emplacement drifts within the lithophysal rocks (BSC 2003e, Table I-2). 
The repository (Figure 1-2) is accessed from ground surface on the east side of Yucca Mountain 
by three 7.62 m diameter entry ramps (the north and south ramps currently exist and are part of 
the Exploratory Studies Facility (ESF) and are driven at 2.5% grade). Once these ramps achieve 
the depth of the Topopah Spring Tuff, they join a number of subhorizontal, 7.62-m diameter 
access mains. These access mains define the outer perimeter of four panels that are composed of 
5.5 m nominal diameter waste emplacement drifts (BSC 2003e). The emplacement drifts are 
accessed at one end from the fresh air intake main via a turnout (Figure 1-3) and intersect the 
exhaust main at the opposite end. The excavations, with the exception of the turnouts, are driven 
2 Lithophysae: hollow, bubblelike cavities in a volcanic rock surrounded by porous rims formed by fine-grained 
alkali feldspar, quartz, and other minerals. Lithophysae are typically a few centimeters to a few decimeters in 
diameter; however, they can be as small as 1 mm in diameter or less to as large as 1 m or more in diameter. 
1-4 June 2004 No. 4: Mechanical Degradation Revision 1 
by tunnel boring machines (TBMs). The access mains are supported with rock bolts and heavy 
wire mesh, the turnout-access mains intersections with rock bolts and shotcrete, and the 
emplacement drifts with thin (2 to 3 mm), perforated, stainless steel sheeting held tight to the 
drift walls using stainless steel, friction-type rock bolts (Figure 1-4). All of the ground support is 
placed over a 240 arc on roof and walls. The general ventilation method employed is to draw 
fresh air into the repository via the ramps and a number of intake shafts, distribute it down the 
intake mains, through the emplacement drifts, and out through a series of exhaust shafts (BSC 
2003e, Table I-2). 
June 2004 1-5 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2003e. 
NOTE: Footprint of emplacement area boundary is shown as a dashed line. This footprint represents the currently 
characterized area in which emplacement drifts can be located. 
Figure 1-2. Repository Layout in Plan View Showing Ramps, Access Intake and Exhaust Mains, 
Emplacement Drifts, Shafts, and Existing Excavations (as Dashed Lines) 
June 2004 1-6 No. 4: Mechanical Degradation Revision 1 
Figure 1-3. Plan View and Cross-Sectional Views of Primary Repository Excavations 
June 2004 1-7 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2003c and BSC 2004d. 
NOTE: Ground support coverage using Bernold-Type stainless steel sheets (240 of circumference above invert) 
and stainless steel friction bolts. 
June 2004 
Figure 1-4. Ground Support Method for Emplacement Drifts Showing Drip Shield 
A rail-based waste emplacement transporter, drawn by electric locomotives, will leave the 
surface waste handling facilities with a waste package on the emplacement pallet loaded within a 
shielded rail car. The transporter will deliver the waste package, via the access mains, to the 
emplacement drift turnout. In the turnout, the waste package will be transferred, in an automated 
mode, by pushing the waste package and pallet from the transporter onto a docking bay at the 
end of the emplacement drift. The waste package will then be picked up from the loading dock 
and transported into the emplacement drift by an emplacement gantry crane (Figure 1-5). The 
emplacement gantry will deliver the waste package and pallet to their resting location on the drift 
invert structure, and will return to the docking bay. Normal retrieval operations, if required, will 
involve a reversal of the emplacement sequence. At closure, a titanium drip shield will be placed 
over the top of the waste packages (Figure 1-4), and all nonemplacement drifts will be backfilled 
with crushed tuff from the TBM operations. 
1-8 No. 4: Mechanical Degradation Revision 1 
1.2.2.1 Preclosure Effects 
Excavation of the repository drifts will result in concentration of in situ stresses around the 
openings. Because the in situ stresses are relatively small in comparison to the rock mass 
strength, little, if any, yield of the rock mass is expected. Thus, the openings will undergo 
primarily elastic deformation, which equilibrates within a short distance (about two tunnel 
diameters) behind the advancing TBM. Light, temporary ground support is placed directly 
behind the TBM for worker protection purposes. Permanent ground support is placed after the 
emplacement drift is completed and the TBM withdrawn, therefore, it will be subjected only to 
deformation and loading that may occur from transient effects such as thermal and seismic 
loading. 
During the preclosure time period, approximately 100 years from initial excavation of the 
emplacement drifts, forced ventilation will be used to remove approximately 90% of the heat 
generated by the waste packages. This heat removal will keep drift wall temperatures cooler 
1.2.2 General Degradation Processes 
The general process of mechanical degradation of emplacement drifts during the preclosure and 
postclosure time frames involves interactions with several engineered and natural barrier 
systems. A general description of the impact or interaction of the mechanical degradation of the 
emplacement drifts is given here. 
Figure 1-5. Shielded Waste Transporter at Docking Bay Showing Waste Package, Waste 
Emplacement Gantry, and Locomotive 
June 2004 1-9 No. 4: Mechanical Degradation (BSC 2004a, Section 6.2) and will result in small thermally related rock mass stress changes (see 
Section 5.3.1). The emplacement drifts will be supported by rock bolts and slotted stainless steel 
sheeting, thus minimizing, if not eliminating, mechanical degradation of the excavations. 
Although this technical basis document is concerned primarily with the postclosure mechanical 
and seismic degradation effects, a number of the Repository Design and Thermal-Mechanical 
Effects (RDTME) KTI agreements either relate specifically to preclosure ground support 
performance or are concerned with rock mass characterization and modeling issues that are 
common to both preclosure and postclosure performance. Since a detailed review of preclosure 
design issues is not given in this document, Table 1-1 is provided to give a cross-reference 
between the KTI agreement resolution summaries found in the appendices and the specific 
source document that provides greater details of the analyses that support the agreement. 
Appendix 
A 
B 
C 
D 
E 
F 
G 
H 
1.2.2.2 Postclosure Effects 
Repository closure will involve installation of titanium drip shields over the waste packages, 
backfilling of access mains and shafts, and the cessation of forced ventilation. This will result in 
a rapid rise in temperature of the drift walls, reaching approximately 140C to 165C within 
NOTE: Document A is BSC 2003d. Document B is BSC 2003c. Document C is BSC 2004b. Document D is BSC 
2003a. 
Data and analysis of rock bridges between joints Sections 3.2, 4.1, and 5.3 
Continuum and discontinuum analyses of ground Section 4.1, 4.2 
support system performance 
Where Addressed 
Sections 3.2 and 4.2 
Sections 3.2, 4.1, 4.2, and 5.3 
Document A 
Sections 4.1, 4.2, and 5.3 
Documents A and B 
Section 3.2 
Document D 
Section 3.2, 4.1, and 5.3 
Documents A, B, C, and D 
Sections 3.2, 4.2, and 5.3 
Document A 
Document A 
June 2004 No. 4: Mechanical Degradation 
Table 1-1. Key Technical Issue AgreementsRelevant Section Supporting Documents 
KTI Agreement 
RDTME 3.05 
RDTME 3.06 
RDTME 3.15 
RDTME 3.16 
RDTME 3.17 
RDTME 3.19 
RDTME 3.02 
RDTME 3.10 
RDTME 3.13 
RDTME 3.04 
RDTME 3.08 
RDTME 3.12 
RDTME 3.09 
RDTME 3.11 
Description 
Technical basis for accounting for effects of 
lithophysae 
Design sensitivity and uncertainty analyses of 
the rock support system 
Modeling joint planes as circular discs, small 
trace length fractures 
Technical basis for effective max rock size 
Determine whether rockfall can be screened 
from PA abstractions 
Critical combinations of in situ, thermal, and 
seismic stresses 
Two-dimensional modeling of emplacement drifts 
Boundary conditions, discontinuum versus 
continuum modeling 
Site-specific properties of the host rock 
Design sensitivity and uncertainty analyses of 
fracture patterns 
Dynamic analyses of ground support 
performance 
Rock movements in the invert 
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Revision 1 Revision 1 
about 20 years after closure. The temperature then slowly decreases over time, with the 
emplacement drift remaining above 96C for approximately 1,000 years (BSC 2004a). 
After closure, the ground support will corrode and lose its function over the period of perhaps 
decades to centuries, resulting in unsupported emplacement drifts through the large majority of 
the postclosure period. 
During the postclosure period, the emplacement drifts will be subjected to the following loading 
conditions (Figure 1-6): the in situ gravitational/tectonic stress state, transient thermally induced 
stresses due to expansion of the rock mass3, and seismic stresses due to potential earthquake 
shaking. Additionally, the rock mass strength, particularly in lithophysal rock, will exhibit a 
degree of time-dependency as a result of typical stress corrosion mechanisms4. The impact of 
these in situ and induced stress components is the potential for rock mass yield in the immediate 
region around the tunnels and some degree of rockfall. The rockfall has potential performance 
impacts on the following systems (BSC 2004e): 
 Mechanical Effects on Engineered Barriers 
- Effects include possible rock particle impacts and dynamic loading of the drip shield, 
either under gravitational or earthquake-induced accelerations. 
- Quasi-static loading and contact from rock particles resting on the drip shield 
following rockfall. 
- Dynamic loading of the drip shield could occur from seismic ground motions applied 
to a previously collapsed drift. 
 Mechanical Effects on In-Drift Environment 
- Rockfall of sufficient volume could result in an insulating blanket surrounding drip 
shields, impairing heat transfer to the rock mass and increasing waste package 
temperatures. 
- Rock particles within the drift will potentially alter the in-drift moisture and chemical 
environment around and upon the engineered barrier system components, and 
therefore need to be accounted for in drip shield and waste package corrosion. 
 Mechanical Effects on Seepage of Groundwater Into the Drift 
- Rockfall will result in changes to the size and shape of the emplacement drifts as a 
function of time, thus altering the seepage flow to the drift. 
3 
4 
A large portion of the thermally induced strains and stress are recoverable as the rock mass cools over time. 
Stress corrosion is a term commonly used in the field of rock mechanics (and metals) that represents timedependent, 
subcritical crack growth that occurs when existing material flaws in the rock are subjected to stresses 
that are near the failure state of the material. This process, which occurs at a more rapid rate in the presence of 
moisture, may result in damage and yield at applied stresses that are less than the short-term strength. Corrosion 
here does not refer to corrosion of metals. 
June 2004 1-11 No. 4: Mechanical Degradation Revision 1 
- Rock particles within the drift will have impacts on the drift capillary strength and 
seepage transport mechanisms and travel paths within the drift depending on particle 
size distribution. 
- Damage around the drift itself will affect the rock mass hydraulic characteristics that 
affect the capillary barrier in the periphery of the drift. 
Figure 1-6. Potential Postclosure Performance Impacts of Rockfall on Engineered Barriers, In-Drift 
Environment, and Groundwater Seepage into Drifts 
A series of NRC KTI agreements deal specifically with the methodology for analysis of and 
estimation of rockfall size distribution and volume as a function of the variability of rock mass 
properties, loading conditions and strength time-dependency. A number of KTI agreements in 
other areas (waste package, drip shield, seepage) address the impacts of rockfall as these 
performance issues. 
June 2004 1-12 No. 4: Mechanical Degradation Revision 1 
1.2.3 Summary of Approach to Addressing Preclosure and Postclosure Issues 
The primary technical issues that have formed the basis for the mechanical degradation 
resolution strategy can be summarized in three categories: (1) geotechnical characterization and 
rock mass property definition, (2) development and validation of numerical modeling tools, and 
(3) design and performance analyses. 
1. Geotechnical Characterization and Rock Mass Property Definition 
 Development of laboratory and in situ testing database of thermal-mechanical and 
time-dependent material properties for intact rock and fractures 
 Development of a relationship between geologic structure (fractures in the 
nonlithophysal rocks and lithophysae in the lithophysal rocks) and rock mass 
material response 
 Determination of the impact of the variability of geologic structure on rock mass 
property size effect, variability, and uncertainty. 
2. Development and Validation of Numerical Modeling Tools (determination or 
development of appropriate models for sensitivity studies of excavation stability and 
rockfall under in situ, thermal, and seismic loading) 
 Continuum versus discontinuum modeling 
- Discontinuum modeling required for representing rockfall 
- Continuum modeling suitable for parameter studies 
 Two- versus three-dimensional models 
- Three-dimensional models required in joint-controlled nonlithophysal rock, 
where response to seismic load is generally anisotropic 
- Two-dimensional models sufficient in generally isotropic lithophysal rock 
3. Design and Performance Analyses 
 Design studies 
- Verification, via empirical and numerical analyses, of the functional and 
operational requirements and specification of ground support 
- Develop initial ground support designs and specifications based on 
requirements 
- Verify design concepts via analysis of rock mass stability in lithophysal and 
nonlithophysal rock for in situ stress and preclosure thermal and seismic 
loading 
1-13 June 2004 No. 4: Mechanical Degradation Revision 1 
- Development of observation and maintenance plans for ground support 
 Performance analyses 
- Stability analysis and rockfall estimate under in situ, thermal, and seismic 
loading 
- Assessment of long-term stability of tunnels under quasi-static loading and 
time-dependent rock mass strength degradation 
The approach taken to address these issues (Figure 1-7) is based on development of a 
progressively more detailed understanding of the mechanical behavior of the lithophysal and 
nonlithophysal rock masses, starting with basic geologic characterization to understand rock 
mass variability, followed by laboratory and in situ testing with closely coupled numerical model 
development and validation. The process was developed from laboratory to field scale, allowing 
testing of the models and development of confidence in their ability to predict complex 
degradation modes. Next, the validated site-specific models were used to conduct sensitivity 
studies to examine thermal and mechanical drift degradation for variation in rock mass structure 
(and properties) under static and transient loading. The KTI agreements that have been 
addressed at each stage in this process are also given in this figure. 
The process was composed of six major program work elements. The approach initially 
involved developing a detailed understanding of the thermal-mechanical properties and 
variability of lithophysal and nonlithophysal rocks, and developing validated numerical models 
that can be used for design and performance assessment. The outcome of this process was a set 
of material models and properties and their ranges that were used as input to sensitivity studies of 
drift degradation in response to seismic events and to time-dependent processes. 
As described in Section 3, the mechanical and thermal properties of the lithophysal rocks are 
sensitive primarily to the lithophysal porosity. Because the lithophysal cavities range in size 
from millimeters to over a meter, and are of widely different shape and distribution within the 
rock mass, sampling and mechanical testing of sufficiently large samples presents a challenge, 
and thus characterization of properties and property variability is an issue. 
To overcome this issue, mechanical testing at increasing scales was used to understand porosity 
and size effects. Calibrated numerical modeling using discontinuum methods were used as a 
means of exploring the variability of lithophysal porosity, size, shape, and distribution on failure 
mechanisms and mechanical property variations. Thus, testing was used to define the basic 
lithophysal rock mass properties and was supplemented by modeling to explore variability 
ranges. 
The sensitivity modeling was aimed at estimating tunnel stability in the preclosure and 
postclosure time frames. In the preclosure period, the primary issue was the verification of 
ground support methods. In the postclosure period, the primary issues were estimation of 
mechanical degradation of emplacement drifts from thermal and seismic loading, or timedependent 
degradation response of the rock mass. 
June 2004 1-14 No. 4: Mechanical Degradation Revision 1 
studies. PSHA = probabilistic seismic hazard analysis. 
NOTE: Process starts with compilation and analysis of basic geotechnical mapping, followed by laboratory and field 
testing and model validation to develop rock mass property estimates for design and performance sensitivity 
Figure 1-7. General Approach to Resolution of the Repository Design and Thermal-Mechanical Effects 
Key Technical Issues 
A review of the six major program work elements is as follows: 
1. Geotechnical CharacterizationFurther analysis of the extensive, existing rock mass 
geologic and geotechnical characterization data from the ESF and the Enhanced 
June 2004 1-15 No. 4: Mechanical Degradation Revision 1 
Characterization of the Repository Block (ECRB) Cross-Drift, as well as surface 
outcrops and boreholes, to estimate the geometrical variability of rock mass fracturing 
and lithophysae. These data are supplemented by new detailed panel mapping of 
lithophysae in the ECRB Cross-Drift. Detailed statistical analysis of the fracture 
geometry in the middle nonlithophysal unit of the Topopah Spring Tuff (Tptpmn) and 
the lithophysae geometry in lower lithophysal unit of the Topopah Spring Tuff (Tptpll) 
is performed to provide the basic rock mass structural input and its variability to the 
modeling and analysis activities (see Section 2.3). 
2. Laboratory Testing and Model CalibrationA large amount of data for thermal and 
mechanical rock properties exists for nonlithophysal rocks. These data have been 
supplemented with laboratory direct shear testing of representative fracture samples. 
Additional compression tests on nonlithophysal samples with varying size and 
saturation levels, as well as static fatigue testing to derive time-dependent strength 
properties, have been conducted. Laboratory tests on lithophysal rock cores with a 
large diameter (290 mm) have been conducted to determine mechanical and thermal 
properties as a function of porosity, temperature, and saturation level. These 
laboratory tests, combined with in situ testing and previous large diameter (267-mm) 
cores, are used to establish the basic engineering properties of the lithophysal and 
nonlithophysal rock masses and their variability with porosity (see Sections 3.2.1, 
4.2.2, and 4.2.3). 
The lithophysal laboratory data (i.e., strength, modulus, and variation with porosity) 
and observations of failure mechanisms are used as the basis for initial calibration of 
discontinuum numerical models5. Two discontinuum numerical approaches are used 
to back-analyze laboratory data and to provide an understanding of the impact of 
lithophysal cavities on deformability and yield mechanisms. These approaches can 
represent the presence of voids as well as the physical reality of complex failure 
mechanisms involving interlithophysae fracturing and compaction of void space. 
PFC2D and PFC3D, which use a micromechanical discontinuum approach for 
representing rock, and UDEC, a standard discontinuum modeling program, are used to 
reproduce stress-strain behavior and failure mechanisms observed from the laboratory 
experiments. The models are also calibrated against static fatigue testing, which 
determines a relationship between the time-to-failure and applied stress level for 
nonlithophysal cores. These data are used to develop a basic mechanistic approach for 
predicting time dependency that is based on a standard stress corrosion crack growth 
model. 
5 
3. In Situ Material Properties Testing and Model ValidationIn situ mechanical and 
thermal testing of lithophysal rocks is performed to determine the size effect (porosity 
and fracture) on rock mass constitutive behavior. These tests are further used for 
Discontinuum numerical models simulate rock masses as actual blocks of solid material separated by fracture 
surfaces that have cohesive tensile and frictional strength. When the rock mass yields, the fractures may 
physically break in shear or tension, creating individual blocks that are free to detach from the surrounding rock 
mass. Continuum numerical models represent the rock mass as a continuous body in which the stiffness and 
yielding effects of fractures are represented via material stress-strain relationships. The rock mass cannot 
represent detachment of individual blocks from the parent rock mass. 
June 2004 1-16 No. 4: Mechanical Degradation Revision 1 
validation of the numerical modeling approaches at increasing size scales. This results 
in numerical modeling approaches that can be used with confidence for extrapolating 
the mechanical response of lithophysal rocks (see Sections 3.2.3 and 4.2.7). 
4. Model Extrapolation for Establishing Property VariabilityDue to the size effect 
on thermal and mechanical properties introduced by lithophysae, it is necessary to test 
large sample sizes (for lithophysal units) to obtain realistic in situ rock mass 
properties. It is impractical to perform a statistically large number of in situ tests to 
determine strength and deformability variability. However, the numerical models, 
suitably calibrated against the laboratory and field data, provide the capability of 
exploring the impact of lithophysal variations such as porosity, lithophysal shape, size 
distribution, and spacing on strength and deformability variation. These models have 
been used to provide an understanding of the basic mechanics of the lithophysal rock 
mass and to define the resulting input property ranges for subsequent design and 
performance calculations. The models are also used as a means of understanding the 
impact of lithophysal porosity on stress redistribution within the rock mass and thus its 
impact on stress-related time dependency (see Section 4.2.3.3). 
5. Best-Estimate Constitutive Models and Property Ranges for Sensitivity Studies 
The testing data and numerical extrapolations are used to define the constitutive6 
models and property ranges for the lithophysal rock mass. The lithophysal rock mass 
properties are subdivided into five categories. These categories are based on 
lithophysal porosity level and are indicative of the in situ rock mass strength and 
moduli variation. Laboratory data from direct shear testing and field characterization 
of joints are used to define property ranges for modeling of nonlithophysal rocks (see 
Section 4.2.4). 
6 
6. Design and Performance Assessment Sensitivity StudiesSensitivity studies of 
excavation stability under in situ, thermal, and seismic loads in the preclosure and 
postclosure time frames have been performed. Numerical model sensitivity studies are 
performed using the range and distribution of rock mass properties determined from 
testing and model extrapolation. The deformation and yield of the rock mass around 
the openings in preclosure time are used to define ground support requirements and 
appropriate support methods. These models are supplemented with practical empirical 
ground support specification methods, where applicable, to determine ground support 
methods and materials that will require minimal or no maintenance over the preclosure 
period. Extensive discontinuum modeling studies are used to examine thermal and 
seismic loading effects on mechanical degradation of the emplacement drifts in the 
postclosure. In nonlithophysal rocks, analysis of the jointing is used to generate a 
large number of three-dimensional, stochastically defined fracture geometries that 
realistically represent the rock mass fracturing documented from field mapping. These 
jointing realizations and associated emplacement drifts are subjected to postclosure 
A constitutive model in this case is a set of mathematical equations that describe the relationship of applied stress 
to strain for the rock mass. These equations include a description of the elastic and failure characteristics of the 
rock mass. Typically, rock is described by yield criteria that account for the strength of the fracture fabric of the 
rock mass. A typical form of yield criteria for rock is the Mohr-Coulomb or Hoek-Brown criteria (e.g., 
Hoek 2000). 
June 2004 1-17 No. 4: Mechanical Degradation Revision 1 
rock mass temperature levels and ground motions representative of seismic events 
with 10-5, 10-6, and 10-7 annual exceedance frequencies. The rock mass is also 
subjected to predicted postclosure temperature distribution. Two-dimensional models 
representing the range of strength categories of the lithophysal rocks are subjected to 
similar thermal and seismic loading. The rockfall generated by these simulations is 
presented in a number of ways, including the size distribution of the particles, the peak 
velocities of each particle and the total mass of rock dislodged from the surrounding 
mass. This information is used to define dynamic and quasi-static loads applied to the 
drip shield as well as the resulting shape of the tunnels (see Section 5.3). 
Finally, the models are used to predict the time-related degradation in the lithophysal 
rocks resulting from quasi-static loading and time-dependent strength reduction. 
Sensitivity studies for the range of rock mass strength categories are used to predict 
the change in shape of the emplacement drifts and the total dislodged rock volume as a 
function of time. This information provides a basis for the estimation of in-drift and 
seepage effects due to mechanical degradation (see Section 5.3.2.2.4). 
A detailed discussion of the methodology for inclusion of drift degradation estimates into the 
total system performance assessment is given in Seismic Consequence Abstraction (BSC 2004e). 
The drift degradation estimates, including rockfall and drift size and shape, are considered in 
both nominal and seismic performance scenario classes7. The nominal scenario class 
considers the estimated time-dependant change in drift shape, and the volume of associated 
rockfall on and around the drip shield is considered. The mechanical damage to the drip shield 
resulting from quasi-static loading of dislodged rock is considered, as well as the impacts of the 
rockfall on heat transfer mechanisms. The seismic scenario class takes into account the rockfall 
induced by possible seismic events. Here, the estimates of rockfall onto the drip shield as a 
function of the annual exceedance frequencies are used to estimate the drip shield damage level 
and the subsequent effects on seepage of ground water through it. Additionally, dynamic loading 
of the drip shield due to seismic shaking of a previously failed, rubble-filled drift are also 
considered. Although not specifically part of the drift degradation effects, the seismic scenario 
also includes estimation of performance consequences of vibratory motion on waste package 
damage resulting from waste package impacts with the pallet, adjacent waste packages, and the 
drip shield. 
1.3 ORGANIZATION OF THE REPORT 
The report format is as follows: 
 Introduction (Section 1)The objectives and scope of this technical basis document, 
and a summary of the mechanical degradation issues and issue resolution strategy. 
7 The TSPA model takes into account nominal and disruptive events scenario classes. The disruptive events 
scenario class includes both seismic and igneous scenarios. The nominal scenario class includes FEPs expected 
to occur over the life of the repository, including seismic events with annual probabilities in the range of 10-4 and 
10-5 per year. The seismic scenario class includes higher consequence, lower probability events with annual 
probabilities in the range of 10-6 to 10-8 per year. 
1-18 June 2004 No. 4: Mechanical Degradation Revision 1 
 Geologic Characteristics of the Repository Host Horizon Relevant to Mechanical 
Degradation (Section 2)The geology of the Topopah Spring Tuff as it relates to 
understanding of the thermal-mechanical rock mass properties and the variability 
introduced by fracturing and lithophysae. 
 Thermal-Mechanic Rock Properties DatabaseReview of Laboratory and In Situ 
Testing for the Nonlithophysal and Lithophysal Rock Masses (Section 3)A review 
of the results from nearly 20 years of laboratory and field testing programs for 
lithophysal and nonlithophysal rocks. 
 Development of Rock Mass Material Modeling Approaches for Nonlithophysal and 
Lithophysal Rocks (Section 4)The development of appropriate mechanical material 
models for lithophysal and nonlithophysal rocks. The generalization of these data into 
modeling approaches for mechanical degradation studies in lithophysal rocks. 
Calibration and validation of discontinuum numerical models against these data and 
extrapolation studies using these models to establish variability in mechanical properties 
of the lithophysal rocks. 
 Analysis of Preclosure and Postclosure Drift Mechanical Degradation under 
Gravitational, Thermal, and Seismic Loading (Section 5)The details of numerical 
parameter 
studies of mechanical degradation under various loading conditions and time 
dependent strength loss. The output from these studies to other disciplines. 
 Summary and Interactions with Engineered and Natural Systems (Section 6)A 
summary of the conclusions and information feeds to other postclosure performance 
systems. 
 References (Section 7)Sources of information used in this document. 
 AppendicesAddress specific RDTME KTI agreements. 
1.4 NOTE REGARDING THE STATUS OF SUPPORTING TECHNICAL 
INFORMATION 
This document was prepared using the most current information available at the time of its 
development. This technical basis document and its appendices providing KTI agreement 
responses that were prepared using preliminary or draft information reflect the status of the 
Yucca Mountain Projects scientific and design bases at the time of submittal. In some cases this 
involved the use of draft analysis and model reports and other draft references whose contents 
may change with time. Information that evolves through subsequent revisions of the analysis 
and model reports and other references will be reflected in the license application as the 
approved analyses of record at the time of license application submittal. Consequently, the 
Yucca Mountain Project will not routinely update either this technical basis document or its KTI 
agreement appendices to reflect changes in the supporting references prior to submittal of the 
license application. 
June 2004 1-19 No. 4: Mechanical Degradation INTENTIONALLY LEFT BLANK 
1-20 No. 4: Mechanical Degradation 
Revision 1 
June 2004 Revision 1 
2. GEOLOGIC CHARACTERISTICS OF THE REPOSITORY HOST HORIZON 
RELEVANT TO MECHANICAL DEGRADATION 
June 2004 
2.1 GENERAL GEOLOGY OF REPOSITORY HOST HORIZON 
The lithostratigraphy and geologic evolution of the Yucca Mountain site is described in Yucca 
Mountain Site Description (BSC 2004f). This section gives a summary of the geologic 
characteristics of the repository host horizon relevant to the analysis of mechanical degradation 
response of the repository excavations. 
Site-specific characteristics of the rock units of the Topopah Spring Tuff that constitute the host 
rock at the repository horizon are found in the geologic mapping of those units in both the main 
drift and ramps of the ESF and the ECRB Cross-Drift. 
The locations of the ESF and ECRB Cross-Drift, and the lithostratigraphic units exposed by the 
tunnels, are illustrated in the geologic cross section (Figure 2-1). The units that comprise the 
host rocks of the repository horizon are zones of the crystal-poor member (Tptp) of the Topopah 
Spring Tuff. The host rocks are shown schematically in Figure 2-2. In descending order (by 
depth), the host rocks consist of the lower part of the upper lithophysal zone (Tptpul), the middle 
nonlithophysal zone (Tptpmn), the lower lithophysal zone (Tptpll), and the lower nonlithophysal 
zone (Tptpln). 
The repository host rock units can be categorized into two general engineering classifications: 
nonlithophysal units (Tptpmn and Tptpln) and lithophysal units (Tptpul and Tptpll), based on the 
relative proportion of lithophysal cavities. 
The nonlithophysal units are generally hard, strong, fractured rocks with matrix porosities of 
10% or less. Fractures that formed during the cooling process are the primary structural features 
found in these units. In contrast, the lithophysal units have significantly fewer fractures of 
significant continuous length (i.e., trace length greater than 1 m), but have relatively uniformly 
distributed porosity in the form of lithophysal cavities. Lithophysal porosity in the Tptpul and 
Tptpll is generally on the order of less than 10% to about 30% by volume. The groundmass that 
makes up the rock matrix in the lithophysal units is mineralogically the same as the matrix of the 
nonlithophysal units, but is heavily fractured with small scale (lengths of less than 1 m) 
interlithophysal fractures in the Tptpll; however, it is relatively fracture-free in the Tptpul. 
2-1 No. 4: Mechanical Degradation nonlithophysal), Tptpll (lower lithophysal), and Tptpln (lower nonlithophysal). 
Revision 1 
June 2004 
Source: Mongano et al. 1999, Drawing 06-46-345 and 0A-46-345. 
NOTE: Repository host horizon includes the major subunits of the Topopah Spring Formation, designated the Tptpul (upper lithophysal), Tptpmn (middle 
Figure 2-1. Geologic Cross Section at the Location of the Enhanced Characterization of the Repository Block Cross-Drift 
2-2 No. 4: Mechanical Degradation Revision 1 
Source: Buesch et al. 1996, Appendix 2; Mongano et al. 1999, pp. 12 to 43. 
Figure 2-2. Structure of the Topopah Spring Tuff Showing the Relative Relationship between 
Fracturing and Lithophysae in the Major Flow Subunits 
June 2004 
2.2 UNDERGROUND REPOSITORY LAYOUT AND LITHOLOGIC 
INTERSECTIONS 
The repository consists of four panels that will cover about 5 km2 within the Topopah Spring 
Tuff (Figure 2-3). The repository layout extends about 5 km in length (northsouth) with the 
widest part being about 2 km (eastwest). The total length of all excavated openings including 
the drifts, turnouts, exhaust mains, exhaust shafts and raises and other miscellaneous openings is 
about 110 km. The emplacement drifts comprise about 66 km of tunnels contained primarily 
within the Tptpll (less than or equal to 81%) and the Tptpmn (less than or equal to 12%). The 
remaining geologic units comprise roughly 7% (Tptpul about 4% and Tptpln about 3%) of the 
emplacement drift area. Overall, the nonlithophysal rocks comprise roughly 15% of the 
2-3 No. 4: Mechanical Degradation Revision 1 
emplacement area, whereas the lithophysal rocks comprise approximately 85% (BSC 2003e, 
Table I-2). 
Source: BSC 2004g. 
Figure 2-3. Overlay of the Lithostratigraphic Units on the Planned Repository Layout 
2.3 ENGINEERING CHARACTERISTICS OF THE ROCK MASS IMPORTANT TO 
MECHANICAL PERFORMANCE OF REPOSITORY EXCAVATIONS 
The rock matrix material is a typical fine-grained volcanic rock of silica content. The structure 
of the rock mass defines the properties and overall mechanical response of the rock mass to 
thermal and mechanical loading. The fracture geometry and properties in the nonlithophysal 
rocks and the degree of porosity (total and lithophysal) in the lithophysal subunits are the 
primary geologic structural features that impact rock mass behavior. 
Geotechnical mapping of fractures has been performed in the ESF and the ECRB Cross-Drift 
(CRWMS M&O 1998a; Mongano et al. 1999). Figure 2-2 presents a schematic of the Topopah 
Spring Formation illustrating the general occurrence of fracturing and lithophysae in the various 
subunits of the repository host horizon. The occurrence of fractures and lithophysae are roughly 
inversely proportional as demonstrated quantitatively in Figure 2-4, where the fracture density 
(fractures with trace length greater than 1 m), determined from detailed line mapping, and the 
approximate percentage of lithophysal porosity in the ECRB Cross-Drift are shown. The density 
June 2004 2-4 No. 4: Mechanical Degradation Revision 1 
of fractures with trace length greater than 1 m is significantly larger in the Tptpmn and Tptpln 
(20 to 35 fractures/10 m), as compared to 5 fractures/10 m or less in the Tptpul and Tptpll. 
Lithophysae, on the other hand, are sparse in the Tptpmn and Tptpln. 
Source: Mongano et al. 1999, Figure 13. 
NOTE: There is an inverse relationship between fracture density and lithophysal porosity. 
2.3.1 Characterization of Fractures 
Fracturing in the TptpmnThe field fracture database, constructed from mapping in the ESF 
and ECRB Cross-Drift during tunneling operations, consists of full-periphery maps and detailed 
line surveys of all fractures with length of one meter or greater. The full periphery maps consist 
of traced fracture lengths drawn directly on invert-to-invert surface maps of the tunnels and 
include the dip and dip direction of the feature and any intersections with other fractures. 
Detailed line surveys consist of mapping the detail geometry and surface characteristics of every 
fracture crossing a line hung along the springline of the tunnel. The details of this mapping are 
provided by Mongano et al. (1999). In total, a database of more than 35,000 fracture 
descriptions is available. Additionally, a small-scale fracture-mapping program was conducted 
to document the characteristics of fractures of less than 1-m length in the nonlithophysal and 
lithophysal rocks. 
In summary, there are four sets of fractures in the Tptpmn with the geometrical and surface 
characteristics provided in Table 2-1. 
Figure 2-4. Composite Plot of Fracture Frequency and Lithophysal Porosity as a Function of Distance 
along the Enhanced Characterization of the Repository Block Cross-Drift 
June 2004 2-5 No. 4: Mechanical Degradation Revision 1 
Table 2-1. General Characteristics of Fracture Sets in the Middle Nonlithophysal Unit 
Set 
1 
Mean 
Azimuth/Dip 
120/84 
2 215/88 
3 302/38 
Comment 
Rough to smooth, planar 
Smooth but curved 
Random fractures with generally 
flat to moderate dip 
Trace Length Median 
from Full Periphery 
Geologic Maps (m) 
3.3 
3.1 
3.6 
3.4 
Mean 
Spacing (m) 
0.48 
1.08 
3.40 
2.46 VPP 329/14 Random fractures with generally 
flat to moderate dip 
Source: DTNs for tunnel mapping include: GS960908314224.020, GS000608314224.006, 
GS960908314224.015, GS960908314224.016, GS971108314224.025, 
GS960708314224.008, GS000608314224.004, and GS960708314224.010. 
NOTE: Trace length medians are taken from a compilation of tunnel mapping and synthetic tunnel 
samples from FracMan. VPP = vapor phase parting. 
The fractures, particularly the high angle sets (sets 1 and 2), have mean trace lengths less than 
the diameter of the emplacement drifts (Figure 2-5), with ends that sometimes terminate either 
against other fractures or in solid rock. Thus, rather than having continuous joint sets, there is 
often a solid rock bridge between joint tracks. A photograph of a typical wall in the ECRB 
Cross-Drift (Figure 2-6), demonstrates an important aspect of the fracturing in the Tptpmn. The 
fracture traces were painted during the detailed line survey, as seen in this photo, and each 
fracture termination was logged as being against another fracture, within solid rock, or 
continuous. The photo shows the common occurrence of fractures that terminate in solid rock 
(T-junctions) as opposed to continuous structures (arrowheads). The subvertical fractures have 
smooth surfaces but often have curved surfaces with large-amplitude (dozens of centimeters) 
asperities and wavelength of meters (Mongano et al. 1999). 
Figure 2-7 shows that low angle vapor phase partings are relatively continuous structures seen 
throughout the Tptpmn. These continuous, but anastomosing fractures are subparallel to the dip 
of the rock unit, and are filled with concentrations of vapor-phase mineralization (primarily 
tridymite and cristobalite). The surfaces are rough on a small scale and, unlike the subvertical 
fractures, have cohesion as a result of the mineral filling. 
The nature of the fracture geometry governs the estimates of the stability and degradation of the 
nonlithophysal rock mass, particularly under the action of seismic shaking, as well as estimates 
of the support function and level of required ground support. Most rock mass classification 
schemes are based on experience of rock masses with continuous joint sets that create regular, 
blocky masses (e.g., Hoek 2000). In the Tptpmn, the relatively short trace lengths and 
nonpersistent joints create relatively few kinematically removable blocks. This is evidenced by 
the fact that only a small number of rock blocks have actually been dislodged in the ECRB 
Cross-Drift (BSC 2004a, Appendix F). They were either dislodged under the action of the TBM 
or were scaled out of the drift back and walls immediately after mining. There have been no 
reported keyblock failures in the ECRB Cross-Drift since excavation, even though only light 
bolting and meshing is used for ground support. 
June 2004 2-6 No. 4: Mechanical Degradation Revision 1 
Source: Mongano et al. 1999, Figure 14. 
Figure 2-5. Fracture Trace Length as a Function of Depth in the Enhanced Characterization of the 
Repository Block Cross-Drift and by Subunit of the Tptp from Detailed Line Surveys 
NOTE: T-junctions on fractures indicate terminations; arrowheads show continuous features. 
Figure 2-6. Fractures in Wall of the Enhanced Characterization of the Repository Block Cross-Drift in 
the Tptpmn 
June 2004 2-7 No. 4: Mechanical Degradation Revision 1 
Figure 2-7. Low-Angle Vapor-Phase Partings in Tptpmn 
Fracture geometry and surface characteristics are required for numerical analysis of mechanical 
degradation of emplacement drifts. Three-dimensional discontinuum numerical methods are 
used to simulate the degradation of the drifts under the action of in situ, thermal, and seismic 
stressing. These models require geometric input of fracture distributions that are statistically 
similar to those observed in the ESF and ECRB Cross-Drift. The mapping data (both full 
periphery and detail line surveys) are used as the basis for generation of a synthetic rock mass 
fracture distribution developed stochastically using the FracMan fracture generation program 
(USGS 1999). Section 4.1.1 discusses the use of FracMan for generation of a statistically 
equivalent fracture domain and the subsequent discontinuum modeling. 
Fracturing in the TptpllShort-length fractures (less than 1-m trace length), coupled with the 
lithophysae, are the most important features that govern stability in the Tptpll. Whereas the 
Tptpul tends to have sparse, small-scale, interlithophysal fracturing (Figure 2-8a), the Tptpll has 
abundant fracturing (Figure 2-8b). The fractures, existing throughout the Tptpll, have a primary 
vertical orientation and have spacing of a few centimeters. 
Thin-section analyses of the fracturing in the Tptpll and the Tptpmn show vapor phase 
alterations on most of the fracture surfaces within the rock mass away from the tunnel wall, 
indicating they are natural fractures (i.e., not mining-induced) that were formed during the 
cooling process (BSC 2004a, Section 6.1.4.1). Near the tunnel wall, it is clear that some of the 
fractures have been disturbed by mining, and that at a small number of locations, new, stressinduced, 
wall-parallel fractures have been created in the immediate springline of the tunnel. 
These stress-induced fractures are observed to a depth of about 0.5 m in some of the large 
diameter (290 mm) core holes drilled in the springline area for rock mechanics testing purposes. 
June 2004 2-8 No. 4: Mechanical Degradation Revision 1 
No matter the origin of the fracturing, it is clear that the Tptpll has a ubiquitous fracture fabric 
that is evident in panel mapping or when large diameter core is removed from boreholes 
(Figure 2-9). The matrix material, although strong and similar to the Tptpmn, has numerous 
fracture surfaces that tend to separate during the drilling process. The result is breakage into 
small blocks, making removal of large lengths of core (and thus laboratory testing of sufficiently 
large samples) very difficult. These fractures, which interconnect the lithophysae, tend to create 
blocks with dimensions on the order of about 10 cm or less on a side. Longer length fractures 
that cut the entire drift are widely separated and have been found to be incapable of producing 
kinematically possible wedges (BSC 2004a). Therefore, the potential mode of failure within the 
Tptpll under seismic or time-dependent yield will be in a raveling mode that creates small block 
sizes. 
Emplacement drift ground support for preclosure in this type of rock mass requires a more-orless 
continuous surface confinement that prevents any loosening of these small blocks. The 
design solution developed for support of this rock mass is a thin, continuous, perforated steel 
sheeting that is bolted directly to the emplacement drift surface (see Figure 1-4). Calculations 
indicate that large blocks do not form in the Tptpll (BSC 2004a, Section 6.4.3). The combination 
of lithophysae and fractures in this zone tend to create small blocks with dimensions on the order 
of about 10 cm or less on a side (BSC 2004a, Section 6.1.4.1). Blocks of this size are not 
capable of breaching the drip shield or waste package under dynamic loading. 
June 2004 2-9 No. 4: Mechanical Degradation NOTE: The Tptpul (a) is characterized by a relatively unfractured matrix between lithophysae, whereas the Tptpll (b) 
is abundant in natural, short-length fractures that interconnect lithophysae. Spacing of the fractures is 
generally less than 5 cm. Hackly surface in (a) is a result of TBM cutting process. 
Figure 2-8. Comparison of Lithophysae and Fracturing in the Tptpul and Tptpll 
2-10 No. 4: Mechanical Degradation 
Revision 1 
June 2004 Revision 1 
Source: BSC 2004a, Figure O-3. 
NOTE: Lithophysae have red L identifiers with cavities outlined in red and rims in green. Spots have blue S 
identifiers with cyan outlines. Lithic clasts have orange C identifiers with gold outlines. The panel shown is 
1  3 m. Geometry of lithophysae, rims, and spots are scaled from the photographs and used to determine 
porosity, size, and shape distributions. Lithophysal porosity averages about 20% within the Tptpll. 
Figure 2-9. Lithophysae, Spots, and Clasts of Tptpll (as Discussed Below) in Panel Map 1493 Located 
on the Right Rib from Station 14+93 to 14+96 
2.3.2 Characteristics of Lithophysae 
Although the character of the lithophysae varies between the Tptpul and Tptpll, the mineralogy 
of the matrix material within both of these units is the same as in the nonlithophysal units. 
The lithophysae in the Tptpul (BSC 2004a): 
 Tend to be smaller (roughly 1 to 10 cm in diameter) 
 Are more uniform in size and distribution within the unit 
 Vary in infilling and rim thicknesses 
 Have a volume percentage that varies consistently with stratigraphic position 
 Are stratigraphically predictable. 
June 2004 2-11 No. 4: Mechanical Degradation Revision 1 
In contrast, the lithophysae in the Tptpll tend to be highly variable in size, ranging from roughly 
1 cm to 1.8 m in diameter (BSC 2004a). They also: 
 Have shapes that are highly variable from smooth and spherical to irregular and sharp 
boundaries 
 Have infilling and rim thickness that vary widely with vertical and horizontal spacing 
 Have volume percentages that vary consistently with stratigraphic position 
 Are stratigraphically predictable. 
A detailed study of the lithostratigraphic features in the lower lithophysal zone exposed in the 
ECRB Cross-Drift has recently been completed. These data are summarized in Drift 
Degradation Analysis (BSC 2004a), and a representative plot is given in Figure 2-9. The data 
package documents the distributions of size, shape, and abundance of lithophysal cavities, rims, 
spots, and lithic clasts, and these data can be displayed and analyzed as (1) local variations, (2) 
along the tunnel (a critical type of variation), and (3) as values for the total zone. 
In addition to the variation in the abundance of features, such as lithophysae along the ECRB 
Cross-Drift, there are variations in the size, shape, and distance between features. These types of 
variations are most easily observed with panel map data8 (Figure 2-9), which have been 
converted into porosity variations as functions of distance along the ECRB Cross-Drift, through 
the entire Tptpll subunit (Figure 2-10). This information provides direct input to mechanical 
degradation studies in the following ways: 
 The panel maps and porosity, size, and shape variations of lithophysae provide the basis 
for numerical estimation of impact of lithophysae on rock mass properties. 
 Rock mass properties in the lithophysal rocks are primarily a function of porosity, and 
the variation in porosity provided by the direct panel mapping allow the variation in rock 
mass properties to be estimated. 
8 A total of 18 1-by-3-m panel maps were developed in the ECRB Cross-Drift from top to bottom of the Tptpll. 
June 2004 2-12 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure O-15. 
NOTE: Porosity of the 5-m averaged large-lithophysae inventory is not included in the total. 
Figure 2-10. Calculated Porosity of Lithophysal Cavities, Rims, Spots, Matrix-Groundmass, and the 
Total Porosity in the Tptpll Exposed along the ECRB Cross-Drift 
This information has been used to develop a model for simulation of the spatial variability of 
lithophysal porosity within the Tptpll (BSC 2004a, Appendix T). The lithophysal porosity 
within the Tptpll follows a general stratiform geometry in a fashion similar to the occurrence of 
lithostratigraphic contacts within the overall Topopah Spring Tuff (i.e., the Tptpmn, Tptpll, etc.). 
The ECRB Cross-Drift transects the entire Tptpll unit, allowing determination of the stratiform 
nature of laterally continuous subzones of lithophysal porosity within the Tptpll. The lateral 
continuity of lithostratigraphic features and the projection of these features along the apparent 
dip of the ECRB Cross-Drift has been used to develop a vertical cross section of the distribution 
of lithophysal porosity within the Tptpll. Figure 2-11 presents two simulated vertical projections 
of lithophysal porosity for 50-m-tall vertical cross sections through the top and lower portions of 
the Tptpll. Each of the cells of these cross sections represent a volume of the Tptpll and have a 
lithophysal porosity associated with it. As seen in these cross sections, the lithophysal porosity 
occurs as stratiform subzones, with the highest values (i.e., 20% or higher) occurring in thin 
bands near the top of the subzone. The lowest lithophysal porosities occur in the lower portions 
of the subzone near the contact with the Tptpln. These simulated cross sections are used in 
Section 5.3 of this report as a basis for examination of the impact of spatial variability on rock 
mass mechanical response. 
June 2004 2-13 No. 4: Mechanical Degradation Revision 1 
No. 4: Mechanical Degradation 2-14 
Source: BSC 2004a, Figure T-5. 
NOTE: Cross section A is a 50 x 200-cell table representing a 1 x 1-m grid, and cross section B is a 20 x 80-cell table representing a 2.5 x 2.5-m grid for the 
simulated section at 17+56 in the ECRB Cross-Drift. Cross sections C and D represent simulated section at 20+14 in the ECRB Cross-Drift. Two 50 x 
200-m Simulated Cross Sections (Aspect Ratio is not 1:1) at Upper (A/B) and Lower (C/D) Sections of the Tptpll. 
June 2004 
Figure 2-11 Illustration of the Process of Sampling and Modeling Spatial Variability Using Lithophysal Porosity Revision 1 
Stress Component 
2.3.3 In Situ Stress State 
The in situ stress state at and in the vicinity of the Yucca Mountain site has been determined by 
hydraulic fracturing by Sandia National Laboratories (CRWMS M&O 1997a) and by Stock et al. 
(1985). A summary of the measurements is given in Table 2-2. 
Table 2-2. In Situ Stress Estimates at Yucca Mountain Site 
1 
Magnitude (MPa) 
~0.023 ~ depth (m) 
2 0.617 ~ 1 
Direction 
Vertical 
N15E 
N105E 3 
Source: DTN: SNF37100195002.001. 
2.4 SUMMARY 
The rock mass that comprises the repository host horizon consists of a number of subunits of the 
Topopah Spring Tuff. From an engineering and mechanical degradation standpoint, these 
subunits can be divided into two broad classifications: lithophysal and nonlithophysal rocks. The 
matrix material of both is similar mineralogically and mechanically, with the distinguishing 
characteristics, from a geomechanics standpoint, being their structural features (i.e., the fractures 
and lithophysae). Due to the importance of these structural features in the mechanical response 
of the rock, it has been necessary to perform detailed geologic and statistical descriptions of the 
geometric characteristics of the fractures and lithophysae. This information has been used to 
provide the basis for input to engineering design and performance modeling of these units, as 
described in Section 5.3. 
0.362 ~ 1 
2-15 June 2004 No. 4: Mechanical Degradation INTENTIONALLY LEFT BLANK 
2-16 No. 4: Mechanical Degradation 
Revision 1 
June 2004 Revision 1 
3. THERMAL-MECHANICAL ROCK PROPERTIES DATABASE 
REVIEW OF LABORATORY AND IN SITU TESTING FOR THE 
NONLITHOPHYSAL AND LITHOPHYSAL ROCK MASSES 
3-1 
3.1 INTRODUCTION 
This section of the report reviews the rock mass properties database for repository host horizon 
units. Two documents provide a detailed review and analysis of this information: Yucca 
Mountain Site Geotechnical Report (CRWMS M&O 1997b) and Subsurface Geotechnical 
Parameters Report (BSC 2003a). The former presents field and laboratory test data available 
through June 1996, while the latter presents the geomechanical data acquired since that time. 
The original report presents data primarily from small-diameter cores and thus provided little 
information on more representatively sized samples of lithophysal rock. A major laboratory and 
field testing effort to obtain data on lithophysal rocks was conducted in 2002 and 2003. This 
section presents an overview of the geomechanical database, and is organized in terms of rock 
type. 
To perform estimates of the mechanical degradation of emplacement drifts subject to stresses 
and time dependent material property changes, it is necessary to determine the basic thermal and 
mechanical properties of intact rock and the rock mass and to estimate their variability within the 
repository host horizon. The properties and geotechnical characteristics that need to be 
determined are: 
. Strength properties (T 
 Mechanical Properties of Intact Rock 
. Elastic moduli (E) and Poissons ratio (v) 
o) (from uniaxial, triaxial compression, and tensile testing). 
 Mechanical Properties of Joints 
. Normal (Kn) and Shear (Ks) stiffness 
. Shear strength (s) 
. Surface roughness (Jr) and dilation angle. 
 Thermal Properties 
. Conductivity (k) 
. Heat capacity (Cp) 
. Expansion coefficient (). 
 Geotechnical Characterization 
. Fracture geometry statistics (dip, dip direction, spacing, length) 
. Fracture surface characteristics (roughness, planarity) 
. Classification indices (Q, RMR, GSI). 
Figures 3-1 and 3-2 illustrate the approaches taken to defining rock mass thermal-mechanical 
constitutive models and properties for design and performance assessment studies. Laboratory 
testing of small cores from surface-based boreholes (diameter of 50 mm) and large cores from 
drilling within the ESF and ECRB Cross-Drift and Busted Butte (diameter of 305 mm and 267 
mm, respectively), as well as field geotechnical characterization, form the bases for initial 
June 2004 No. 4: Mechanical Degradation Revision 1 
properties estimates and materials model definition. For nonlithophysal rocks (Figure 3-1), 
which are typically strong and fractured volcanic materials, a standard geomechanical approach 
to rock properties estimation and numerical model definition is taken (e.g., Hoek 2000). 
Laboratory testing and ESF and ECRB Cross-Drift field mapping, using geotechnical 
characterization methods, are used to perform estimates of the in situ rock mass properties for 
use in ground support design parameter studies. Rockfall analyses require that the rock mass be 
modeled as a discontinuum material with explicitly defined fractures; therefore, for these studies, 
a detailed, statistical description of the rock mass fracturing and fracture properties is required 
for direct input to three-dimensional numerical models. The FracMan program (USGS 1999) is 
used for the fracture geometric modeling and field estimates, and direct shear testing of large 
joints in the laboratory is used for estimating surface properties. 
Definition of the material properties of, and mechanical constitutive modeling approach for, 
lithophysal rocks requires a different approach (Figure 3-2). The use of empirical techniques for 
estimating rock mass properties is not particularly applicable to these rock masses since there is 
little available experience in excavating similar rock types. Due to the presence of the 
lithophysal cavities, the rock mass properties are also both porosity- and size-dependent. In 
other words, the properties of the material are a function of both the size of the sample being 
tested, and the size, shape and degree of lithophysal cavities that the sample contains. 
Figure 3-1. Development Strategy for Rock Properties Database and Modeling Strategy for 
Nonlithophysal Rocks 
June 2004 3-2 No. 4: Mechanical Degradation Revision 1 
Figure 3-2. Development Strategy for Rock Properties Database and Modeling Strategy for Lithophysal 
Rocks 
To address these issues, an approach is taken based on laboratory and field-scale testing at 
increasingly larger scales to define the size effect and to provide data for model calibration and 
validation. An integral aspect of this approach is the comparison and validation of discontinuum 
numerical models that can simulate the basic mechanical response of the lithophysal tuff to 
applied stresses. 
Once it is demonstrated that the model can account for the observed response, it is used for 
conducting parametric compression testing studies on simulated samples of lithophysal tuff to 
investigate the impact of variability of lithophysae shape, size, and porosity on the range of 
expected mechanical properties. In this manner, the numerical models are used to extend the 
field and laboratory testing and establish the range of material properties for design and 
performance studies. This technique, combined with the laboratory testing, is used to establish 
the upper and lower bounds of lithophysal rock properties. In addition, the rock mass properties 
of large-scale samples of Tptpll, taking into account the impact of spatial variability of 
lithophysal porosity, are examined numerically. Comparison of the mechanical properties ranges 
to in situ ESF and ECRB Cross-Drift fracturing observations and thermal stress-induced spalling 
in the Drift Scale Test (DST) are used as validation tests of the model and properties ranges (see 
Section 4.2.7). All of these studies are used to define the base set and range of rock mass 
properties for use in performance assessment. 
3.2 GEOMECHANICAL PROPERTIES OF YUCCA MOUNTAIN ROCK 
3.2.1 Mechanical Intact Rock Properties of Nonlithophysal and Lithophysal Rocks 
In the late 1970s and through the mid-1980s, many samples were tested, from units extending 
from the upper-most parts of the Paintbrush Tuff down through the lower regions of the Crater 
Flat Tuff. From the beginning, an approach was adopted to assign a baseline set of conditions 
June 2004 3-3 No. 4: Mechanical Degradation Revision 1 
and then study the effects of other conditions (i.e., sample related, environmental, and inherent 
rock characteristics) relative to that baseline. 
Initially, samples were difficult to obtain, so the baseline conditions were defined as: each test 
specimen was machined as a right-circular-cylinder, with a nominal diameter of 25 mm (1 in) 
and a 2:1 length: diameter ratio, and tested in a water-saturated state at room temperature, 
atmospheric pressure, and a nominal axial strain rate of 10-5/s. The results from these test series 
(Olsson and Jones 1980; Nimick et al. 1985; Price and Jones 1982; Price, Jones, and 
Nimick 1982; Price and Nimick 1982; Price, Nimick, and Zirzow 1982; Price, Spence, and 
Jones 1984) revealed that there is some lateral (i.e., within a unit) and vertical (i.e., unit to unit) 
variability. However, the variabilities in the elastic and strength properties of the tuffs (all 
having similar chemical constituents) are predominantly a function of the tuff porosity 
(Figure 3-3; Price 1983; Price and Bauer 1985). Figure 3-3 shows a compilation of the 
unconfined compressive strength and Youngs modulus as a function of total porosity for test 
samples with a diameter of 50.8 mm. To account for the impact of lithophysae on strength and 
deformability, cores with a diameter of 254 mm (Figure 3-4) were sampled (Price et al. 1985). 
These cores were extracted from the Tptpul from Busted Butte, which is adjacent to Yucca 
Mountain. These results, combined with testing of Tptpul and Tptpll cores with diameters of 
305 and 267 mm, are discussed further in Section 4.2. 
Confined (triaxial) compression testing has been conducted on 25.4 and 50.8 mm samples taken 
from surface boreholes from the various welded and nonwelded tuff units (Figure 3-5). 
Environmental parameters used during the testing, in addition to confining pressure, include 
temperature and saturation (either room dry, heated dry, or vacuum saturated). A detailed 
description of the test data can be found in Subsurface Geotechnical Parameters Report (BSC 
2003a). The results show significant variability, presumably due to the variability in porosity of 
samples within the Tptpmn. Sample temperatures of 150C resulted in a slight decrease in 
strength compared to those samples tested at room temperature. 
Indirect tensile strength testing was conducted on 50.8 mm diameter core samples from the 
Tptpmn and Tptpll under saturated conditions. The results show mean values of 
10.8 MPa 4.02 MPa (14 samples) in the Tptpmn and 8.33 MPa 2.93 MPa (24 samples) in the 
Tptpll (BSC 2003a, Table 8-35). 
The next key study (Price 1986) examined the effect of sample size on the elastic and strength 
properties of the middle nonlithophysal zone of the Topopah Spring tuff (Figure 3-6).9 
Sometime later, a model relating strength, sample size, and functional porosity was developed 
from these data (Price et al. 1993) in combination with the early fits of strength versus functional 
porosity. 
To further explore the effect of sample size for nonlithophysal material, a single large block of 
material from the nonlithophysal section of the Tptpll near its lower boundary with the Tptpln 
was obtained from Busted Butte (outcropping adjacent to Yucca Mountain, Figure 3-7). 
9 As seen in Figure 3-8, this original work on sample size influence on properties of the Tptpmn has been recently 
supplemented by additional size-effect studies on nonlithophysal rock from the base of the Tptpll subunit. 
June 2004 3-4 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2003a, Figures 8-22 and 8-11. 
NOTE: Porosity is composed of matrix and lithophysal porosity. All measurements are from a 50.8 mm diameter 
core that is saturated and at room temperature with a length to diameter ratio of 2:1 and a strain rate of 10-5. 
Small cores from lithophysal zones generally contain only small amounts of lithophysal porosity, and thus 
the above tests are not indicative, in general, of properties of the lower and upper lithophysal units. 
Figure 3-3. Intact Unconfined Compressive Strength (a) and Youngs Modulus (b) for Topopah Spring 
Subunits as a Function of Effective Porosity 
June 2004 3-5 No. 4: Mechanical Degradation Revision 1 
Figure 3-4. Photograph of Busted Butte Sample from the Upper Lithophysal Zone 
June 2004 3-6 No. 4: Mechanical Degradation Source: BSC 2003a, pp. 8-106 and 8-107. 
Figure 3-5. Plot of Axial (1) Versus Confining (3) Stress for 50.8 mm Diameter Specimens of the 
Tptpmn: Room Dry and Room Temperature (top), Saturated at 150C (bottom). 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 3-7 Revision 1 
Source: Price 1986, Figure 11. 
Figure 3-6. Effect of Sample Size on the Unconfined Compressive Strength of Welded Tuff from the 
Middle Nonlithophysal Zone 
NOTE: The block shown was obtained from the nonlithophysal portion of the Tptpll near its lower boundary with the 
Tptpln, Busted Butte. 
Figure 3-7. Development of Rectangular Specimens for Matrix Size Effect and Anisotropy Study 
June 2004 3-8 No. 4: Mechanical Degradation Revision 1 
Several studies (Martin et al. 1993a; Martin et al. 1993b; Martin et al. 1995; Martin et al. 1997a; 
Martin et al. 1997b) have produced indications that the strength properties of the tuffs are 
somewhat time-dependent. A total of nine nonlithophysal rock samples obtained from the 
Tptpmn at Busted Butte were used to conduct static fatigue experiments with saturated rock 
conditions at 150C. Except at very fast rates of deformation (i.e., an axial strain rate of 10-3/s), 
the strengths of the tuffs were found to decrease with decreasing strain rate (from 10-5/s to 
10-9/s), although the average change was relatively minor (about a 10% decrease in strength per 
decade change in strain rate). In addition, constant stress (creep) experiments at fairly high 
stresses (most tests at 100 MPa and higher) resulted in very little strain accumulating after 
several million seconds. The existing test data are described in Section 5.3.3.2.4 (Table 5-8) and 
in detail in Drift Degradation Analysis (BSC 2004a, Appendix S). 
A total of 110 samples with size ranging from 26 to 223 mm diameter were cut and tested to 
examine both size effect and mechanical anisotropy. The mechanical anisotropy studies 
included testing of 62 to 51 mm samples drilled at three mutually perpendicular directions from 
the same block of material. 
The results of the sample size on unconfined compressive strength are shown in Figure 3-8, 
where the unconfined compressive strength is plotted as a function of the sample volume (as a 
log-log plot), and is compared to the data given in Figure 3-6 for the Tptpmn. The vertical offset 
of the two lines is indicative of the slightly different average strength of the Tptpll and Tptpmn 
matrix material, although the size effects for the two are virtually identical. The mechanical 
anisotropy is demonstrated in terms of the average values for the Youngs moduli from each of 
the perpendicular orientations. As seen in Figure 3-9, there is a maximum anisotropy of 
approximately 10.5% in the average matrix moduli, which is considered to be a second order 
effect in comparison to lithophysal and fracturing effects. 
A series of tests were run on 50.8-mm nonlithophysal Tptpll samples from the same outcrop 
boulder to examine the impact of saturation level on unconfined compressive strength. It is 
impossible to accurately control moisture content at specific levels of saturation for a rock 
sample, so a number of tests were performed to fully dry and saturate the samples before 
allowing them to equilibrate at room humidity conditions (Table 3-1). As seen in this table, the 
presence of moisture has a significant effect on unconfined compressive strength, particularly 
whether the samples are under heated-dry or exposed to humid air conditions. Complete drying 
of samples increases the mean strength of the samples tested by approximately 20%. 
June 2004 3-9 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure E-20. 
NOTE: Results from the 2003 testing of Tptpln/Tptpll samples (DTN: SN0306L0207502.008) are compared to 
previous testing of samples from the Tptpmn (Price 1986). The 2003 data, although from the Tptpll, are 
composed of matrix material and contain no observed lithophysae. 
Figure 3-8. Results of Size Effect Study Showing Variation in Sample Unconfined Compressive 
Strength as a Function of Sample Volume 
Source: BSC 2004a, Figure E-21. 
NOTE: Averages of values from 62 samples shows maximum of 10.5% anisotropy in the average modulus. 
Figure 3-9. Anisotropy in Youngs Modulus of Nonlithophysal Tptpll Matrix for Three Mutually 
Perpendicular Coring Directions 
June 2004 3-10 No. 4: Mechanical Degradation Revision 1 
Table 3-1. Impact of Moisture Conditions on Unconfined Compressive Strength of Nonlithophysal Tptpll 
Samples 
Test 
Condition 
1 
2 
3 
4 
Mean Strength 
(MPa) 
213 
176 
158 
149 
Moisture Condition 
Samples dried by slow heating to 200C, tested at 200C 
Samples dried by slow heating to 200C, then slowly cooled in dry environment, 
exposed to room humidity for about 30 minutes, and tested at room temperature 
Samples allowed to equilibrate with room humidity, tested at room temperature 
Samples water saturated, tested at room temperature 
Source: BSC 2004a, Table E-12. 
NOTE: Strengths are mean values from testing of 51-mm diameter samples at each moisture condition. 
This strength decrease in the presence of moisture is consistent with other testing of silicic rocks 
and is a typical stress-corrosion mechanism involving chemical alterations due to moisture in 
flaws within the samples. The compression test data reported here is, unless otherwise noted, all 
at room humidity conditions. 
Following a conservative design approach, performance calculations performed for ground 
support or postclosure effects assume average strength conditions from room temperature testing, 
with data ranges to cover fully saturated conditions. 
Large Core Laboratory Testing of Lithophysal RocksIn an attempt to obtain more 
representative samples of lithophysal rocks, a series of large core (305-mm diameter) samplings 
of the Tptpul and Tptpll were taken in the ESF and ECRB Cross-Drift. The objective of the 
drilling was to obtain samples that had a number (approximately 3 to 5) of lithophysae across a 
sample diameter. Drilling of such large core in this material is quite difficult due to the tendency 
of the core to break in shear when a large or poorly placed lithophysae is encountered. 
Figure 3-10 shows a series of these samples as prepared for laboratory compression testing. 
Approximately 20 samples were tested at varying environmental conditions, including air dry, 
saturated under vacuum, and at 200C. Plots of the unconfined compressive strength and 
Youngs modulus as functions of the approximate levels of lithophysal porosity10 are given in 
Figure 3-11. In these results, the unconfined compressive strength and Youngs Modulus are 
relatively insensitive to lithophysal porosity above approximately 20%, with rapid increase in 
both for lithophysal porosities below about 15%. This trend fits the general exponential 
relationship between these parameters and porosity as seen in Figure 3-3, and can generally be 
viewed as an extension of this response at higher porosity levels. As was shown previously via 
field measurement in the ECRB Cross-Drift (Figure 2-10), lithophysal porosity through the 
Tptpll averages approximately 15%, with a range from about 10% to 25%. The laboratory-scale 
unconfined compressive strength of the Tptpll can vary from as high as about 25 to 30 MPa to as 
low as about 10 MPa, while the Youngs modulus can vary from 20 to 5 GPa. 
10 Lithophysal porosity is approximate, as it was estimated from core surface measurements. 
June 2004 3-11 No. 4: Mechanical Degradation Revision 1 
Figure 3-10. Photographs of Large Lithophysal Core Samples (290 mm/11.5 diameter) from the Tptpll 
and Tptpul (a) and a Sample in Unconfined Compression (b) 
June 2004 3-12 No. 4: Mechanical Degradation Revision 1 
Source: DTNs: SN0208L0207502.001; SN0211L0207502.002; SN0305L0207502.006 (for laboratory data). 
Lithophysal Rocks 
A convenient way of presenting these data for design purposes is in terms of the relationship 
between the unconfined compressive strength and Youngs modulus (Figure 3-12). Some of the 
primary mechanical properties that are input data to numerical models are the unconfined 
compressive strength and Youngs modulus. Figure 3-12 illustrates the interrelationship of these 
parameters, the total range of their variation, and that they vary in a generally linear fashion, 
Figure 3-11. Unconfined Compressive Strength (a) and Youngs Modulus (b) as Functions of Estimated 
Lithophysal Porosity from Large Core (290 mm diameter) Compression Tests from 
June 2004 3-13 No. 4: Mechanical Degradation Revision 1 
irrespective of the lithophysal porosity values. As will be discussed later, this entire range of 
data is subdivided into a series of categories for design parameter studies. Discontinuum 
numerical models, calibrated to this data, are used to examine the impact of variability of 
lithophysal porosity, shape, size and distribution on the variability of these parameters for design 
purposes. 
Source: DTNs: SN0208L0207502.001; SN0211L0207502.002; SN0305L0207502.006 (for laboratory data). 
NOTE: Linear fits made to 11.5-in. and 10.5-in. core tests. 
Figure 3-12. Unconfined Compressive Strength as a Function of Youngs Modulus from Large Core 
(290 mm and 267 mm diameter) Compression Tests from Lithophysal Rocks 
June 2004 
3.2.2 Summary of Laboratory Testing of Lithophysal and Nonlithophysal Cores 
In summary, a large amount of data has been collected to date on tuffaceous samples from Yucca 
Mountain at a baseline set of conditions. These data have shown that the variabilities in elastic 
and strength properties are not random functions of lateral or vertical position, but primarily a 
function of porosity and its spatial distribution. 
Though there are obvious trends in the Youngs modulus and strength data when plotted against 
porosity, the data have a significant scatter, based on sample porosity variability. The secondary 
effect that creates the scatter is the distribution of the porosity within the sample. Other 
investigations have examined the effects of many other conditions (i.e., sample related, 
environmental and inherent rock characteristics); for example, sample size, saturation, pressure, 
temperature, deformation rate, attenuation, anisotropy have all been studied. 
3-14 No. 4: Mechanical Degradation Revision 1 
In Section 4, the development of mechanical material models and numerical modeling 
techniques are described for both lithophysal and nonlithophysal units. In lithophysal rocks, the 
trends observed in the relationship of mechanical properties to porosity are used as a basis for 
development of the models. In situ measurements of the variability of lithophysal porosity are 
used as the basis for establishing a linkage between the measured mechanical properties and the 
estimated field properties ranges. For nonlithophysal rocks, discontinuum modeling approaches 
are used in which intact blocks and fractures are modeled explicitly. The intact core testing data 
described in this section is used to assign the mechanical properties of intact blocks, whereas 
fracture shear strength properties (described in Section 3.2.4) define the fracture response. In 
conclusion, the intact and large core rock mechanical property information collected over the last 
two decades has provided adequate characterization and property values for design and analysis 
of the behavior of Yucca Mountain tuffs. 
3.2.3 In Situ Mechanical Testing of Nonlithophysal and Lithophysal Rocks 
In situ testing provides data on the mechanical characteristics of the rock mass at a scale 
commensurate with the excavation dimension. Field compression testing has been performed in 
both the nonlithophysal and lithophysal rocks. For nonlithophysal rocks (Tptpmn), field testing 
is restricted to measuring the deformation modulus, as the rock mass strength is too large to 
conduct strength measurements. However, in the lithophysal rocks, the in situ rock mass 
strength is sufficiently low that both the deformation modulus and unconfined compressive 
strength can be measured. 
The Plate Loading Test, conducted as part of the larger DST, was used to apply load to the drift 
wall adjacent to the DST in the Tptpmn unit. The results of the testing show an estimated rock 
mass modulus from 11.4 GPa to 29.5 GPa for ambient and thermally perturbed fractured tuffs, 
respectively (George et al. 1999). 
A series of three pressurized slot tests (PSTs) were conducted to perform deformation modulus 
and strength measurements in lithophysal rock units (Table 3-2). PST#1 was conducted in the 
poorest quality Tptpll material characterized by large lithophysae and heavily fractured matrix at 
the transition between the Tptpmn and the Tptpll. PST#2 was conducted in good quality Tptpul 
in the ESF and PST#3 was conducted in what was considered to be average Tptpll repository 
conditions in the ECRB Cross-Drift. All three tests included compressions at ambient 
temperature and one test included compressions at elevated temperature. 
3-15 June 2004 No. 4: Mechanical Degradation Revision 1 
Table 3-2. Summary of In Situ Slot Compression Tests in Lithophysal Rock Units 
Location 
ESF 57+77 
Test 
PST#1 
Thermal- 
Mechanical Unit 
Tptpll 
PST#2 Tptpul 
Tptpul 
ESF 63+83 
ESF 63+83 
Location 
Wall 
Wall 
Wall 
Floor 
Temperature 
ambient 
ambient 
temperature90C 
ambient ECRB 21+25 PST#3 Tptpll 
Source: Sobolik 2002a, Table 1; Sobolik 2002b, Table 1; Schuhen and Sobolik 2003, Table 1. 
NOTE: Metric stationing is used throughout the ESF, so that Station 57+77 is located 5,777 m 
from the start of the tunnel. 
The flatjack slot test first involves placing steel flatjack bladders into parallel, thin, sawcut 
excavations in the sidewall or floor of a tunnel (Figure 3-13). The parallel sawcuts isolate an 
approximately 1-m3 rock specimen between them, which is subsequently compressed by 
pressurizing the flatjacks within the slots. Typically, an instrumentation borehole (305 mm in 
this case) is drilled midway between the slots to allow observation of the interior of the rock 
sample and to monitor deformations during the flatjack pressurization. The flatjack slots are 
excavated larger than the flatjacks to minimize end effects due to block attachment to the 
surrounding solid rock. The flatjacks pressure is raised in a series of pressure cycles to monitor 
hysteresis effects and time-dependent strain at increasing levels of applied stress. 
(b) Slot Test 2 Showing Central Instrumentation Hole and Parallel Slots 
Figure 3-13. Photographs of (a) Preparation of Slot Test 3 in the Floor of the Enhanced Characterization 
of the Repository Block, Tptpll Unit, Lowering of Flatjack into Place in Sawcut Slot and 
Figure 3-14 shows flatjack pressure versus instrumentation hole diametral deformation for the 
tests. The loading results show a typical elastic-plastic response in which a linear loading slope 
is followed by yield and plastic deformation. Yielding of the rock was typically in shear, 
emanating from the central borehole, and resulting in rockfall in the form of small rock particles 
in the central borehole during yield. The results, summarized in Table 3-3, show that the rock 
mass has deformation modulus and strength that lie at the lower end of the design range given in 
June 2004 3-16 No. 4: Mechanical Degradation Figure 3-12, but consistent with the same general relationship of strength to modulus as observed 
in the large core laboratory tests. However, the results illustrate the difficulties with performing 
large in situ slot tests. PST#1 was located in the poorest quality lithophysal rock immediately 
adjacent to the contact of the Tptpmn. The low value of modulus obtained indicates that the skin 
of rock surrounding the tunnels at this particular sidewall locations, is likely in a preyielded state 
due to mining-induced stress and excavation effects induced by the TBM. PST#2 resulted in 
shear along a preexisting fracture outside the block (without failing the block itself), and PST#3 
resulted in spalling of the tunnel floor prior to actually failing the rock block between the 
flatjacks. 
Source: Sobolik 2002a, Figure 7; Sobolik 2002b, Figures 7 and 9; Schuhen and Sobolik 2003, Figure 7. 
Figure 3-14. Composite of Flatjack Pressure versus Central Hole Diametral Strain for the Three 
Pressurized Slot Tests 
Table 3-3. Summary of Mechanical Properties Results from the Pressurized Slot Tests 
Test 
PST#1 
PST#2 
PST#2 
PST#3 
Source: Sobolik 2002a, Table 1; Sobolik 2002b, Table 1; Schuhen and Sobolik 2003, Table 1. 
NOTE: PST#1 in poorest quality Tptpll, PST#2 in good quality Tptpul, and PST#3 in typical repository Tptpll. 
Strength here is the peak flatjack pressure reached during the test. PST#2 failed on the fracture 
Strength (MPa) 
6a 
NA 
11a 
7a 
Location 
ESF 
ESF 
ESF 
ECRB Cross-Drift 
outside block, and PST#3 spalled the tunnel floor above flatjacks. 
a Results do not account for presence of central hole in failure load. 
Tuff Unit 
Tptpll 
Tptpul 
Tptpul 
Tptpll 
Condition 
Ambient 
Ambient 
Heated, >80C 
Ambient 
E (GPa) 
0.5 0.3 
3.0 0.5 
1.5 0.5 
1.0 0.3 
No. 4: Mechanical Degradation 3-17 
Revision 1 
June 2004 Revision 1 
3.2.4 Mechanical Properties of Fractures 
3.2.4.1 Direct Shear Test Results 
Five direct shear tests were performed on core samples with a diameter of 305 mm and with 
internal fracture surfaces from the Tptpmn unit. Photographs of the direct shear samples are 
shown in Figure 3-15. Two tests were performed on subvertical cooling joints, and three tests 
were conducted on more or less horizontal vapor-phase partings (VPPs). The two fracture types 
have a physically distinct appearance, and testing of these fractures resulted in equally distinct 
fracture behaviors (Table 3-4). 
Figure 3-15. Photographs of Direct Shear Samples from Rough Vapor Phase Partings (a) and Smooth 
Cooling Joints (b) 
Cohesion and friction angle parameters were determined from repeated tests on the same fracture 
surface (Table 3-4). The advantage of this is that the same fracture is being tested in the same 
equipment setup. The disadvantage is that because of repeated tests on the same fracture surface, 
degradation of the asperities changes the fracture behavior on subsequent loading tests. 
June 2004 3-18 No. 4: Mechanical Degradation Table 3-4. Direct Shear Test Summary of Tptpmn Fractures 
Joint Type Test ID 
65A-643 Cooling 
65A-657 Cooling 
65A-642 VPP 
65A-646 VPP 
65A-647 VPP 
Source: DTN: GS030283114222.001. 
Length 
(mm) 
238.76 
142.24 
241.30 
226.06 
246.38 
NOTE: JRC = the measured value of joint roughness coefficient; CC = the correlation coefficient of a 
linear fit through the data. 
Cohesion 
Area (m2) JRC (MPa) 
2.00 0.04 
1.00 0.02 
15.00 0.06 
16.00 0.03 
10.00 0.05 
-0.01 
0.08 
0.72 
0.66 
0.84 
Table 3-5 illustrates that cooling fractures have lower cohesion, lower peak friction angle, and 
much lower peak dilation angle than the VPP fractures. 
Table 3-5. Summary Statistics of Direct Shear Tests of Fracture Peak Strength 
Count Mean 
2 
3 
2 
3 
2 
3 
Unit 
Cooling Tptpmn 
Tptpmn 
Cooling Tptpmn 
Tptpmn 
Cooling Tptpmn 
Tptpmn 
Joint 
VPP 
VPP 
VPP 
Direct Shear Rock Joint Dilation Angle at Peak Stress 
1.6 
13.7 
2.33 
0.15 
3.61 
2.12 
2.55 
1.22 
1.6 
14.0 
Source: DTN: GS030283114222.001. 
Error 
Deviation / 
Mean 
Direct Shear Rock Peak Cohesion (MPa) 
Standard Standard 
Deviation 
1.97 
0.12 
0.06 
0.09 
0.04 
0.05 
0.32 
0.74 
Direct Shear Rock Peak Friction Angle () 
0.01 
0.04 
0.4 
1.9 
0.3 
1.1 
33.4 
44.0 
Median Minimum Maximum 
0.03 
0.72 
33.0 
44.5 
0.08 
0.84 
-0.01 
0.66 
33.7 
45.7 
33.1 
41.9 
4.1 
16.3 
-1.0 
12.1 
The final normal stress and shear stress measured in the last shearing for each fracture was used 
to determine the degraded friction angle values assuming the cohesion was zero. These degraded 
shear results are shown in Table 3-6. The peak and degraded friction angles are roughly the 
same for the smooth subvertical cooling fracture, whereas the degraded friction angle for VPP 
fractures was slightly higher than the peak value. 
Table 3-6. Summary Statistics of Direct Shear Fracture Degraded Strength 
Unit 
Tptpmn 
Tptpmn 
Joint 
Cooling 
VPP 
Direct Shear Rock Degraded Friction Angle () 
Count Mean 
2 
3 
Source: DTN: GS030283114222.001. 
33.4 
46.9 
No. 4: Mechanical Degradation 
Error 
Standard Standard Deviation / 
Deviation Mean 
2.7 
3.0 
1.9 
1.7 
0.08 
0.06 
3-19 
Friction 
Angle () 
33.7 
33.1 
45.7 
41.9 
44.5 
Median Minimum Maximum 
33.4 
46.5 
35.3 
50.0 
31.5 
44.1 
Revision 1 
CC 
1.00 
1.00 
0.99 
1.00 
0.99 
June 2004 3.2.4.2 Rotary Shear Tests 
A series of 22 rotary shear tests have been conducted on natural rock fractures from core samples 
in the repository host units. The samples ranged in size to 76 mm and tests were conducted at 
room dry, room temperature and at 175C. Rotary tests involve application of constant normal 
stress and torque to undercored samples (undercoring creates a pipe-like sample with an inner 
and outer radius). The shear stress induced on the surfaces of the joint result in slip and dilation 
of the surfaces can be determined with continued displacement. Table 3-7 provides summary 
statistics of the results of this testing; further discussion can be found in Subsurface Geotechnical 
Table 3-7. Summary Statistics of Fracture Strength Using Rotary Shear 
Parameters Report (BSC 2003a). 
Rock Unit 
Tptpul/ll 
Peak CC* 
0.92 
Count 
6 
Peak 
Friction 
Angle 
40 
Peak 
Cohesion 
(MPa) 
1.48 
Temperature 
( C) 
Room 
0.85 0.79 8 Tptpmn/ln Room 42 
0.96 1.38 8 Tptpmn/ln 175 44 
NOTE: CC* refers to the correlation coefficient of a linear fit through the data. 
Source: DTNs: SNL02112293001.001; SNL02112293001.003; SNL02112293001.005; SNL02112293001.006; 
SNL02112293001.007. 
(Eq. 3-1) 
June 2004 
3.2.4.3 Empirical Estimate of Peak Dilation Angle 
Barton (in Duan 2003, p. 40) used the following equation to empirically estimate peak dilation 
angles11,  peak, for Yucca Mountain joint sets: 
peak = 1 JRC log. . . 
JCS .. . 
2 
11 
.  n . 
where JRC is joint roughness coefficient, JCS is joint wall compressive strength, and n 
is the 
effective normal stress. A range of laboratory scale JRC values for joint sets is estimated 
(Table 3-4) by making JRC measurements from roughness traces of the rotary shear and direct 
shear laboratory fractures, making field JRC measurements of fractures in the ESF, and 
correlating JRC values to U.S. Bureau of Reclamation field roughness statistics (R1 to R6). 
Adopting a common value of joint wall normal strength (JCS) of 100 MPa, a normal stress 
(n) range of 4 to 8 MPa, and the joint roughness coefficient (JRC) range shown in Table 3-8, the 
range of peak dilation angles is derived. The empirically derived field dilation angle estimates 
are consistent with those determined from the laboratory testing (Table 3-5). 
Dilation angle is defined as the ratio of the displacements perpendicular and parallel to the fracture surface when 
subjected to shear. This value provides an indication of the roughness of the joint surface. 
3-20 No. 4: Mechanical Degradation 
Revision 1 
Ultimate 
Cohesion Ultimate 
CC* 
0.76 
Ultimate 
Friction 
Angle 
32 
(MPa) 
1.51 
0.90 0.60 39 
0.91 0.24 44 Revision 1 
Nonlithophysal 
Rock 
Table 3-8. Estimate of Peak Dilation Angles for Topopah Spring Formation Fracture Sets 
Cooling Joint Sets 1 and 4 
Cooling Joint Set 2 
VPP Joint Set 3 
Lithophysal Fractures 
Joint Roughness 
Coefficient Range 
2 to 4 
4 to 8 
12 to 16 
12 to 20 Lithophysal 
Rock 
NOTE: Peak dilation angles calculated from Equation 3-1 for JCS = 100 MPa and n = 4 to 8 MPa. 
3.2.5.1 Thermal Conductivity 
Thermal conductivity is a proportionality constant that relates the heat transfer (conduction) rate 
per unit area in a material to the normal temperature gradient. Thermal conductivity of rock is 
used in predicting temperature changes in the rock mass after waste emplacement. 
Peak Dilation Angle Range 
1.4 to 2.2 
2.8 to 4.4 
8.4 to 8.8 
8.4 to 11.0 
June 2004 
3.2.5 Thermal Properties of the Repository Host Rocks 
Thermal properties (i.e., thermal conductivity, heat capacity or specific heat, and coefficient of 
thermal expansion) of lithostratigraphic rock units at the repository host horizon are important 
parameters used in the design and performance assessment because they are used in calculating 
the transient rock mass temperature and thermally induced stress. Their values are estimated 
primarily based on laboratory and field measurements. These measurements are essential in 
providing not only the site-specific values of thermal properties but also the information of their 
spatial variability and dependencies on temperature, porosity and fracture, and moisture content. 
Additionally, these testing efforts have also assisted in development of theoretical models that 
describe spatial variation of thermal properties and correlation between rock mass thermal 
properties, intact rock thermal properties, and other rock properties such as porosity. 
Numerous laboratory tests using small specimens containing few voids and/or fractures show 
that intact rock thermal characteristics of lithophysal and nonlithophysal of rocks are similar 
(CRWMS M&O 1997b, Tables 5-11, 5-13, 5-15, and 5-16), indicating that similar methods may 
be used for acquiring intact rock thermal properties. However, rock mass characteristics of these 
rocks are quite different due to their different dominant features. These impacting factors are 
reflected in the difference in their rock mass thermal properties, suggesting, for example, the use 
of different methods of acquiring rock mass thermal properties for the lithophysal rocks versus 
those for the nonlithophysal rocks. Due to the associated relatively high cost and logistics 
involved, only a limited number of field thermal tests have been conducted (BSC 2002a, 
Sections 6.2.3.5, 6.3.1.4, and 6.3.3.6.5). Thus, estimation of rock mass thermal properties and 
their spatial variations based solely on the available field testing data may not be sufficient. This 
points to the desirability of using theoretical models to aid in describing the correlation between 
intact rock and rock mass thermal properties and the spatial variations. 
Factors that may affect thermal properties include temperature, porosity, fracture, moisture 
content, specimen size or scale, mineral content, and loading condition. The effects of these 
factors on thermal properties have been a primary focus of investigation in the characterization 
of Yucca Mountain host rocks. 
3-21 No. 4: Mechanical Degradation Revision 1 
Due to the heterogeneity and discontinuities in the rock units that will host the repository, 
thermal conductivity of the rock of interest is both scale- and direction-dependent. When 
increasing the size of testing specimens, the degree of heterogeneity and the impact of 
discontinuities increase, thus affecting the thermal conductivity of the rock specimen. It is 
important to understand the effect of scale on the rock thermal conductivity and the difference 
between the intact rock and the rock mass, so a correct value of thermal conductivity can be used 
in the design. 
Intact Rock Thermal ConductivityThermal conductivity of intact rock was estimated based 
on laboratory thermal conductivity measurements using small specimens. These specimens had 
nominal dimensions of 50.8 mm in diameter and 12.7 mm in length (Brodsky et al. 1997, Section 
2.1, Table 2-1), and contained few voids or fractures. The effects of these discontinuities on the 
measurements were considered minimal. A large number of laboratory thermal conductivity 
measurements have been conducted since the late 1990s. The results are summarized in Table 39 
for the four rock units at the repository host horizon. 
Table 3-9. Intact Rock Thermal Conductivities for Repository Units 
Below 100C (W/mK) Above 100C (W/mK) 
Dry Dry Air Dry Saturated Rock 
Units 
Tptpul 
Mean 
1.97 
Standard 
Deviation 
0.11 
Standard 
Deviation 
0.21 
Standard 
Deviation 
0.22 
0.11 0.12 0.45 2.33 Tptpmn 
0.04 0.08 0.13 2.13 Tptpll 
Mean 
1.20 
1.68 
1.65 
N/A 
Mean 
1.07 
1.51 
1.45 
N/A 
Mean 
1.06 
1.60 
1.54 
N/A 
Standard 
Deviation 
0.12 
0.49 
0.03 
N/A N/A N/A N/A Tptpln N/A 
June 2004 
Source: CRWMS M&O 1997b, Tables 5-11 and 5-13; DTN: SNL01A05059301.005. 
Rock Mass Thermal ConductivityThermal conductivity of the rock mass is the effective value 
of thermal conductivity that relates the heat conduction rate to the normal temperature gradient in 
a rock mass. It accounts for the effects of voids, fractures, and any heterogeneity or 
discontinuity on thermal conductivity. Correlations between intact rock or rock matrix thermal 
conductivity, porosity (both matrix and lithophysae), and rock mass thermal conductivity are 
developed in Subsurface Geotechnical Parameters Report (BSC 2003a, Section 8.3.3.2). These 
efforts included both conducting experimental tests and developing theoretical approaches. 
A limited number of experimental tests have been conducted to estimate rock mass thermal 
conductivity. They included laboratory measurements using large specimens and field 
measurements in the DST and the ECRB Cross-Drift. The DST measurements were of the 
Tptpmn unit, while the ECRB Cross-Drift measurements were of the Tptpll unit. Table 3-10 
provides a range of rock mass thermal conductivity for the Tptpmn and Tptpll units obtained 
from the field measurements. Compared to those listed in Table 3-9, it is seen that the in situ 
values are within the ranges observed in the laboratory measurements on small specimens. 
3-22 No. 4: Mechanical Degradation Revision 1 
Table 3-10. Rock Mass Thermal Conductivities from Field Measurements 
Tptpmna 
Range of Thermal Conductivity (W/mK) 
1.69 to 1.95 
1.73 to 2.18 Tptpllb 
Rock Unit 
a
b 
Source: DTNs: LL980411104244.061; LL980902104244.070; 
UN0106SPA013GD.004; UN0201SPA013GD.007. 
DTNs: SN0206F3504502.012; SN0206F3504502.013; 
SN0208F3504502.019. 
Alternative analytical approaches have also been developed to estimate rock mass thermal 
conductivity (BSC 2002b). The rock mass thermal conductivity, estimated from analytical 
correlations for the various repository host horizons, is summarized in Table 3-11. The 
analytical approaches, which account for matrix void volume, compares well to field 
measurements in Table 3-10. Compared to those listed in Table 3-9 for the intact rock, it is 
obvious that porosity and moisture content have significant effect on the rock mass thermal 
conductivity. For conservatism, ranges of thermal properties that encompass the rock mass 
thermal conductivity are used in design and performance assessment. 
Table 3-11. Rock Mass Thermal Conductivities for Repository Units 
Saturated (W/mK) Dry (W/mK) 
Rock Units 
Tptpul 
Standard 
Deviation 
0.24 
0.27 Tptpmn 
0.25 Tptpll 
Mean 
1.18 
1.42 
1.28 
1.49 
Mean 
1.77 
2.07 
1.89 
2.13 
Standard 
Deviation 
0.25 
0.25 
0.25 
0.27 0.28 Tptpln 
3.2.5.2 Heat Capacity 
Heat capacity of a substance is defined as the amount of energy required to raise the temperature 
of a unit mass of the substance by one-degree (Nimick and Connolly 1991, p. 5). It is an 
important parameter used in thermal analysis to evaluate temperature changes in rock after waste 
emplacement. For solid materials, heat capacity is strongly dependent on temperature. For the 
temperature range of interest in the design and performance assessment, heat capacity for the 
repository rock units is estimated for three temperature ranges, 25C to 94C, 95C to 114C, 
and 115C to 325C, corresponding to the preboiling, transboiling, and postboiling regimes, 
respectively. 
Only a limited number of laboratory and field measurements have been made to estimate rock 
heat capacitance (product of heat capacity and density). These measurements covered only a few 
rock units. Instead, the heat capacity values used in the design and performance assessment are 
Source: DTN: SN0208T0503102.007. 
The thermal conductivity ranges provided in Table 3-11 are used as the basis for parametric 
analysis of rock mass temperature distributions in the postclosure period (see Section 5.3.1). 
3-23 June 2004 No. 4: Mechanical Degradation largely based on the calculated values obtained from analytical methods. The estimates have 
been compared with laboratory and field measurements and correlate sufficiently well as to 
validate the estimate model and the resulting values. These methods are presented in Heat 
Table 3-12. Rock Grain Heat Capacities for Repository Units 
Average Average Standard 
Deviation 
Capacity and Thermal Expansion Coefficients Analysis Report (BSC 2003f). 
Rock Grain Heat CapacityThe calculated average values of rock grain heat capacity for the 
four repository rock units are presented in Table 3-12. These values were estimated based on 
available data on mineral abundance and mineral heat capacity. 
Rock 
Units 
Tptpul 
Tptpmn 
Tptpll 
Tptpln 
T = 25C to 94C T = 95C to 114C 
Standard 
Deviation 
90 
110 
100 
70 
90 
110 
100 
70 
870 
870 
870 
870 
780 
780 
780 
780 
Source: DTN: SN0307T0510902.003. 
NOTE: All measurements are in (J/kgK). 
Rock Mass Heat CapacityRock mass heat capacity is the effective value of heat capacity that 
accounts for the effect of air-filled voids and of water that exists in the voids. 
Efforts to measure the rock mass heat capacity were made. Rock mass volumetric heat capacity 
or heat capacitance of the Tptpll unit was estimated from the thermal measurements in the ECRB 
Cross-Drift. The estimated rock mass volumetric heat capacity values range from 1.96  106 to 
2.30  106 J/m3K (DTNs: SN0206F3504502.012; SN0206F3504502.013; 
SN0208F3504502.019). Given a bulk density value of 2,360 kg/m3 for the Tptpll unit, the rock 
mass heat capacity is estimated to range from 831 to 975 J/kgK. 
The calculated values of rock mass heat capacity for the four repository rock units using the 
analytical methods are summarized in Table 3-13. These values were estimated for the 
preboiling, transboiling, and postboiling regimes, based on the available data on rock matrix 
porosity and saturation, lithophysal porosity, rock grain heat capacity, and density. 
Table 3-13. Rock Mass Heat Capacities for Repository Units 
Average Average Standard 
Deviation 
Rock 
Units 
Tptpul 
Tptpmn 
Tptpll 
Tptpln 
NOTE: All measurements in (J/kgK). 
T = 25C to 94C T = 95C to 114C 
Standard 
Deviation 
1000 
900 
1000 
800 
300 
300 
300 
300 
940 
910 
930 
900 
Source: DTN: SN0307T0510902.003. 
3600 
3000 
3300 
2800 
3-24 No. 4: Mechanical Degradation 
Revision 1 
Average 
T = 114C to 325C 
Standard 
Deviation 
110 
130 
120 
90 
990 
990 
990 
990 
Average 
T = 114C to 325C 
Standard 
Deviation 
300 
300 
300 
300 
990 
990 
990 
990 
June 2004 Revision 1 
It is appropriate to use the rock mass heat capacity in the design and performance assessment if 
the phase change over the transboiling regime cannot be accounted for in the analysis. 
Otherwise, the rock grain heat capacity should be used because the modeling accounts for the 
heat capacity effects in boiling of pore water. The estimates have been compared with laboratory 
and field measurements and sufficient correlation is found to validate the estimate model and the 
resulting values. 
3.2.5.3 Coefficient of Thermal Expansion 
Thermal expansion is a mechanical response in the form of strain because of the change of 
temperature. The coefficient of thermal expansion (CTE) of rock is strongly dependent on 
temperature. It is an important parameter in thermal-mechanical analysis to predict thermally 
induced rock displacements and stresses and to evaluate stability of repository openings and 
performance of installed ground support during heating. 
Intact Rock CTEIntact rock CTE was estimated based on laboratory thermal expansion 
measurements using small specimens. A large number of thermal expansion measurements have 
been made on specimens taken from the rock units at the repository host horizon. Most of the 
measurements were conducted on dry or saturated specimens over a temperature range of 25C 
to over 300C. Table 3-14 summarizes the measured intact rock CTE for the four repository 
rock units. 
Rock Mass CTERock mass CTE is the effective thermal expansion that rock mass experiences 
when subjected to a change in temperature. It accounts for the effects of voids, fractures, 
moisture content, and any heterogeneity or discontinuity that affect the thermal expansion. 
Estimation of rock mass CTE is based on field or large core (diameter of 305 mm) thermal 
expansion measurements. Two major field tests, which involved the measurements of rock mass 
CTE, are the Single Heater Test and the DST. Both tests are located in the Tptpmn unit. The 
results from these measurements are summarized in Table 3-15. 
There are no field thermal expansion measurements available in the Tptpll unit. The best data 
available on rock mass CTE for this rock unit are those from laboratory thermal expansion 
measurements on specimens with a nominal diameter of 305 mm (12 in.). The results from these 
laboratory measurements are also presented in Table 3-15. 
After comparing the rock mass CTEs presented in Table 3-15 to those listed in Table 3-14 for the 
intact rock, it is apparent that the former are lower than the latter. The difference decreases as 
temperature increases, which indicates that the effect of fractures or voids on CTEs diminishes as 
more fractures or voids are closed by rock deformation as a result of temperature increase. From 
the perspective of ground support design, use of the intact rock CTE is conservative. 
3-25 June 2004 No. 4: Mechanical Degradation Mean 
Mean 
Dry 
Std. Dev. 
Mean 
Saturated 
Std. Dev. 
Tptpmn 
Mean 
Dry 
Std. Dev. 
Mean 
June 2004 
Temperature 
Range 
Mean 
Saturated 
Std. Dev. 
Mean 
Dry 
Std. Dev. 
Mean 
Saturated 
Std. Dev. 
Mean 
Dry 
Std. Dev. 
Mean 
Saturated 
Std. Dev. 
Mean 
Dry 
Std. Dev. 
Mean 
Saturated 
Std. Dev. 
Mean 
Dry 
Std. Dev. 
Temperature 
Range 
Rock 
Units 
Tptpul 
Tptpmn 
Tptpll 
Tptpln 
Rock 
Units 
Saturated 
Std. Dev. 
Tptpul 
Saturated 
Std. Dev. 
Tptpll 
Mean 
Dry 
Std. Dev. 
Mean 
Saturated 
Std. Dev. 
Tptpln 
Mean 
Dry 
Std. Dev. 
Source: CRWMS M&O 1997b, Tables 5-15 and 5-16; DTN: SNL01B05059301.006. 
3-26 No. 4: Mechanical Degradation 
Table 3-14. Intact Rock Coefficients of Thermal Expansion for Repository Units 
200 
225C 
25.60 
7.08 
29.34 
10.73 
15.53 
1.02 
14.57 
2.04 
15.42 
2.22 
15.14 
3.26 
N/A 
N/A 
12.78 
1.53 
100 
125C 
CTE on Heat-Up (10-6/C) 
CTE on Cool-Down (10-6/C) 
150 
175C 
12.95 
1.76 
13.51 
2.57 
11.74 
0.47 
10.95 
0.52 
11.73 
1.76 
10.75 
1.01 
N/A 
N/A 
11.56 
2.75 
150 
175C 
100 
125C 
10.22 
0.69 
9.52 
0.52 
8.73 
2.04 
9.50 
0.27 
9.37 
2.78 
9.12 
0.57 
N/A 
N/A 
9.58 
1.07 
200 
225C 
75 
100C 
7.91 
0.65 
8.89 
0.39 
7.93 
0.94 
8.95 
0.24 
7.03 
1.31 
8.77 
0.54 
N/A 
N/A 
8.79 
0.47 
225 
250C 
50 
75C 
7.00 
0.33 
8.43 
0.36 
7.78 
1.90 
8.45 
0.30 
7.20 
1.09 
8.15 
0.47 
N/A 
N/A 
8.24 
0.57 
250 
275C 
25 
50C 
7.59 
0.01 
7.41 
0.42 
7.20 
0.84 
6.89 
1.45 
7.09 
0.45 
6.41 
0.75 
N/A 
N/A 
6.55 
1.29 
275 
300C 
33.46 34.81 28.46 19.87 
2.90 3.38 1.92 2.44 
30.06 34.20 29.24 21.56 
8.76 15.44 9.61 3.32 
25.84 36.20 38.28 27.79 
4.41 5.05 2.14 1.45 
22.55 28.39 30.08 24.82 
4.27 6.30 5.33 2.25 
17.91 19.05 19.71 17.30 
3.92 4.90 5.31 3.93 
21.69 22.11 20.16 17.15 
8.17 8.25 4.78 2.71 
N/A N/A N/A N/A 
N/A N/A N/A N/A 
15.07 15.38 15.52 15.02 
1.63 2.28 2.89 2.70 
125 
150C 
10.76 
0.32 
10.86 
1.34 
10.11 
0.87 
10.12 
0.36 
9.87 
0.69 
9.87 
0.68 
N/A 
N/A 
10.65 
2.17 
175 
200C 
175 
200C 
16.73 
3.19 
19.38 
6.89 
12.96 
0.70 
12.09 
1.01 
13.20 
1.85 
12.55 
1.80 
N/A 
N/A 
11.90 
2.35 
125 
150C 
17.14 23.61 35.28 
2.61 5.79 5.21 
16.63 22.27 30.33 
4.29 7.57 11.03 
12.75 14.51 17.93 
0.84 1.32 3.02 
11.88 13.72 17.20 
2.78 3.42 5.10 
12.26 13.66 16.75 
1.65 1.38 3.16 
11.47 15.80 22.06 
3.63 5.54 14.24 
N/A N/A N/A 
N/A N/A N/A 
11.98 13.03 14.08 
2.56 2.61 2.06 
225 
250C 
32.83 
3.35 
32.35 
8.56 
20.60 
2.04 
19.45 
3.47 
17.80 
3.29 
25.19 
27.61 
N/A 
N/A 
13.87 
1.11 
75100C 
11.91 13.93 
0.41 1.30 
10.82 13.40 
0.91 2.41 
10.65 11.48 
0.47 0.63 
9.93 10.73 
1.07 1.86 
9.92 11.56 
0.54 2.77 
8.88 10.10 
2.50 2.87 
N/A N/A 
N/A N/A 
9.95 10.85 
1.09 1.87 
9.84 
6.95 9.68 
1.80 0.81 
9.14 9.83 
0.52 0.40 
8.38 9.34 
1.25 
Revision 1 
275300C 
53.94 
3.49 
48.83 
18.41 
50.39 
7.55 
41.56 
7.92 
26.93 
8.26 
33.40 
17.99 
N/A 
N/A 
17.78 
4.38 
3550C 
250 
275C 
43.98 
8.99 
40.16 
17.22 
31.23 
3.75 
27.24 
6.23 
20.65 
4.80 
26.15 
13.65 
N/A 
N/A 
15.28 
1.94 
5075C 
10.84 
0.01 0.26 
0.45 
8.50 9.16 
0.57 0.64 
7.03 8.12 
2.39 2.33 
N/A N/A 
N/A N/A 
5.24 9.20 
0.23 0.63 Table 3-15. Rock Mass Coefficients of Thermal Expansion for Repository Units 
Specimen Source 
Single Heater Testa 
Drift Scale Testb 
12 Specimensc 
Rock Unit 
Tptpmn 
Tptpll 
b BSC 2002a, Table 6.3.3.6-5. 
c DTNs: SN0208L01B8102.001; SN0211L01B8102.002. 
Mean 
4.14 
2.36 
5.88 
2.03 
2.41 
4.19 
4.40 
7.44 
9.81 
12.55 
6.50 
6.60 
10.04 
15.34 
Source: a BSC 2002a, Table 6.2.3.5-1. 
Temperature 
70C 
117C 
160C 
50C 
75C 
100C 
125C 
150C 
175C 
200C 
80C 
120C 
160C 
200C 
No. 4: Mechanical Degradation 3-27 
Revision 1 
Standard 
Deviation 
N/A 
N/A 
N/A 
1.29 
0.93 
2.07 
1.96 
0.45 
0.80 
N/A 
1.49 
1.73 
1.69 
5.58 
June 2004 INTENTIONALLY LEFT BLANK 
3-28 No. 4: Mechanical Degradation 
Revision 1 
June 2004 Revision 1 
4. DEVELOPMENT OF ROCK MASS MATERIAL MODELING APPROACHES 
FOR NONLITHOPHYSAL AND LITHOPHYSAL ROCKS 
Section 3 reviewed the basic laboratory and in situ thermal and mechanical testing data that have 
been generated by the Yucca Mountain Project for the repository host horizon units. These data, 
typically on small-scale samples, need to be generated to provide input properties and property 
ranges for design and performance analyses. This section describes the integration of geologic 
mapping and geotechnical characterization studies with the laboratory and field testing to 
produce field-scale rock mass properties. Constitutive modeling approaches are also described. 
June 2004 
4.1 MECHANICAL DEGRADATION MODELING APPROACH FOR 
NONLITHOPHYSAL ROCK 
The nonlithophysal rocks are strong, hard materials. The degradation behavior of tunnels in 
these rock units is controlled by the occurrence of keyblocks, or kinematically removable 
wedges, which can dislodge and fall under the action of external loading. Dislodging of these 
keyblocks does not necessarily lead to extensive failure and may simply result in isolated rock 
falls. Thus, isolated blocks may become dislodged, yet the excavation remains stable. 
Keyblocks in the 5-m-diameter ECRB Cross-Drift are first evident in the crown at about 
Station 10+50 in the Tptpmn unit. Most of the keyblocks in this region are of minor size and 
have typically been forcibly removed by scaling operations immediately after excavation, but 
prior to ground support installation. Keyblocks are possible in this area because of the increased 
presence of planes of weakness (i.e., a vapor-phase parting) in the near horizontal orientation that 
intersects with two opposing near vertical joint planes. The largest resultant void is 
approximately 0.5 m3 at approximately Station 11+55 as shown in Figure 4-1. No unstable 
keyblocks (i.e., those that have fallen out at a later time due to gravity) have been observed in the 
field (BSC 2004a). 
The approach taken here to represent the degradation response of nonlithophysal rocks is to 
explicitly model the fractured, blocky response of the material to allow a direct calculation of 
rockfall and opening shape change as a function of loading. This approach requires that the 
stochastic nature of the fracturing be captured in the modeling. Two items are required to 
successfully implement this approach: a tool for producing representative fractured volumes of 
rock, and numerical models that can simulate the physical, three-dimensional collapse modes of 
a blocky rock mass subjected to seismic and other loadings. A sufficient number of mechanical 
simulations12 using representative fracture realizations are necessary to adequately describe the 
full range of stochastic response of the rock mass. Uncertainty in the estimate of rockfall arises 
from three sources: 
1. The uncertainty in the knowledge of the fracture geometry 
2. The uncertainty in mechanical properties of the fractures 
3. Uncertainty in the applied loadings. 
12 A sufficient number of fracture realizations is described in Section 5.3. 
4-1 No. 4: Mechanical Degradation Revision 1 
Figure 4-1. Evidence of Key-Block Occurrence in the Enhanced Characterization of the Repository 
Block Cross-Drift, Station 11+55 
June 2004 4-2 No. 4: Mechanical Degradation Revision 1 
Uncertainty in the fracture geometry is inherently accounted for through stochastic 
representation of fracture geometries using the FracMan fracture simulation model and by 
conducting a sufficiently large number of analyses with randomly selected fracture patterns. 
Uncertainty in fracture mechanical properties is investigated by varying the properties over their 
expected ranges. Uncertainty in the applied loading is accounted for by the use of measured in 
situ stresses, inclusion of thermal loading histories, and use of 15 sets of probabilistically defined 
ground motions that account for the range of uncertainties in the ground motion (BSC 2004c). 
This section describes the development of the stochastic fracture geometry model that provides 
the input for the three-dimensional discontinuum stability model. The model and modeling 
approach is described in Section 5.3. 
4.1.1 Development of Fracture Geometries for Nonlithophysal Rock 
Analysis of seismic response and rockfall in emplacement drifts in fractured, nonlithophysal rock 
is, in general, a three-dimensional problem requiring the rock mass to be represented as an 
explicitly fractured assemblage. To achieve this objective, the 3DEC, three-dimensional 
discontinuum program (BSC 2002c) is used to model the mechanical response of a rock block 
assemblage subjected to in situ, thermal, and seismic loads. The 3DEC program allows direct 
input of the fracture geometry in creation of a synthetic rock mass composed of an assemblage 
of blocks within which emplacement drifts may be simulated. The details of the 3DEC model 
are described in Section 5.3. 
The blocks of nonlithophysal rock are significantly stronger than the in situ and thermally 
induced stresses, and thus the problem of modeling this material is essentially one of elastic 
blocks separated by fracture surfaces. Therefore, in modeling of the stability of the tunnels and 
the rockfall that may occur from the applied load, the fracture geometry and surface properties 
become of primary importance. A methodology for defining statistically representative fractures 
is therefore required as a direct input to the 3DEC program. In particular, the input fracture 
geometry must provide an adequate representation of the orientation, length, spacing and 
continuity of fractures and their variability, as this controls the size and number of removable 
blocks that surround the tunnel. Additionally, the surface characteristics, including roughness, 
planarity, and alteration/infilling define the shearing and tensile resistance of the fractures under 
load. 
The development of a stochastically defined fracture system, representative of the actual rock 
mass is accomplished using the FracMan program (USGS 1999). The existing fracture mapping 
database, described in Section 2.3.1, provides the basic input to the FracMan program, which 
develops sets of planar, circular fractures that conform to the statistical variability of the 
geometric characteristics of the input data. Statistical models are fitted to the various geometric 
characteristics of each fracture set in the database, followed by generation of representative 
fracture sets. These representative fractures are then back-checked against the statistical 
variability and geologic realism of the original sets (i.e., field data) to achieve an acceptable 
facsimile. Details of this process are described in Drift Degradation Analysis (BSC 2004a). 
A three-dimensional representative rock mass cube, 100 m on a side, is generated using FracMan 
for each Topopah Spring subunit. Each fracture is described by its centroid coordinate, dip, dip 
4-3 June 2004 No. 4: Mechanical Degradation direction, and radius. These geometric properties are used as direct inputs to the 3DEC program 
for development of a block geometry within which emplacement drifts can be randomly 
excavated. 
4.1.2 ExampleFracture Geometry Generation for the Middle Nonlithophysal Unit 
Because the large-scale fracture control of block geometry is most prevalent in the 
nonlithophysal rock, and in the Tptpmn in particular, an example of the FracMan methodology 
for construction of an equivalent fracture model for this unit is given. The analysis for the 
Tptpmn uses a classical approach to identify sets based on orientation only (Mongano et al. 
1999; CRWMS M&O 2000). 
The detailed line survey data are used to condition FracMan to develop representative fracture 
trace lengths and spacings. Table 4-1 displays the mean orientation of the sets, a comparison of 
average fracture radius converted to diameter and average trace length, and intensity (average 
spacing) from FracMan and average spacing from the detailed line surveys. 
Table 4-1. Comparison of Data from Detailed Line Survey, Full-Periphery Geologic Maps, and FracMan 
Output for the Tptpmn 
Inter-Fracture 
Distance (m) 
Set Number 
Set 1 
Observed 
Orientation 
(Strike/Dip) 
120/84 
(210/06) 
Set 2 215/88 
(305/02) 
Set 3 302/38 
(212/52) 
Observed FracMan 
0.48 
1.08 
3.40 
2.46 
0.79 
1.29 
3.16 
1.48 
FracMan 
Orientation 
(Strike/Dip) 
125/84 
214/86 
299/43 
327/08 329/14 
(239/76) 
Vapor-Phase 
Parting 
Source: DTNs for tunnel mapping include GS960908314224.020, GS000608314224.006, GS960908314224.015, 
GS960908314224.016, GS971108314224.025, GS960708314224.008, GS000608314224.004, and 
GS960708314224.010. 
NOTE: Trace length medians are taken from a compilation of tunnel mapping and synthetic tunnel samples from 
FracMan. 
A direct comparison between actual full periphery geologic maps from the ESF to synthetic full 
periphery geologic maps from FracMan is given in Figure 4-2. This comparison ensures that the 
synthetic fracture geometries are not only quantitatively validated, but similar from a geological 
perspective as well. Details of the quantitative comparison of FracMan results to the full 
periphery and detailed line surveys is given in Drift Degradation Analysis (BSC 2004a). 
4-4 June 2004 No. 4: Mechanical Degradation 
Trace Length 
Median from Full 
Periphery Geologic 
3.3 
3.1 
3.6 
3.4 
Maps (m) 
Trace Length 
Median from 
FracMan (m) 
2.8 
2.9 
3.7 
3.5 
Revision 1 Revision 1 
Source: (a) DTNs: GS990408314224.004; GS000608314224.006; GS960908314224.015; GS960908314224.016; 
(b) BSC 2004a. 
NOTE: The purpose of this figure is to illustrate the geologic structure contained on a full periphery geologic map. 
The annotated information on this figure is not intended to be legible. 
Figure 4-2. Comparison of Full Periphery Geologic Maps from the Tptpmn in the Exploratory Studies 
Facility (a) with Simulated Full Periphery Geologic Maps from the FracMan Cube (b) 
June 2004 4-5 No. 4: Mechanical Degradation Revision 1 
4.2 MECHANICAL DEGRADATION MODELING APPROACH FOR LITHOPHYSAL 
ROCK 
June 2004 4-6 
4.2.1 Material Model Requirements 
The lower lithophysal unit (Tptpll)13 is characterized by intense, small-scale fracturing. Joint 
sets are not as clearly defined as in the middle nonlithophysal (Tptpmn) unit. Average joint 
spacing is less than 1 m, and at certain locations this spacing is much smaller, on the order of 
0.1 m. In addition to fracturing on different scales, the lithophysal rock mass is characterized by 
the presence of almost uniformly distributed holes (lithophysae) of varying size (from less than 
1 cm to greater than 1 m in diameter). The lithophysae account for up to 30% of the rock mass 
volume. The size of the internal lithophysae structure and fracture spacing is much smaller than 
the drift size (i.e., 5.5-m diameter). Since the lithophysae size is small in comparison to the drift 
diameter, the impact of lithophysae on mechanical effects on drift stability can be represented 
numerically using average properties for the rock mass that take into account the porosity 
without actual modeling of individual lithophysae. Based on the fracture and lithophysae 
network, it is possible to empirically deduce that failure of this material will produce constituent 
block sizes that are controlled by the spacing of these features (i.e., blocks on the order of tens of 
centimeters on a side). 
Figure 4-3 is a core sample from the Tptpll with a diameter of 290 mm that has been removed 
from the core barrel and is still moist from drilling water. The moisture highlights the intensive 
fracture system that exists in the rock mass. These fractures are largely of a cooling origin, as 
evidenced by vapor phase alteration along the fracture faces (i.e., they are not of a mining 
origin). Drilling of core in this material without special techniques is problematic due to stresses 
placed on the core by the rotating barrel, which tend to pull it apart along these very rough 
fracture surfaces. This photograph clearly shows the ultimate small rock block sizes that would 
be created in the event that failure occurs in the lithophysal rock. Therefore, it is not the goal of 
the degradation modeling effort to attempt to predict block size (as was the case in the 
nonlithophysal rocks), but to predict the ultimate volume of detached rock, the shape of the 
excavation, and the quasi-static loading applied to the drip shield. 
13 Although the Tptpll is emphasized in this document since a majority of the repository is located in this unit, the 
results also apply to the Tptpul, which exhibits similar rock mechanical properties. 
No. 4: Mechanical Degradation Revision 1 
Figure 4-3. Large (290 mm) Diameter Core from Tptpll 
NOTE: Moisture on core surface clearly shows intensive internal fracture and lithophysae network. 
There is no preferred direction in the orientation of fracturing or lithophysae distribution that 
would justify the necessity of inclusion of anisotropy in analysis of the mechanical response of 
excavations within the Tptpll (BSC 2004a). Heterogeneity in rock mass properties resulting 
from spatial variability of the lithophysal porosity is considered in two ways in stability analyses. 
First, a series of base case parameter studies are conducted on the drift scale in which 
homogenous rock properties are assumed for a given drift cross section, but using properties that 
cover the range of potential mechanical properties. Second, analyses are also conducted in 
which the spatial variability of lithophysal porosity is accounted for by explicitly representing 
nonhomogenous distributions of rock mass properties. These analyses take into account the 
stochastic variability of rock properties but require a larger number of realizations to cover 
extreme cases. With homogeneous properties, using a range of mechanical property categories, 
conservative results can be obtained with fewer analyses. Therefore, the simple, bounding range 
approach with constant properties (method 1 above) is used for the majority of calculations, and 
explicit modeling of the estimated spatial variability (method 2) is used in a checking mode to 
verify the conservative nature of the bounding range approach. Use of a two-dimensional 
modeling approach is justified for lithophysal rocks based on the lack of anisotropy in the rock 
mass and the fact that the block sizes created upon failure are very small (essentially granular in 
nature) with respect to the size of the emplacement drift. Finally, since a two-dimensional 
modeling approach is used, the initial, in situ horizontal stress component is projected into the 
plane perpendicular to the axis of the emplacement drifts. This stress component is 
approximately 0.5 times the vertical component. 
4-7 June 2004 No. 4: Mechanical Degradation Revision 1 
To represent drift degradation mechanisms and rockfall in lithophysal rock, the mechanical 
model and the numerical method in which it is embedded must have the following capabilities: 
 The model must provide a general capability of modeling in situ stress, thermal, and 
seismic loading of the rock mass. 
 The model must represent the effects of porosity and matrix preexisting fracturing on the 
elastic and strength properties of the material. 
 The model must allow internal fracturing and detachment of the rock mass (i.e., rockfall) 
to occur in response to gravity, thermal effects, and seismic shaking. 
These requirements imply the necessity of a discontinuum approach to represent the rock mass. 
4.2.2 Mechanical Material Model and Numerical Analysis Approach Development for 
Lithophysal Rocks 
4.2.2.1 Generalization of Lithophysal Rock Mass Properties into a Number of Rock 
Quality Categories 
In the lithophysal units, lithophysal porosity is the primary physical feature impacting rock 
quality conditions (and therefore rock mass strength and stiffness). As seen in Figure 4-4, the 
laboratory data shows a range in unconfined compressive strength from approximately 10 to 
30 MPa with a corresponding range in Youngs modulus from approximately 5 to 20 GPa. The 
estimated sample lithophysal porosity varies from approximately 10% to 30% over this range, or 
is roughly comparable to the range of in situ values defined from mapping in the ECRB Cross- 
Drift (Figure 2-10). Thus, the core sampling used for the laboratory testing spans roughly the 
same range of lithophysal porosity (if not lithophysae size and shape) as observed throughout the 
ECRB Cross-Drift. 
For convenience of analysis, the mechanical rock properties range, as shown in Figure 4-4, is 
subdivided into five rock strength categories that cover the entire range of large-core 
laboratory testing and in situ testing results. Table 4-2 presents these strength and moduli ranges 
derived by subdividing the laboratory data into five categories with an unconfined compressive 
strength increment of 5 MPa. The associated Youngs modulus for each unconfined compressive 
strength is derived from the linear data fit to the 290-mm core data given in Figure 4-4. The 
approximate equivalent lithophysal porosity for each of these ranges is given in Table 4-2 
(BSC 2004a, Section E.4). 
June 2004 4-8 No. 4: Mechanical Degradation Source: DTNs: SN0208L0207502.001; SN0211L0207502.002; SN0305L0207502.006. 
NOTE: Base case strength and modulus based on fit to 11.5-in. (290-mm) core data. Data for this plot are from 
Figure 3-11. 
Figure 4-4. Relationship of Unconfined Compressive Strength to Youngs Modulus from Large Core 
Testing of Lithophysal Rock (Figure 3-11) and Assignment of Five Average Quality 
Categories 
Table 4-2. Suggested Range of Mechanical Properties Selected for Base-Case Design and 
Performance Analyses 
Category 
1 
2 
3 
4 
5 
Unconfined 
Compressive 
Strength (MPa) 
10 
15 
20 
25 
30 
Source: BSC 2004a, Table E-10. 
NOTE: aThe calculation of Youngs modulus values is documented in BSC 2004a, Section E.4.1. 
bApproximate lithophysal porosity estimates provided in BSC 2004a, Section E.4.1. 
No. 4: Mechanical Degradation 
Estimated Youngs 
Modulusa 
(GPa) 
1.9 
6.4 
10.8 
15.3 
19.7 
4-9 
Approximate Lithophysal 
Porosity From Laboratory Testsb 
(%) 
35 +/- 8 
28 +/- 6 
21 +/- 4 
13 +/- 5 
7 +/- 7 
Revision 1 
June 2004 Revision 1 
It is considered that, by conducting numerical analyses with this entire range of data, that all 
levels of rock quality and rock mass response from lowest to highest porosity ranges and size 
effects can be covered. A histogram constructed to provide the information on the percentage of 
each rock category in the Tptpll as a whole is shown in Figure 4-5. The distribution of 
lithophysal porosity is based on the lithophysal mapping data in the ECRB Cross-Drift (BSC 
2004a, Appendix O). It is shown that approximately 90% of the mapped rock quality is equal or 
better than the Category 3 rock. Categories 1 and 2 are related to localized, stratiform, high 
porosity zones that occur particularly near the top of the Tptpll. Although of low abundance, 
Categories 1 and 2 are included in the range of rock properties for analysis to test the 
conservative bounds of rock strength. Figure 4-5 also shows photographs of example panel 
maps that illustrate lithophysal porosity levels characteristic of Categories 3, 4, and 5. 
The rock strength categories that form the basis for parametric stability calculations are based on 
a linear fit to the strength-modulus data from 305-mm diameter core testing. Properties are given 
in Section 4.2.3.3. 
June 2004 4-10 No. 4: Mechanical Degradation Revision 1 
Source: (a) BSC 2004a, Appendix O, Section O.6.6. 
NOTE: Lithophysal porosity data are based on 183 5-m traverse locations from the ECRB Cross-Drift between 
stations 14+44 and 23+26. Examples of approximate rock strength category levels taken from 1-by-3-m 
panel maps: Category 3 (a) with lithophysal porosity of approximately 19%; Category 4 (b) with lithophysal 
porosity of 13.3%; and Category 5 (c) with lithophysal porosity of 8.5%. 
Figure 4-5. Distribution of Lithophysal Porosity and Estimated Rock Properties Categories for the Tptpll 
in the Enhanced Characterization of the Repository Block Cross-Drift 
June 2004 4-11 No. 4: Mechanical Degradation Revision 1 
4.2.2.2 Continuum- and Discontinuum-Based Approaches to Representing Rock Masses 
A standard approach for solving excavation stability problems in geotechnical engineering is the 
use of numerical models based on continuum mechanics assumptions (Figure 4-6). Such an 
approach is quite effective if the rock mass, in response to stressing, eventually arrives at a state 
of mechanical stability and where the primary purpose of the modeling is the computation of 
stress redistribution around an opening or determination of the final displacement profiles. 
However, difficulties are encountered if a continuum model is used for prediction of a 
mechanical system (i.e., a tunnel) that does not arrive at a stable condition upon excavation. 
Source: BSC 2004a, Figure 7-7. 
NOTE: The continuum approach models yield of the rock through use of a material model that enforces plasticity 
relations (note marked elements). Rock breakage and separation are not possible in this approach. The 
discontinuum approach also represents the rock mass using similar material models but provides the 
capability for the rock mass to fracture and break apart on potential fracture surfaces. The illustration on the 
right contains a simple rectangular representation of a drip shield. 
Figure 4-6. Continuum (Left) and Discontinuum (Right) Approaches to Modeling Drift Stability 
Continuum models use stress-strain relations to describe the mechanical behavior of a material. 
In rock, the mechanical effects of fractures and other features are typically assumed to be taken 
into account within a standard form of plasticity model without explicitly representing the 
features (e.g., see Hoek 2000). The specific characteristics of the rock mass, such as the rock 
type, the spacing and continuity of the fractures, and the roughness and alteration of the fracture 
surfaces, are taken into account by empirical adjustment of properties determined from testing of 
intact rock cores. 
A linearly elasticperfectly plastic material model with Mohr-Coulomb yield criteria is a 
constitutive model often used to represent mechanical behavior of a rock mass (e.g., Hoek 2000). 
Because the material strength of a perfectly plastic Mohr-Coulomb material under general 
conditions does not decrease as a function of plastic deformation, this model will show 
June 2004 4-12 No. 4: Mechanical Degradation Revision 1 
indications of material yielding (i.e., plastic deformation) in different portions of the model, but 
will never actually predict instability that leads to rockfall. In order to predict rockfall, it is 
necessary to use a form of material model in which strength degrades as a function of 
deformation after the peak-strength of the material has been reached. Because representation of 
fracture and breakage in a continuum representation is speculative, a typical continuum modeling 
approach is not the primary method used to represent mechanical degradation and rockfall. 
Continuum models are, instead, used as a means of comparison to other approaches. 
4.2.2.3 Development of a Discontinuum Approach to Representing the Mechanical 
Response of the Lithophysal Rock Mass 
The estimation of rockfall requires that the modeling technique and mechanical material model 
be capable of representing fracture and separation of the intact rock mass into individual blocks 
of material. In particular, an estimate of the size distribution of particles is desired. Thus, the 
modeling technique must be based on use of a discontinuum numerical method (i.e., a method in 
which the rock is represented as blocks separated by fracture surfaces in which slip and 
separation of contacting rock blocks can be estimated). 
The development of a mechanical material model and estimate of property ranges for the 
lithophysal material is based on use of the laboratory database described in Section 3 (primarily 
compression testing on cores with diameters of 290 mm) supplemented by numerical model 
extrapolation using the PFC and UDEC discontinuum programs. These extrapolation approaches 
are conducted in parallel as alternative numerical methodologies. 
The database available for properties definition and model development includes: (1) uniaxial 
and triaxial compression and tensile testing of nonlithophysal rock; and (2) uniaxial compression 
testing of large-scale cores and in situ blocks of lithophysal tuff. As discussed in Section 3, the 
material properties of lithophysal rock are size-dependent due to the influence of the lithophysal 
pore space. Due to the difficulties in both obtaining and testing large samples of this rock, the 
laboratory testing data base is limited. However, it is possible to calibrate appropriate numerical 
models to reproduce the laboratory data. These models can then be used to extrapolate from the 
laboratory to the in situ scale for examination of the impact of the variability of lithophysal size, 
shape, and distribution on the range of rock mass strength and deformability. Thus, the model is 
used as a means of extending the laboratory data base. 
This strategy, which combines laboratory testing, model calibration, and extrapolation for 
estimation of mechanical properties ranges is illustrated in Figure 4-7. In this section, an 
approach is described in which physics-based discontinuum numerical modeling programs 
PFC2D (BSC 2004h), PFC3D (BSC 2004i), and UDEC (BSC 2002d)are used as numerical 
laboratories to simulate and test the basic deformation and failure response mechanisms of 
lithophysal tuff. These programs were chosen due to their ability to simulate the physics of 
deformation and fracture of a bonded granular matrix that contains void space of varying shape, 
size, and porosity. Using two different approaches provides a check and greater confidence in 
the modeling. The UDEC program is additionally used as it allows constituent grains that are 
nonspherical in shape, and thus overcomes some simplifications used in the PFC approach. 
Specifically, it allows greater flexibility in modeling failure mechanisms under triaxial 
compression. 
June 2004 4-13 No. 4: Mechanical Degradation Revision 1 
Figure 4-7. General Approach for Estimation of Lithophysal Rock Property Bounds 
The approach has two steps. First, the programs are validated (Section 4.2.3) against the existing 
laboratory compression data, where it is demonstrated specifically that a detailed understanding 
of the basic physical mechanisms of the rock mass behavior can be obtained without resorting to 
empiricism or complex constitutive modeling. Next, the model is used to extend the laboratory 
data by conducting numerical experiments on simulated samples of lithophysal tuffs at various 
physical conditions of porosity, lithophysae shape and distribution, as well as various levels of 
confinement and applied stress. The outcomes of the modeling are estimates of the range of rock 
mass strength and stiffness for varying conditions of lithophysal porosity, size, shape, and 
distribution. Additionally, the results provide a means of understanding the size-scaling and 
variability issues introduced by lithophysae shape and distribution and their impact on rock mass 
properties and failure criteria. 
The material model developed from the testing and PFC/UDEC extrapolation is embedded into a 
drift-scale UDEC model for conducting the parametric performance analyses of emplacement 
4-14 June 2004 No. 4: Mechanical Degradation Revision 1 
drift stability. UDEC and PFC have a significant history of use in the mining and defense 
industries for analysis of tunnel stability and rock fracture simulation under in situ and seismic 
loading. The PFC2D, PFC3D, and UDEC programs have been qualified for use on the Yucca 
Mountain Project. 
4.2.3.1 The Particle Flow Code Model 
The PFC approach represents rock as a number of small, rigid, spherical grains that are bonded 
together at their contacts with shear and tensile strength, as well as a grain-to-grain friction angle 
after the contact bond has been broken. Details on the mechanics of the PFC program are 
provided in Itasca SoftwareCutting Edge Tools for Computational Mechanics (Itasca 
Consulting Group 2002). The deformability of the contacts between particles is represented by 
normal and shear stiffness at the contact point. Porosity is developed naturally in the model by 
control of the shape and size of void space between chains of bonded grains. The contact 
properties and porosity distribution are referred to as microstructural properties. Thus, the 
input conditions necessary for the model are very simple, only contact strength and stiffness. 
However, extremely rich constitutive behavior may develop naturally based on void porosity and 
the few straightforward input properties and their variability throughout the rock. 
When load is applied to the grain assembly, forces are transmitted across contacts. If the shear 
or tensile strength of the contact is reached, failure will occur, and the adjacent particles are free 
to slide past one another, or to separate. In either case, a fracture is formed and the forces must 
reorient in some fashion, thus redistributing loads. Realistic failure mechanisms may then 
develop which can be compared to those observed in the laboratory. Calibration of the model 
against laboratory testing is accomplished via sensitivity studies in which the contact strength 
and stiffness values are varied and the macroscopic stress-strain response is compared to that 
monitored. 
The UDEC approach, although similar, is different specifically in that the grains may be of any 
arbitrary shape and size, and the contacts between grains are not point force contacts, but contact 
across a plane. Additionally, the UDEC grains may be deformable rather than rigid. The 
importance of this distinction is described later. 
4.2.3 Validation of the Particle Flow CodeA Micromechanical Model Representation of 
the Mechanical Behavior of Lithophysal Rock 
June 2004 
4.2.3.2 Particle Flow Code Model Calibration 
The PFC model has been calibrated against the laboratory testing data discussed in Section 3. 
Sample numerical compression experiments were conducted for nonlithophysal and lithophysal 
tuffs using the same matrix properties (Figure 4-8). These matrix mechanical properties were 
derived by calibration of strength and modulus from the laboratory testing for a midrange 
lithophysal void porosity of approximately 15% (Table 4-2). 
4-15 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004j, Section 6.5.6, Figure 6.5-20. 
NOTE: Stress strain curves have the same scale, so strength and modulus difference in lithophysal sample is 
evident. Red and black lines on samples are tensile and shear fracture, respectively, that have formed 
within the rock samples as they are loaded. Nonlithophysal sample shows typical macroscopic shear 
failure mechanism whereas lithophysal sample fails due to tensile splitting between voids. Calibration 
results are shown in Figure 4-9. 
Figure 4-8. Particle Flow Code Calibration Compression Experiment Samples and Their Respective 
Uniaxial Stress-Strain Curves for (a) Nonlithophysal and (b) Lithophysal Rock 
June 2004 4-16 No. 4: Mechanical Degradation Revision 1 
The nonlithophysal samples show a typical shear failure mechanism as evidenced by the 
coalescence of extension cracks to form major shear fractures. The model also shows highly 
linear (elastic) response just to the point of rock failure, followed by a brittle postpeak failure 
response due to the uniformity of bond strength set in the samples. This replicates the observed 
laboratory behavior of nonlithophysal samples. 
Once the basic behavior of the nonlithophysal rock is represented, the samples can be populated 
with pores. Initially, simple circular holes were added to the nonlithophysal model to represent 
the lithophysal rock. The circular holes were added with a random spatial distribution to the 
model, and the correspondence between laboratory-derived strengths, modulus and porosity were 
examined. The model showed good comparison with general magnitude and trend of laboratory 
strength and modulus data (Figure 4-9). 
A conclusion of this initial work with a simplistic void porosity model is that the primary 
strength-decreasing effect of the lithophysae is due to the formation of tensile splitting between 
neighboring lithophysae under compressive load. As porosity increases, the spacing between 
lithophysae decreases, and thus a greater propensity for tensile splitting at lower applied forces 
results due to the smaller solid bridges between voids. 
The tensile splitting mechanism results in increasingly less-brittle postpeak response with 
increasing porosity. Additionally, the same matrix strength provides a reasonable fit to both 
nonlithophysal and lithophysal laboratory data (i.e., the void porosity is the primary driver in the 
mechanical properties reduction and not mineralogical differences in lithophysal and 
nonlithophysal rocks). 
Further numerical testing was performed in which all of the actual lithophysal panel maps 
generated in the ECRB Cross-Drift (e.g., BSC 2004a, Appendix O) were discretized and used as 
compression test specimens for the calibrated model. As seen in Figure 4-10, the complex 
shape, size and distribution of lithophysal voids results in the same general failure mode (tensile 
splitting between voids) and same general trend of strength and modulus to void porosity 
(Figure 4-9); however, the variability tends to be greater primarily due to shape and distribution 
of voids. This is particularly true for samples whose lithophysal porosity is lower (e.g., around 
10% to 15%). A few large voids with uneven distribution within a sample can result in lower 
strength, whereas widely spaced voids in a finite sample size can result in higher strength for the 
same sample with uniformly distributed voids. 
June 2004 4-17 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2003a, Figures 9-22 and 9-23. 
NOTE: Greater variability in the strength of the PFC samples for a given porosity is seen when true lithophysae 
shapes (stenciled) are introduced. This effect is particularly large at low porosities due to the greater 
variability of distribution and solid bridge lengths between lithophysae in a given sample volume. Test data is 
from large-core lithophysal tests and is the same as in Figure 3-12. 
Figure 4-9. Plots Showing Data from Large Core Compression Testing of Tptpul and Tptpll Compared 
to Particle Flow Code Simulations Using Circular Shaped Lithophysae as well as Actual 
Stenciled Shapes from Enhanced Characterization of the Repository Block Panel Maps 
June 2004 4-18 No. 4: Mechanical Degradation Source: BSC 2004j, Section 6.5.6, Figure 6.5-20. 
Figure 4-10. Examples of Particle Flow Code Compression Tests Using Simulated Rock Specimens 
Developed by Stenciling Field Panel Maps in the Enhanced Characterization of the 
Repository Block 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 4-19 Revision 1 
The result of all strength testing on both the simplest, circular void cases and the more complex, 
realistic shapes can be summarized as follows: 
 The PFC model shows that the mechanism for strength reduction from nonlithophysal to 
intact rock results from tensile splitting of solid rock bridges between lithophysal voids. 
The smaller these rock bridges (i.e., the greater the porosity), the greater the strength 
reduction. 
 The PFC model and laboratory data show a reasonable agreement with respect to 
strength and modulus change with lithophysal porosity. Strength and modulus show a 
generally logarithmic decrease with lithophysal porosity. 
 Models with circular voids that are randomly distributed through the rock mass show 
less scatter and more uniform relation of mechanical properties to lithophysal porosity. 
 Models with lithophysal voids with complex, irregular size and shape distribution show 
the same general trends as the circular void models, but with greater scatter of the 
strength or modulus. This scatter may partially account for the obvious size effect 
observed in the laboratory and in situ scale testing. 
The conclusions from the PFC model calibration are that a viable tool has been developed for 
simulation of the mechanical response of lithophysal tuff to stressing and that this tool can be 
used, in addition to field and laboratory testing, to study variability in material response. 
4.2.3.3 Extrapolation of Lithophysal Mechanical Response Using the Particle Flow 
Code Program 
The calibrated PFC program is used as an extrapolation tool to examine the impact of 
lithophysae size, shape, porosity, and distribution on mechanical properties. Parametric studies 
have been conducted with simple (circular and triangular) lithophysae shapes, as well as with 
actual, complex shapes and distributions digitized directly from ECRB Cross-Drift panel maps. 
The results of these studies, shown in Figure 4-11, are used as a guide to understanding the 
variability of mechanical properties for a given porosity range. As seen in this plot, the results of 
the calibrated PFC shape studies from all ECRB Cross-Drift panels are overlaid on the large-core 
laboratory data for lithophysal rock showing the relationship of unconfined compressive strength 
to Youngs modulus. Estimated upper and lower bounds of all the properties, including the 
results of the PFC extrapolations, are given on this plot. These upper and lower bounds are 
simply drawn to encompass all of the laboratory and extrapolated data, but the lines representing 
the bounding values are dashed outside the range of measured values. The range of data 
generated by the PFC extrapolations for size, shape, and distribution variability fall within the 
range of the laboratory testing with the exception of the highest levels of lithophysal porosity 
(e.g., less than 20%). It is seen that the saturated strength generally forms the lower bound of the 
data range, with a minimum strength value of approximately 10 MPa. The lower bound has been 
cut off at this 10-MPa strength level. This cutoff value is based on comparison of observations 
of current drift stability in the ECRB Cross-Drift and the ESF to drift-scale modeling studies 
performed assuming in situ stress conditions. Models indicate that extensive yield of the 
springlines of the ECRB Cross-Drift would be observed if the unconfined compressive strength 
June 2004 4-20 No. 4: Mechanical Degradation Revision 1 
were below 10 MPa (BSC 2004a, Appendix E.4.1]). This is obviously not the case as the ECRB 
Cross-Drift shows stable conditions with generally unsupported sidewalls. Therefore, the PFCextrapolated 
strength values below the laboratory data range appear to be related to local zones 
of high porosity. The impact of the spatial variability of lithophysal properties on drift-scale 
rock mass properties is discussed in Section 5.3. 
Source: BSC 2004a, Appendix E.4.1. 
NOTE: The bounding curves are empirically derived to include all large core laboratory data, including 267-mm 
diameter saturated cores from Busted Butte. Base-case properties characteristic, derived from 305-mm 
diameter cores, are also shown. Extrapolated (PFC) strength values below 10 MPa are localized and 
inconsistent with observations of lack of yielding in ECRB Cross-Drift and ESF. 
Figure 4-11. Estimated Upper and Lower Bounds of Unconfined Compressive Strength and Youngs 
Modulus for Lithophysal Rock 
4.2.4 Extrapolated Triaxial Behavior of Lithophysal Rock and Estimation of Rock Mass 
Constitutive Response 
The previously described PFC calibration and analyses describe the basic comparison of the 
unconfined compression behavior and modulus of lithophysal rocks. As was discussed in 
Section 3, triaxial laboratory experiments have not been conducted on representative lithophysal 
samples due to the necessary size of the sample needed for lithophysal rock and the associated 
difficulties in obtaining pressure vessels and confining jacketing systems for samples with 
cavities. 
June 2004 4-21 No. 4: Mechanical Degradation Revision 1 
An understanding of the confining pressure response is necessary for verification or for 
determination of a proper yield criteria for the lithophysal rocks to develop an understanding of 
how the basic rock mass failure parameters (rock mass cohesion, angle of internal friction, and 
dilatancy) are affected by lithophysal porosity. UDEC, a discontinuum program, similar in many 
respects to PFC, is used for this purpose due to its more general particle shape capability. The 
circular particle geometry employed by PFC simplifies the numerical algorithms, making it a 
good tool for conducting many parameter studies as shown in the previous calibration studies. 
However, the particle geometry also restricts the ability to examine postpeak failure 
mechanisms in detail due to dilational response of a circular particle modelthis is not the case 
for the more general UDEC approach. 
A series of uniaxial (extension and compression), triaxial compression experiments was 
conducted on the modeled samples, with circular lithophysal voids added randomly to create 
porosities of 10.3%, 17.8%, and 23.8%. These porosity values cover the approximate range of 
lithophysal porosity that spans the range of rock strength categories observed in the field panel 
maps. Figure 4-12 shows an example stress strain response for the case of 17.8% lithophysal 
porosity. The stress strain response for tensile testing and compression at a number of confining 
pressures are shown. Adjacent to each of the stress-strain curves is a figure of the sample in the 
failed state. Rock failure results from fracturing between the voids in a similar fashion as that 
demonstrated previously for the PFC modeling. 
The general behavior is typical for brittle rock materialsincreasing strength with confinement 
and conversion of the failure mode from axial splitting (in unconfined compression) to shear 
failure as the confining pressure increases. The material response is elastic-brittle at low 
confinement and elastic-plastic at higher confining levels. 
June 2004 4-22 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2003a, Figure 9-34. 
NOTE: Failure state of samples showed failed state for a particular confining pressure: counterclockwise from left 
are direct tension, unconfined, 1, 3 and 5 MPa confining stress, respectively. Circular voids distributed 
randomly throughout the sample. 
Figure 4-12. Stress-Strain Response and Failure Mechanisms for Lithophysal Porosity of 17.8% as 
Predicted by the UDEC Model 
June 2004 
4.2.5 Estimation of Linear and Nonlinear Failure Envelopes 
The UDEC peak strength values for the numerical testing results can be used to construct 
traditional failure envelopes for the nonlithophysal as well as lithophysal samples. Figure 4-13 
shows the peak strength values plotted in principal stress space with approximate Mohr-Coulomb 
(linear) and Hoek-Brown (nonlinear) failure envelopes fit to the results. Note that in each case, 
multiple UDEC simulations were made for each minimum principal (confining) stress level in 
which different random distributions of UDEC grain structure were used. The Mohr-Coulomb 
and Hoek-Brown strength parameters derived from the fits to these data are given in Table 4-3. 
As seen, the primary effect of increasing lithophysal porosity is a reduction in the unconfined 
compressive and tensile strength components. There is little apparent impact of lithophysal 
porosity on the internal angle of friction of the material, which appears reasonable from a 
physical standpoint. These data can be used for definition of failure properties of the rock mass 
for performance analyses. 
4-23 No. 4: Mechanical Degradation As seen in Figure 4-13, the best-fit Hoek-Brown envelopes (Hoek 2000, p. 179) for the UDEC 
simulations, yield nearly the same principal stress relations as the linear Mohr-Coulomb 
Estimated 
Hoek-Brown 
mi 
6.6 
7.7 
7.3 
5.0 
Estimated 
Hoek-Brown 
UCS (MPa) 
58.5 
25.1 
16.5 
14.0 
Estimated 
Cohesion 
(MPa) 
14.9 
6.4 
4.1 
3.9 
envelopes in the range of confining stresses given here. 
Source: BSC 2003a, Figure 9-40. 
NOTE: HB = Hoek-Brown 
Figure 4-13. Hoek-Brown Failure Envelopes Based on UDEC Simulations for Various Lithophysal 
Porosities 
Table 4-3. Summary of Average Strength and Modulus for Variation In Lithophysal Porosity and Mohr- 
Coulomb (Linear) and Hoek-Brown (Nonlinear) Yield Criteria Parameters as Derived from 
UDEC Simulations 
Lithophysal 
Porosity 
(%) 
0 
10 
17 
24 
Estimated 
UCS (MPa) 
58.7 
25.1 
15.5 
13.2 
Source: BSC 2003a. 
NOTE: Friction angle and cohesion are derived from best-fit linear envelope to data. Hoek-Brown s value is 
assumed to be 1 in deriving the Hoek-Brown estimated UCS values in this table, as the simulations are 
assumed to represent intact samples. Lithophysal porosities in this table can be related to in situ Tptpll 
rock mass properties categories via Figure 4-5. UCS = unconfined compressive strength. 
(GPa) 
19.8 
14.2 
11.2 
9.3 
Estimated 
Youngs 
Modulus 
Estimated 
Friction 
Angle () 
36 
36 
35 
29 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 4-24 Revision 1 
4.2.6 Summary of Material Model Development 
The following summarizes the material model and properties for analysis: 
 A set of base-case rock mass strength and moduli were developed from large core 
laboratory testing. The properties were subdivided into a series of categories that span 
the range of in situ lithophysal porosities/rock qualities. These categories were related 
to approximate lithophysal porosities observed in existing tunnels at the Yucca 
Mountain site (Figure 4-5), providing a link between observation and rock quality. This 
data provides the basis for design and performance parameter analyses and for 
establishing an understanding of uncertainty in the analyses. 
 The PFC and UDEC programs were calibrated to reproduce the basic strength and 
moduli as functions of lithophysal porosity as well as to provide a basic understanding 
of the mechanics of yield in this material. The models were used to extrapolate the 
mechanical properties for lithophysal rock masses over the range of in situ conditions 
determined from field mapping. These generated-data provide an estimate of the range 
of variability of the properties. 
 Standard forms of yield conditions for rock masses (Mohr-Coulomb and Hoek-Brown) 
were fit to the laboratory and model extrapolation data, and their standard strength 
parameters were determined. This information, combined with the unconfined 
compressive strength and modulus ranges, provides the basis for development of 
drift-scale stability simulation models. 
The material model can now be embedded into a general numerical method that is capable of 
simulating drift-scale problems and estimate rock mass degradation. This is described in the 
next section. 
4.2.7 Development and Validation of a Drift-Scale Modeling Method for Lithophysal 
Rock Using the UDEC Program 
Although the UDEC program has been shown to provide reasonable agreement to large-scale 
laboratory compression testing of lithophysal rock, it is not practical to attempt to model a full 
drift-scale problem with explicitly represented lithophysal cavities. The model would be 
excessively large with correspondingly large computational run times. However, it is not 
necessary to attempt to model individual lithophysae as long as the overall mechanical response 
of the rock mass behaves according to the material models and their strength and deformability 
parameters derived in the previous section. 
Development of the drift-scale modeling approach for lithophysal rock requires calibration. The 
method used here is, first, to calibrate the rock mass material model and mechanical properties 
such that the UDEC discontinuum program (without explicit voids) is able to reproduce the 
observed large-scale laboratory moduli and strength behavior. Next, the model is validated 
against (1) observations of failure mechanism in the laboratory, (2) field observations of tunnel 
response in the ECRB Cross-Drift, and (3) thermally induced fracture development in the DST 
conducted in the ESF in the Tptpmn unit. 
June 2004 4-25 No. 4: Mechanical Degradation Revision 1 
Figure 4-14 provides a flow chart illustrating the calibration and validation strategy for the driftscale 
UDEC model. The following section describes the validation of the UDEC model and 
exploration of its limitations. 
4.2.7.1 Drift-Scale Model Calibration 
The drift-scale UDEC lithophysal model represents the rock mass as an assembly of polygonal, 
elastic blocks (Figure 4-15) bonded together across their boundaries to form a coherent solid. 
The goal is to provide a rock mass in which the overall mechanical behavior of the mass is 
consistent with the material model developed for the lithophysal rock, yet allow internal 
fracturing to form and blocks to loosen and detach as the evolving thermal and dynamic stress 
state dictate. In other words, the fractures are invisible to the model until yielding begins. 
Figure 4-14. General Approach to Validation of Mechanical Material Model for Lithophysal Rocks 
June 2004 4-26 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 7-20. 
NOTE: Blocks are bonded at their contacts with a cohesion and tensile strength. When these break, the contacts 
become purely frictional. Specimen is sampled from equivalent rock mass representing the Tptpll. 
Figure 4-15. UDEC Lithophysal Rock Specimen Composed of Many Irregular Blocks with Roughly 
Equal Side Lengths 
June 2004 4-27 No. 4: Mechanical Degradation Revision 1 
Following calibration, the model can be used to conduct additional simulations under biaxial 
compression and tension to produce the yield criteria for the material, which can then be 
compared to that generated using the grain-scale models with explicitly modeled lithophysae in 
Section 4.2.4. Further details on the UDEC model calibration can be found in Drift Degradation 
Analysis (BSC 2004a). 
4.2.7.2 Validation of the Numerical Model 
Once calibrated, the UDEC model and its constituent properties require validation against field 
observations and testing. The model was validated against laboratory failure mechanisms and 
drift scale response through: 
 Comparison of lithophysal test specimen failure mechanisms in the laboratory 
 Comparison of the prediction of drift scale fracturing in the Tptpll at ECRB Cross-Drift 
depth to observations of tunnel sidewall fracturing in the ECRB Cross-Drift 
 Comparison of roof spalling in the DST in the Tptpmn during thermal overdrive 
experiments to UDEC model predictions 
 Comparison of several different numerical modeling techniques to UDEC for a field 
simulation of a liner-reinforced tunnel to dynamic loading from a blast. 
The details of the validation comparisons are given in Drift Degradation Analysis (BSC 2004a). 
The validation tests showed that: 
 The model is capable of reproducing the basic axial splitting fracturing and failure mode 
of lithophysal samples observed in uniaxial compression, while reproducing the proper 
strength and moduli. 
 The model is capable of reproducing the springline shear fracturing observed 
occasionally in the Tptpll in the ESF and ECRB Cross-Drift. Figure 4-16 shows a 
photograph of the wall-parallel fractures occasionally observed in the large-diameter 
holes drilled for sampling at the springline. Although not common, this form of 
localized yielding has been observed in a few boreholes in the Tptpll. This form of 
fracturing behavior, typical of local in situ stress-related yield in hard rocks, is found 
only directly at the springline and extends to a depth of about 0.5 m into the sidewall. 
Due to the high vertical-low horizontal stress state, the greatest stress concentration 
occurs at this location directly at the springline. The UDEC model reproduces the local 
wall-parallel fractures and yield for strength Category 1, representative of localized, high 
porosity material. The model shows no fracturing for strength categories 2 or higher. 
This is consistent with the typical condition within the ESF and ECRB Cross-Drift, 
which shows no observation of sidewall yield. This observation is also consistent with 
the general lithophysal porosity mapped within the ECRB Cross-Drift (Figure 4-5), 
showing that the mean porosity is indicative of a strength category of 3, with about 97% 
being greater than a Category 1. In other words, the small number of observations of 
4-28 June 2004 No. 4: Mechanical Degradation Revision 1 
sidewall spalling is consistent with the relatively infrequent occurrence of low strength 
categories. 
 The model is capable of reproducing the roof crown spalling behavior observed during 
the thermal overdrive portion of the DST in the Tptpmn. In this test, the rock mass 
temperatures were driven to approximately 200C at the conclusion of the test, at which 
time minor spalling in the form of plate-shaped fragments was observed in the crown of 
the tunnel. The UDEC model was calibrated against the large-sample strength 
(Figure 3-5) and in situ modulus (from plate bearing tests) of the Tptpmn in a similar 
manner to that described above for the Tptpll. The DST was modeled by imposing the 
in situ stress state and temperature history (as determined from field temperature 
measurements) onto the model. The roof crown-parallel fracturing extent and apparent 
mechanism were reproduced at the proper temperature levels without need for 
adjustment of properties (Figure 4-17). It was found that the fracturing was the result of 
the large horizontal, thermally induced stresses in the immediate roof. The crown is 
placed in a state of uniaxial compression, and when this stress reaches the strength of the 
material, extensional, wall-parallel fracturing occurs over a limited roof span. The 
system equilibrates when fracturing extends to a short depth. 
Source: BSC 2004a, Figure 7-28. 
NOTE: The model shows that sidewall fracturing and yield does not occur for strength categories 2 or higher, which 
is consistent with the general observation of no sidewall fracturing. Localized fracturing is occasionally 
observed to a sidewall depth of about 0.5 m. The model reproduces wall-parallel springline fracturing to the 
proper sidewall depth for the lowest strength category only. 
Figure 4-16. Validation of the UDEC Numerical Approach for Lithophysal Rock against Observation of 
Springline Wall-Parallel Fracturing in the Exploratory Studies Facility and Enhanced 
June 2004 
Characterization of the Repository Block 
4-29 No. 4: Mechanical Degradation Revision 1 
Source: DTNs: MO9807DSTSET01.000; MO9906DSTSET03.000; MO0001SEPDSTPC.000; 
MO0007SEPDSTPC.001; MO0107SEPDSTPC.003; MO0202SEPDSTTV.001; MO0002ABBLSLDS.000. 
NOTE: Temperatures (a), fractures (b), minimum principal stress (c), and maximum principal stress (d). 
Temperatures were obtained from field thermal measurements; fracture and stress levels were predicted 
from the UDEC model. Conditions represent 3 years of heating in the heated drift. The temperature 
contours and induced fractures in (a) and (b) are outcomes of the model. 
Figure 4-17. Results from Heated Drift Analysis with Temperature Conditions after 3 Years of Heating 
June 2004 4-30 No. 4: Mechanical Degradation Revision 1 
4.3 ROCK MASS MODELING APPROACH FOR GROUND SUPPORT ASSESSMENT 
AND DESIGN 
Postclosure modeling methods used for nonlithophysal and lithophysal rock masses, described in 
Sections 4.1 and 4.2, respectively, are based on the use of discontinuum numerical methods. 
These methods represent the rock mass as being composed of elastic rock blocks separated by 
fracture planes. Strength and modulus properties are assigned separately to the blocks and the 
fractures and the model accounts automatically for the interaction of blocks across the fractures 
and shearing or tension failure of the fractures during the simulation of applied stresses. 
Another, more common approach to modeling fractured rock, is as a continuum material in 
which fractures are not modeled explicitly. Instead, the continuum model is assigned rock mass 
elastic and strength properties that include the overall effect of fractures and solid rock. Both 
approachescontinuum and discontinuumproduce the same approximate results (i.e., stress 
distributions and deformation, depth of yield, overall assessment of stability) as long as 
equivalent properties are used. The type of modeling approach used depends on the objectives of 
the modeling. Discontinuum methods, as used for postclosure analysis, are required in this 
application since the objective is to calculate estimates of size and amount of rockfall and the 
ultimate profile of the tunnels in response to in situ, thermal and seismic loading. Since 
continuum models do not explicitly represent fractures and their impact on rockfall, they are not 
capable of providing the information required for assessment of postclosure issues. However, 
for general assessment of excavation stability and ground support design, the continuum 
approach is reasonable, and has the advantage of simpler model setup and shorter analysis time. 
A number of the RDTME Key Technical Issue agreements deal with preclosure issues and, 
specifically, with the assessment of rock mass properties and modeling methods for ground 
support design and analysis (e.g., Appendices B, D, F, G, and H). In these appendices, a 
summary of analyses performed with continuum-based numerical models for assessment of 
preclosure drift stability and ground support design is given. Rock mass properties for the 
lithophysal rock for these analyses are derived from the analyses presented in Section 4.2, and, 
specifically, the rock mass categories and strength and modulus ranges summarized in 
Figure 4-11. 
Rock mass properties for the fractured nonlithophysal rock for these analyses were derived from 
the commonly used estimation method of Hoek and Brown (e.g., see Hoek, 2000). This method 
involves deriving estimates of rock mass mechanical properties from geotechnical rock mass 
classification, along with results from laboratory mechanical strength testing. The detailed 
derivation of the rock mass properties is described in detail in Subsurface Geotechnical 
Parameters Report (BSC 2003a, Section 8.5.2), and summarized here. 
Using the Q rock quality designation system described by Hoek (2000), the U.S. Bureau of 
Reclamation and U.S. Geological Survey conducted geotechnical mapping of the Tptpmn in the 
ESF and ECRB Cross-Drift during the geological mapping immediately behind the advancing 
TBM. From the rock quality classification, the HoekBrown geologic strength index (GSI) was 
determined. Rock mass moduli, rock mass strength, and both HoekBrown and MohrCoulomb 
failure criteria parameters are derived from the GSI. Details of the relationships between the 
GSI, the rock mass modulus, compressive strength, and failure criteria parameters can be found 
June 2004 4-31 No. 4: Mechanical Degradation in Subsurface Geotechnical Parameters Report (BSC 2003a) or in the description of Hoek 
(2000). The results of these calculations show that the GSI for the Tptpmn follows a typical 
normal distribution with mean value of 59, a median rock mass modulus of 17 GPa and median 
rock mass compressive strength of 44 MPa. Table 4-4 provides a summary of the estimated in 
situ rock mass properties derived from the HoekBrown approach. For parametric evaluation, 
the variability of rock mass quality has been subdivided into five categories at 10th, 30th, 50th, 
70th, and 90th percentile of the cumulative probability of occurrence within the Tptpmn. 
Table 4-4. Rock Mass Mechanical Properties Estimates for the Tptpmn 
Rock Quality 
Category 2 
30% 
Rock Mass Parameter 
Associated Cumulative 
Probability of 
Occurrence 
Rock Mass Cohesion, C 
(MPa) 
Rock Mass Internal 
8.7 
42 Angle of Friction,  
(degrees) 
Rock Mass 
39.7 
Compressive Strength, 
cm 
13.7 
(MPa) 
Rock Mass Modulus, Em 
(GPa) 
Rock Mass Quality 
Parameter, Qp 
Rock Quality 
Category 1 
10% 
7.6 
40 
33.5 
10.3 
2.1 
50 Hoek-Brown Geologic 
Strength Index, GSI 
Source: BSC 2003a, Table 8-41. 
3.6 
55 
4-32 No. 4: Mechanical Degradation 
Revision 1 
Rock Quality 
Category 5 
90% 
11.8 
47 
57.7 
26.2 
12.6 
Rock Quality 
Category 4 
70% 
10.4 
45 
49.5 
20.2 
7.7 
62 
Rock Quality 
Category 3 
50% 
9.5 
44 
44.4 
16.7 
5.3 
59 67 
June 2004 Revision 1 
5.1 INTRODUCTION 
As discussed in Section 1, there are four primary mechanisms for drift degradation in the 
preclosure and postclosure time periods. These include rock mass yield and potential instability 
resulting from: (1) stresses induced by in situ gravitational loading, (2) stresses induced by waste 
package heat generation, (3) stresses and shaking induced by seismic ground motion, and 
(4) time-dependent strength loss from stress corrosion mechanisms. Drift degradation is defined 
here as the physical impact of the potential mechanical yield of the rock mass surrounding the 
emplacement drifts to application of these stresses or time-dependent strength change. A series 
of numerical calculations are discussed here that are aimed at predicting the drift degradation in 
physical terms that are easily understood: 
 The extent and character of the fracturing around the emplacement drift tunnels as a 
function of time and stressing 
 The character of rockfall that may occur as a result of this fracturing and the associated 
gravitational and seismic accelerations, including the size distribution of particles, the 
mass of the particles, and their velocities 
 The evolution of the shape of the emplacement drifts and the dislodged rock as functions 
of loading and time. 
To provide these results, it is necessary to use numerical modeling techniques that allow 
representation of preexisting rock structure and that represent fracturing explicitly, as described 
in Section 4. Thus, reliance is placed on the use of discontinuum numerical methods that allow 
displacement on preexisting rock mass structure and rock fracturing in a realistic fashion. 
The following sections give a discussion of the modeling of emplacement drift stability and 
rockfall subject to postclosure loading. In addition, a description of the analysis of 
time-dependent drift stability is given. Due to the different modeling techniques required for 
nonlithophysal and lithophysal rock masses, the results are described separately. 
5. ANALYSIS OF PRECLOSURE AND POSTCLOSURE DRIFT DEGRADATION 
UNDER GRAVITATIONAL, THERMAL, AND SEISMIC LOADING 
June 2004 
5.2 PRECLOSURE GROUND SUPPORT ISSUES 
Mechanical degradation analysis of the unsupported emplacement drift openings for an assumed 
100 years of normal operation (Williams 2003) has shown that the drifts will remain intact and 
not collapse without engineered ground support (BSC 2003c). Those calculations consider static 
gravitational and thermal loads. The emplacement drift ground support, consisting of perforated 
stainless steel sheets (2 mm to 3 mm thick) and stainless steel friction rock bolts, is provided 
primarily for worker safety and to ensure the retrieval option. 
The impact of in situ stress and preclosure rock temperatures on stability of unsupported tunnels 
was examined, as well as seismic events of 5  10-4 and 1  10-4 per year frequency of 
occurrence (or return periods of 2,000 and 10,000 years, respectively). Section 5.3 provides 
details of these calculations. Rockfall induced by preclosure seismic events was also estimated 
5-1 No. 4: Mechanical Degradation Revision 1 
to support the evaluation of the safety class of the ground support. These analyses were used to 
identify the largest potential rock size that could fall on a waste package (assuming no ground 
support). The results of these calculations were used to confirm that the waste package design 
was sufficient to preclude a breach of the waste package for the resulting rockfalls, and, thus, the 
ground support was classified as not important to safety (BSC 2004k). 
Although this document primarily addresses postclosure performance issues, a number of the 
appendices discuss the ground support analyses in detail. Table 1-1 provides a list of NRC KTI 
agreements and the sections or appendices that discuss the associated issues. Most of these 
agreements deal with rock properties and modeling methods that apply to both the preclosure and 
postclosure drift degradation and seismic analyses. 
5.3 ANALYSIS OF PRECLOSURE AND POSTCLOSURE DRIFT DEGRADATION 
5.3.1 Emplacement Drift Degradation from In Situ and Thermal Stressing Effects 
Thermal-mechanical modeling was performed to define drift stability under combined in situ and 
thermally induced stresses. Temperatures within the rock mass are determined from thermal 
analysis conducted using the NUFT program (LLNL 2002), which accounts for the details of 
heat transfer mechanisms within the drift, including heat removal due to ventilation in the 
preclosure period. The NUFT approach is two-dimensional, and thus assumes a cross section 
through a series of infinitely long emplacement drifts. Thus, this type of approach adequately 
represents the developing temperature distribution around emplacement drifts located centrally 
within the repository. 
To examine three-dimensional geometric effects (which include both repository edge effects and 
topographic influences) on the temperature distribution, the FLAC3D (BSC 2002e) regional and 
local-scale models are used (Figure 5-1). Here, the heat flux into the rock mass at each tunnel, 
obtained from the NUFT program, is distributed over the plane of the repository as a function of 
time. The resulting temperature distributions are then used in drift-scale two- and threedimensional 
models to determine the thermally induced stress state as a function of time. The 
details of the thermal-mechanical calculations are described in detail in Drift Degradation 
Analysis (BSC 2004a, Section 6.2 and Appendix C). 
Three major cases of the drift-scale thermal calculation were carried out: 
 Case 1Base-case calculation with 1.45 kW/m initial heat load and 50-year preclosure 
ventilation (90% heat removal ratio, BSC 2004l, Section 5.1). Preclosure maximum 
may be 100 years, but base case is 50 years of ventilation. 
 Case 2Sensitivity calculation for thermal properties of repository rock material (Tptpll) 
with 1.45 kW/m initial heat load, 50-year preclosure ventilation, and 90% heat removal 
ratio. Values of thermal conductivity and specific heat one standard deviation less than 
the mean values were used: 
June 2004 5-2 No. 4: Mechanical Degradation Revision 1 
- Thermal Conductivity (DTN: SN0208T0503102.007): 1.64 W/mK (= 1.89 W/mK  
one standard deviation (0.25 W/mK)) for wet conditions and 1.03 W/mK (= 1.28 
W/mK  one standard deviation (0.25 W/mK)) for dry conditions). 
- Heat Capacity: 811 J/kgK (= 954 J/kgK  one standard deviation (143 J/kgK)). 
 Case 3Sensitivity calculation for heat removal ratio. A heat removal ratio of 70% was 
used for the preclosure ventilation. This level is significantly below the estimated heat 
removal ratio during ventilation and examined only for the purpose of investigating 
preclosure ventilation rates. Figure 5-2 shows the linear heat load into the rock mass for 
each of these cases. 
June 2004 5-3 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure C-1. 
NOTE: (a) View is toward south. (b) Repository panel outline is superimposed. 
Figure 5-1. Aerial View of the Yucca Mountain Site and Digital Elevation Calculation Created from 
Topographic Information (a) and View of the Regional-Scale FLAC3D Thermal-Mechanical 
Model Constructed from the Digital Elevation Calculation and Available Geologic 
Information (b) 
June 2004 5-4 No. 4: Mechanical Degradation Revision 1 
Temperature histories at the drift crown for all the cases of the thermal calculations are presented 
in Figure 5-3. The results exhibited the temperature increase from base case (Case 1) to 
sensitivity calculations (Cases 2 and 3). In particular, Case 3 showed a significant temperature 
increase at the preclosure period. The peak temperature for Case 1 was 138C at around 
75 years, while Cases 2 and 3 were 161C and 153C at around 75 years, respectively. 
Source: BSC 2004a, Figure 6-31. 
NOTE: The no ventilation curve is from BSC 2003g. Cases 1 and 2 use the 90% heat removal curve, while Case 3 
uses the 70% heat removal curve. 
June 2004 
Figure 5-2. Heat Decay Curves for Thermal Calculations 
5-5 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 6-32. 
NOTE: This graphic assumes a 1.45 kW/m linear heat load and 50 years of preclosure ventilation. 
Figure 5-3. Temperature History at the Drift Crown Due to the Linear Heat Load Presented in Figure 5-2 
A comparison of temperature histories in the drift crown for Case 1, as determined in the 
drift-scale calculation (NUFT) and the coupled regional- and local-scale calculations (FLAC3D), 
is quite good, with the conditions in the middle of the repository representing the most 
conservative conditions. Details of these calculations are given in Drift Degradation Analysis 
(BSC 2004a). 
A series of thermal-mechanical calculations are performed initially for the range of lithophysal 
and nonlithophysal mechanical rock properties (Table 4-2) to examine the potential for yield and 
degradation due to in situ stress and thermal loading alone.14 The drift-scale UDEC model 
described in Section 4.2.7 is allowed to equilibrate under initial stress conditions, followed by 
application of the rock mass temperature history derived from the NUFT program. The 
nonlithophysal rock mass shows no thermally induced yield because the stresses are significantly 
below the yield condition of the rock mass (BSC 2004a). An example of the model output from 
the lithophysal rock, Case 1 thermal history is shown in Figure 5-4. Here, the temperature 
distribution and major and minor principal stress trajectories are plotted at 80 years (peak 
temperature and thermally induced stress for case of 50-year forced ventilation and closure) and 
10,000 years near the completion of cool-down. 
The in situ gravitational and thermally induced principal stress history paths for a series of 
radially oriented points, starting at the drift crown and springline and continuing into the rock 
mass, are given in Figures 5-5, 5-6, and 5-7 for the lithophysal rock mass. These figures 
represent the results of a number of thermal-mechanical simulations in which the rock mass 
temperatures given by Case 1 above are applied to an emplacement drift excavated and 
14 Impact of time-related strength degradation mechanisms are examined in Section 5.3.3.2.4. 
June 2004 5-6 No. 4: Mechanical Degradation Revision 1 
equilibrated at the in situ stress condition. The path that the stress state then takes as a function 
of time as the rock mass is heated and cooled down over a 10,000-year time period are then 
given. The rock mass is assumed to have the moduli representative of the lowest quality 
lithophysal rock (Category 1, Figure 5-5), the highest quality lithophysal rock (Category 5, 
Figure 5-6) and the base case quality lithophysal rock (Category 3, Figure 5-7). 
The calculations assume an elastic rock mass, but the associated failure envelope is 
superimposed on the principal stress history paths to allow estimation of the extent of failure into 
the rock mass as a function of time. As seen in these figures, the lowest quality (highest 
porosity-approximately 3% of the Tptpll) lithophysal rock indicates a small zone of yield of 
approximately 0.25 m directly at the springline of the tunnels from in situ stress. The depth of 
this yield zone changes only slightly from the thermal stresses and does not impact the overall 
stability of the tunnels. This small yield zone is the same as that discussed in Section 4.2.7.2 
regarding calibration of the UDEC modeling approach. 
Simulations for the lower porosity lithophysal rock (Category 5approximately 30% of the 
Tptpll) indicate minor yielding of less than 0.25 m in the crown, whereas the base-case quality 
lithophysal rock (approximately 35% of the Tptpll) shows no failure. The conclusions from 
these analyses show that no significant yield or ground collapse mode is evident for 
emplacement drifts in either lithophysal or nonlithophysal rocks during the preclosure or 
postclosure time frames due to in situ or thermally induced stressing alone (i.e., in the absence of 
time-dependent strength change). 
June 2004 5-7 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 6-142. 
NOTE: Peak stress occurs at springline and small amount (0.25 m or less; seen as white zone with absence of 
stress vectors at wall) of yield is predicted from initial stress state. Depth of yield does not increase from 
heating alone. 
Figure 5-4. Comparison of Temperatures, Major and Minor Principal Stress Trajectories at 80 (Peak 
Temperature) and 10,000 Years for Lithophysal Rock, Lowest Quality, Category 1 
June 2004 5-8 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figures 6-144 and 6-143. 
NOTE: Only minor yield (approximately 0.25 m depth) occurs in drift springline area for this lowest strength 
category. This yield is predicted to occur prior to initiation of heating. Estimated failure envelope is shown 
for comparison to the stress conditions. Initial point is at preheating stress state, followed by path through 
10,000 years of heat-up and cool-down. 
Figure 5-5. Elastic Principal Stress Path Histories for Points at Increasing Depth from Emplacement 
Drift Crown (a) and Springline (b) for Lithophysal Rock, Modulus from Lowest Quality 
Category 1 
June 2004 5-9 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figures 6-146 and 6-145. 
NOTE: Only minor yield (less than about 0.25 m depth) occurs in drift crown area for this highest lithophysal 
strength category. This yield is predicted to occur after 80 years of heating. Larger stress change occurs in 
crown for higher modulus category as compared to Figure 5-5. Estimated failure envelope is shown for 
comparison to the stress conditions. Initial point is at preheating stress state, followed by path through 
10,000 years of heat-up and cool-down. 
Figure 5-6. Elastic Principal Stress Path Histories for Points at Increasing Depth from Emplacement 
Drift Crown (a) and Springline (b) for Lithophysal Rock, Modulus from Highest Quality 
Category 5 
June 2004 5-10 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figures 6-84 and 6-85. 
NOTE: No yield from in situ or thermally induced stresses occurs. Estimated failure envelope is shown for 
comparison to the stress conditions. Initial point is at preheating stress state, followed by path through 
10,000 years of heat-up and cool-down. 
Figure 5-7. Elastic Principal Stress Path Histories for Points at Increasing Depth from Emplacement 
Drift Crown (a) and Springline (b) for Nonlithophysal Rock, Strength Category 3 
June 2004 5-11 No. 4: Mechanical Degradation Revision 1 
5.3.2 Analysis of Drift Degradation from Combined In Situ, Thermal, and Seismic 
Loading 
5.3.2.1 Nonlithophysal Rocks 
5.3.2.1.1 Introduction 
The approach to modeling rock mass degradation and rockfall from combined in situ, thermal, 
and seismic loading in nonlithophysal rocks is illustrated in Figure 5-8. This approach involves 
modeling of emplacement drifts excavated within the stochastically defined, representative 
fractured rock mass volume, followed by application of in situ, thermal, and seismic load. 
A large number of parameter studies are conducted in which the tunnel location (and fracture 
geometry), rock fracture surface properties, and loading conditions are varied to derive a 
conservative range of performance response (in terms of rockfall mass, volume, and opening 
shape) that is indicative of the possible geologic conditions at depth. This data is then fed to 
analysis of the postclosure structural stability and determination of location and areal extent of 
localized yielding of the drip shield. This localized yielding is further abstracted into a damage 
model for of the titanium (e.g., tearing or stress corrosion cracking) with associated estimates of 
the drainage of seepage waters through the drip shield (BSC 2004e). The seismic consequence 
abstraction provides the final abstraction of damage consequences for drip shield and waste 
package to the total system performance assessment model. 
Figure 5-8. Approach to Analysis of Drift Degradation and Rockfall from Combined In Situ, Thermal, 
and Seismic Loading 
June 2004 5-12 No. 4: Mechanical Degradation Revision 1 
3DEC Model Development 
As described in Section 4.1, the FracMan program is used to develop the basic fracture input data 
for the 3DEC program. A 100-by-100-by-100-m representative fractured rock mass volume is 
generated and composed of the matrix blocks defined by approximately 90,000 fractures. A total 
of 50 emplacement drifts are randomly located and excavated within this rock mass such that 
the stochastic nature of the jointed medium and its impact on rockfall is adequately sampled. 
A random emplacement drift centroid coordinate is chosen within the cube, and a 
25-by-25-by-25-m volume, oriented at the emplacement drift 72 azimuth, is extracted to contain 
the model emplacement drift (Figure 5-9). Within each emplacement drift, a rigid, rectangular 
block representing the drip shield is affixed to the invert of the tunnel. This drip shield block is 
placed only to facilitate the recording of block impacts (location and time of impact, block mass, 
velocity, and block shape) to the drip shield. For each rock that impacts the drip shield, and is 
not meant to represent its mechanical response, that modeling is accomplished in a separate 
calculation (BSC 2004m). 
An algorithm was developed for applying the FracMan fracture geometry to the 3DEC model. 
Previous versions of the 3DEC program were set up to only efficiently handle through-going 
joint planes. The new algorithm allows incomplete fractures to be cut within a block, or to 
terminate against other fractures, thus creating realistic fracture patterns within the rock mass.15 
The resulting blocks within the 3DEC model are fully deformable as they are subdivided into 
tetrahedral finite difference zones, whose constitutive behavior may be elastic or conform to a 
desired constitutive law. Here, the blocks were assumed to be elastic due to the high intact 
strength of nonlithophysal rock, although later this assumption is reviewed in detail. The 3DEC 
model uses a fully dynamic solution algorithm to solve the laws of motions for the blocks, 
subject to contact restraints with surrounding blocks. The gridpoints that lie along the fracture 
surfaces act as contact points across which forces are transmitted, subject to shear and tensile 
yield conditions. 
The circular disc-shaped fractures prescribed by the FracMan output are initially cut to completely transect the 
blocks in which they occur. In a second step, the 3DEC model bonds those portions of the fracture that fall 
outside the fracture radius with the strength of the intact rock, thus simulating a partially-fractured block. During 
a simulation, the bonded portion of the fracture is free to fail in shear or tension if the stresses dictate, thus 
creating multiple blocks. 
5.3.2.1.2 
15 
June 2004 5-13 No. 4: Mechanical Degradation Source: BSC 2004a, Figure 6-41. 
Figure 5-9. 3DEC Model Geometry and Cross Sections 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 5-14 Revision 1 
5.3.2.1.3 Ground Motion 
Site-specific ground motions were developed for the Yucca Mountain site through use of a 
formal process of expert elicitation resulting in development of a probabilistic seismic hazard 
assessment (PSHA). A summary of the PSHA can be found in Characterize Framework for 
Seismicity and Structural Deformation at Yucca Mountain, Nevada (BSC 2004n). Site-specific 
ground motion time histories for four levels of annual probability of exceedance, 10-4, 10-5, 10-6, 
and 10-7, were examined in Drift Degradation Analysis (BSC 2004a), with results described 
here. The development of the time histories are described in Development of Earthquake 
Ground Motion Input for Preclosure Seismic Design and Postclosure Performance Assessment 
of a Geologic Repository at Yucca Mountain, NV (BSC 2004c). The 10-4 ground motion level is 
for preclosure consideration, while the 10-5, 10-6, and 10-7 ground motion levels are for 
postclosure conditions. The 10-4 preclosure ground motion levels are also used in the analysis of 
access and emplacement drift seismic response for preclosure safety assessment and for analysis 
of ground support safety classification (BSC 2004k). 
For higher frequency spectral accelerations (5 to 10 Hz) and an annual exceedance probability of 
10-4, results of the probabilistic seismic hazard analysis for Yucca Mountain indicate the ground 
motion hazard derives primarily from earthquakes in the magnitude range of 5.0 to 6.5 occurring 
at distances less than 15 km from the site. For lower-frequency spectral accelerations (1 to 2 Hz) 
at the same annual exceedance probability, the hazard shows, in addition to nearby sources, a 
significant contribution from earthquakes in the magnitude range of 7.0 to 8.0 occurring at an 
epicentral distance of about 50 km. For annual exceedance probabilities of 10-6 and 10-7, nearby 
earthquakes in the magnitude range 5.5 to 7.0 are the dominant sources contributing to ground 
motion hazard at both higher and lower spectral accelerations. Discussion of the derivation of 
the site specific ground motion time histories can be found in Technical Basis Document No. 14: 
Low Probability Seismic Events. 
A total of 15 sets of ground motion time histories were developed at the repository horizon for 
each annual postclosure hazard level. The multiple sets ensure a reasonable distribution of 
spectral shapes and time history duration. For each set of ground motion, two horizontal 
components (H1 and H2) and one vertical component (V) of acceleration, velocity, and 
displacement are supplied. Figure 5-10 shows the H1 velocity time history for four annual 
hazard levels. Only one ground motion time history set was provided for the preclosure hazard 
levels because of the deterministic-based approach for preclosure consideration. 
The amplitude of the peak ground acceleration, velocity, and displacement, and the 
seismic-induced far-field stress for one of the ground motion sets from each hazard level are 
provided in Table 5-1. This table is used to demonstrate the typical ground motion parameters 
for the three hazard levels considered. 
It is apparent that the preclosure ground motion levels have lower amplitude vibrations and 
hence lower induced rock mass stresses compared to the postclosure ground motion levels. The 
peak values for each postclosure hazard level vary depending on the hazard level. The complete 
data sets of the ground motion are contained in Drift Degradation Analysis (BSC 2004a). 
June 2004 5-15 No. 4: Mechanical Degradation Revision 1 
Source: DTNs: MO0211TMHIS104.002, MO0306SDSAVDTH.000, MO0301TMHIS106.001, 
MO0403AVTMH107.003; MO0402AVDTM105.001; BSC 2004a. 
Figure 5-10. Examples of Ground Velocity Time Histories (H1) with Truncated Duration for Analysis 
June 2004 5-16 No. 4: Mechanical Degradation Revision 1 
Table 5-1. Peak Ground-Motion Parameters 
Peak 
Displacement 
(cm) 
44.44 
Peak 
Velocity 
(cm/s) 
38.38 
Peak 
Acceleration 
(g) 
0.39 
Ground 
Motion 
Component 
H1 
H2 45.3 43.78 0.37 
Annual Hazard 
Level 
10-4 31.73 47.51 0.47 V 
H1 20.06 104.58 2.77 
H2 14.37 83.31 2.50 
13.00 70.88 2.63 V 
10-5 Ground 
Motion Set 1 
H1 26.76 244.14 7.42 
H2 26.78 195.41 6.74 
13.75 111.29 4.90 V 
10-6 Ground 
Motion Set 1 
H1 58.68 535.26 16.28 
H2 58.72 428.42 14.79 
Seismic Induced Stress 
Corresponding to Peak 
Velocity (Pa)* 
2.20  106 
2.51  106 
4.50  106 
6.00  106 
4.78  106 
6.71  106 
1.40  107 
1.12  107 
1.05  107 
3.07  107 
2.46  107 
2.83  107 36.86 298.44 13.15 V 
10-7 Ground 
Motion Set 1 
Source: DTNs: MO0211TMHIS104.002, MO0306SDSAVDTH.000, MO0301TMHIS106.001, 
MO0403AVTMH107.003, MO0402AVDTM105.001. 
5.3.2.1.4 
NOTE: Seismic-induced stress is calculated based on elastic wave equations (Itasca Consulting Group 2002, 
Manuals/3DEC/Optional Features/Section 2: Dynamic Analysis, Section 2.5). 
Ground motion at the location of the repository, in the form of three mutually perpendicular 
velocities (two horizontal and one vertical), are applied to the lower surface of the model, in 
terms of equivalent stress time histories. The motions, as supplied to the drift degradation 
modeling, include the effects of the free surface reflections, and, thus, the 3DEC model does not 
need to account for topography. 
Nonreflecting vertical and upper model boundaries in 3DEC allow the wave to pass through the 
model, and free-field boundaries on the vertical sidewalls of the model prevent damping and 
distortion along the vertical sidewalls of the incoming wave. No material damping,16 in addition 
to that supplied by sliding on fracture surfaces, is supplied to the model. Prior to use of the 
model for examination of drift degradation, seismic wave propagation of models without tunnels 
was run to ascertain that the wave passed through the model without significant distortion. 
June 2004 
Combinations of Ground Motion and Fracture Modeling Region 
A goal of these analyses is to provide an estimate of seismically induced rockfall that is derived 
from an adequate sampling of the variability of fracture geometries and ground motion time 
histories. As described previously, a 100-m cube was constructed from FracMan for providing 
the fracture network required in 3DEC analysis. A random selection of 105 emplacement drift 
centroid locations was conducted. These 105 centroid locations combined with the 15 sets of 
16 0.3% of critical damping was used in a few analyses for numerical stability purposes. 
5-17 No. 4: Mechanical Degradation Revision 1 
ground motion data serve as the basis for sampling for numerical analysis. A simple Latin 
Hypercube sampling scheme was used for the pairing of ground motion and fracture modeling 
region (DTN: MO0301SPASIP27.004). A total of 50 sets of paired fracturing realizations 
(i.e., drift centroid locations) and ground motion data17 were made for each postclosure annual 
exceedance frequency. For each of these analyses, a base case of block and fracture material 
properties were used so that the variability of the rockfall response was then a function of the 
fracture geometry and ground motion variability only. The base case rock and fracture 
properties are given in Table 5-2. 
5  104 
Intact rock deformation 
properties 
Intact bridge strength 
properties 
Table 5-2. Base-Case Material Properties for 3DEC Analysis 
Joint strength properties 
0.1 
41 
0 
Joint cohesion (MPa) 
Joint friction () 
Joint dilation () 
Joint normal stiffness, Kn (MPa/m) 
Joint shear stiffness, Ks (MPa/m) 
Youngs modulus (GPa) 
Poissons ratio 
Bulk modulus (GPa) 
Shear modulus (GPa) 
Cohesion (MPa) 
Friction angle () 
Tensile strength (MPa) 
5  104 
33.03 
0.21 
19.2 
13.6 
47.2 
42 
11.56 
NOTE: Values of cohesion and friction angle were derived from preliminary data with a 
slight deviation from the reported values in Drift Degradation Analysis (BSC 
2004a, Section E3). An impact analysis was conducted with no difference in 
the results for rockfall prediction (BSC 2004a, Appendix Q). Joint dilation (BSC 
2004a, Table E-3) is set to zero for the base-case analysis. With no dilation, 
joints are modeled as perfectly planar and smooth, resulting in a conservative 
(i.e., higher) estimation of rockfall. 
5.3.2.1.5 Example of Seismic Analysis for Case of 10-6 Annual Exceedance Probability 
The results for a complete set of 3DEC analyses, subjected to the postclosure hazard level of 
10-6 annual probability of exceedance ground motion, are presented in this section. The analyses 
reviewed here are for nonthermal loading conditions; however, the impact of thermal load, in 
addition to gravitational and seismic stresses, is discussed later. A summary of the 10-4 
simulations at other annual exceedance frequencies can be found in Drift Degradation Analysis 
(BSC 2004a). 
17 The adequacy of 50 analyses for representation of the variability of rockfall (at each exceedance level) was 
verified by calculation of the cumulative mean and standard deviation of rockfall parameters for successive 
analyses. As described in Drift Degradation Analysis (BSC 2003a), the mean and standard deviation of rockfall 
mass, for example, does not change after approximately 30 runs. 
June 2004 5-18 No. 4: Mechanical Degradation Revision 1 
The block representing the drip shield is anchored at the invert, and is included in the model to 
record the information of the locations and relative velocities for the rockfall impact. 
Figure 5-11 shows a typical block impacting the drip shield in the 3DEC dynamic simulation. 
Note that fallen blocks are automatically deleted after impacting the drip shield. The deletion is 
to facilitate a conservative approach of recording of all possible rockfall on the drip shield. If the 
blocks are not deleted for the heaviest of rockfall cases, the drip shield will be covered with 
fallen rocks so that some of the rockfall at the later stage of seismic shaking will not impact the 
drip shield. The simulation without deletion of the rock blocks after the impact was examined to 
define the conservatism in this present approach, indicating less rockfall impact without this 
deletion scheme. 
The results of the 50 3DEC simulations are summarized in Table 5-3. All of the simulations 
predicted some level of rockfall associated with the 10-6 seismic shaking, resulting in an average 
of 398 m3 of rockfall per kilometer of drift length. The associated impact parameters for these 
blocks from the analyses include the following: 
 Rock block volume falling on the drip shields (unit in cubic meters) 
 Relative impact velocity of rock block to the drip shields (unit in meters per second) 
 Impact location. 
Value 
50 
2,797 
497.7 
1,250 
2,238 
398.2 
Table 5-3. Summary of 3DEC Rockfall Prediction for 10-6 Annual Probability of Exceedance Hazard 
June 2004 
Parameter 
Source: BSC 2004a, Table 6-13. 
Simulations Completed 
Total Number of Rockfall 
Total Volume of Rockfall (m3) 
Total Length of Drift Simulated (m) 
Number of Blocks per km 
Volume of Rockfall per km (m3) 
5-19 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 6-56. 
NOTE: 3DEC Simulation 22, 10-6 ground motion 5, at t = 5.24 s. 
observed. 
Figure 5-11. Simulation of Rockfall Impact to the Drip Shield 
The distribution of the data for each parameter (i.e., block mass, relative impact velocity, impact 
angle, impact momentum, and impact energy) is presented using histograms (Figures 5-12 
to 5-16). Also included in each histogram plot is the cumulative frequency of occurrence. 
Due to the gravity effect, most of the rockfall will occur in the range of 48 to 132 (based on the 
drip shield coordinate system) as confirmed in Figure 5-14. The impact momentum and impact 
energy, both functions of block mass and impact velocity, were calculated as the required outputs 
for drip shield structural response calculation. Summary statistics for these parameters are 
provided in Table 5-4. The maximum rockfall block mass predicted is 28 MT with median block 
size of 0.13 MT. The predicted results (Table 5-4) show large variance and high skewness with 
the exception of impact velocity, as confirmed by the shape of the histograms. The block mass, 
impact angle, impact momentum and impact energy show the trend of exponential distribution 
with most of the data concentrated on the low end of the data range. The impact velocity shows 
a typical bell shape for the normal distribution. The distribution centers around 3 m/s with a 
standard deviation of approximately 1.7 m/s (BSC 2004a). 
The relative low impact velocities indicate that block fallout is mainly due to gravitational 
free-fall. Differential acceleration or energy trapping to induce high ejection velocity is not 
5-20 June 2004 No. 4: Mechanical Degradation Revision 1 
Table 5-4. Statistical Summary of the Rockfall Impact Parameters, 10-6 Annual Probability of 
Exceedance Hazard, Variability in Fracture Geometry and Ground Motion 
Mean 
Median 
Skewness 
Range 
Minimum 
Maximum 
Impact Momentum 
(kgm/s) 
1,217 
337 
3,464 
11 
79,001 
2 
79,003 
3,403,555 
Impact Energy 
(Joules) 
2,350 
576 
7,704 
12 
163,657 
0 
163,657 
6,573.633 
Relative Impact 
Velocity (m/s) 
3.23 
2.97 
1.74 
1.06 
12.03 
0.07 
12.10 
N/A Sum 
Impact Angle 
() 
136 
124 
93 
0.87 
359 
0 
360 
N/A 
Figure 5-12. Histogram for Block Mass (10-6 Annual Probability of Exceedance Hazard) 
Block Mass 
(MT) 
0.43 
0.13 
Standard Deviation 1.30 
11.61 
28.19 
0.02 
28.22 
1200.43 
Source: BSC 2004a, Table 6-14. 
Source: BSC 2004a, Figure 6-59. 
June 2004 5-21 No. 4: Mechanical Degradation Source: BSC 2004a, Figure 6-60. 
Figure 5-13. Histogram for Relative Impact Velocity (10-6 Annual Probability of Exceedance Hazard) 
Source: BSC 2004a, Figure 6-61. 
Figure 5-14. Histogram for Impact Angle (10-6 Annual Probability of Exceedance Hazard) 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 5-22 Source: BSC 2004a, Figure 6-62. 
Figure 5-15. Histogram for Impact Momentum (10-6 Annual Probability of Exceedance Hazard) 
Source: BSC 2004a, Figure 6-63. 
Figure 5-16. Histogram for Impact Energy (10-6 Annual Probability of Exceedance Hazard) 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 5-23 Revision 1 
5.3.2.1.6 
428 
39.4 Total Volume of Rockfall (m3) 
800 Total Length of Drift Simulated (m) 
535 Number of Blocks per km 
Volume of Rockfall per km (m3/km) 
Source: BSC 2004a, Table 6-22. 
Summary of Nonlithophysal Rockfall from Preclosure and Postclosure 
Ground Motion 
Analyses similar to that discussed in the previous section were also completed for 10-4, 10-5, and 
10-7 motions. The exception is that only one ground motion set was provided for the preclosure 
cases, and 32, rather than 50, analyses were made. 
Figure 5-17 shows a comparison of histograms of block mass for all three ground motion results. 
Essentially, the results show that all motions result in the same general distribution of block sizes 
with mean block masses of less than 0.2 MT, and an exponentially decaying distribution. 
A comparison of the rockfall statistics for the preclosure and postclosure events is given in 
Table 5-5. The important comparison statistic here (due to the variable number of runs) is the 
number of blocks per km, which shows increase from 535 to 2,840 blocks in going from the 10-4 
to 10-7 annual exceedance frequency. 
Statistic 
32 
Total Number of Rockfall 
Table 5-5. Comparison of Rockfall Statistics for Preclosure and Postclosure Events 
10-7 10-4 
44 50 Runs Completed 
3,219 
678.3 
1,100 
2,926 
616.6 
Ground Motion 
10-6 10-5 
50 
2,797 1,764 
497.7 255.4 
1,250 1,250 
2,238 1,414 
398.2 204.3 
June 2004 
49.3 
5-24 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 6-87.1. 
5.3.2.1.7 Summary of Drift Profile 
The three-dimensional depiction of an emplacement drift after 10-6 ground motion analyses 
provides a physical feel for the impact of seismically induced rockfall on drift profile. 
Figures 5-18 and 5-19 show the rockfall profile for the case showing the greatest amount of 
rockfall and for the case showing 50th percentile of quantity of rockfall. 
Two particular cases of the 50 analyzed show larger amount of rockfall due to the fact that the 
two long, subvertical fracture sets strike at a low angle to the axis of the emplacement drift. 
This, coupled with the presence of closely spaced subhorizontal vapor phase partings or random 
sets, allows a number of roof and sidewall blocks to detach over a significant plan view area. As 
seen in the 50th percentile case, isolated wedges formed, again, by two subvertical and 
subhorizontal sets can occasionally occur. 
Figure 5-17. Comparison of Histograms of Block Mass from All Postclosure and Preclosure Ground 
Motion 
June 2004 5-25 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 6-114. 
NOTE: Gray tube is emplacement drift as represented in the 3DEC model. This realization shows a case that has a 
1 in 50 chance of occurrence. 
Figure 5-18. Drift Profile Showing Blocks That Become Detached from Drift Wall during Simulation for 
10-6 Hazard Level, Case with Greatest Amount of Rockfall of 50 Total Realizations 
June 2004 5-26 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 6-115. 
NOTE: This realization shows the mean case of 50 realizations. 
Figure 5-19. Drift Profile Showing Blocks that Become Detached from Drift Wall During Simulation for 
10-6 Hazard Level, 50th Percentile Case 
5.3.2.1.8 Summary of Sensitivity Studies of the Input Parameters and Model 
Conditions 
There are four major input data sets to the three-dimensional discontinuum analysis: ground 
motion, joint geometrical properties, joint and intact mechanical properties, and thermal stress 
history. Sensitivity studies of these input parameters were conducted to establish uncertainty in 
the predictions of rockfall and to identify the important controlling parameters. The following 
section provides a summary of the conclusions of these parameter studies. Details are presented 
in Drift Degradation Analysis (BSC 2004a). 
Ground Motion SensitivityA total of 15 sets of ground motion data were used for each hazard 
level in the postclosure consideration to ensure a reasonable distribution of spectral shapes and 
time history durations. Sensitivity studies on the peak ground velocity, energy contents, 
duration, and orientation of horizontal motions in the 3DEC model to the rockfall prediction 
were examined and reviewed in the previous discussion. Correlation of the rockfall volume with 
the peak ground velocity and energy content in term of Arias Intensity shows that no strong 
relationship is observed for either of the parameters. The ground motion time histories were 
truncated at 5% and 95% energy content to shorten the time required to conduct the dynamic 
analyses. The analyses showed that the majority of the rockfall occurs coincident with the 
June 2004 5-27 No. 4: Mechanical Degradation arrival of the strong motion, which is typically within the first 15 seconds of shaking, and that 
truncating the ground motion had minimal impact on the amount of rockfall. However, the 
results were inspected at the end of the simulation, and, if it was determined that the simulation 
was terminated prematurely (i.e., there was indication of loose blocks), the simulation was 
continued until all loose blocks resulted in rockfall. Likewise, it was found that rotating the 
horizontal ground motion components by 90 had minimal impact on rockfall. This is 
understandable since the horizontal motions have, in general, similar peak amplitudes. 
Joint Geometrical PropertiesThe variability of joint geometrical properties is incorporated in 
the application of FracMan to generate a 100-m cube fracture network. A total of 50 drift 
locations were selected from the 100-m cube fractured rock mass for the 3DEC analyses. 
Results from the analyses of 50 drift locations (or fracture modeling regions) reasonably explore 
the impact of the variability of joint geometrical properties (BSC 2004a, Appendix K). 
Angle 
Joint Joint 
Category Cohesion (Pa) 
1 1.0  105 
34.4 2 1.0  105 
Joint Dilation Peak Friction Joint Shear 
Angle 
1.4 31.4 
4.4 
11 
5.3  109 
1.1  1010 
1.7  1010 3 1.0  105 
Source: BSC 2004a, Table 6-24. 
Time Degradation of Joint PropertiesThe potential exists for time-dependent degradation of 
the rock mass surrounding the tunnels. In the nonlithophysal rock, the potential source of 
Fracture Surface Property Variation and Fracture Strength DegradationThe base-case 
joint properties, listed in Table 5-2, were based on the rotary shear tests of the cored specimen as 
derived in Drift Degradation Analysis (BSC 2004a, Attachment V). Additional direct shear tests 
(Section 3.2.4.1) have been completed, and results from these tests are used to provide the range 
of variation tested in the sensitivity studies. With limited joint test results currently available and 
given the fact that the use of rotary shear devices in rock mechanics is not common, some of the 
parameters in the base case, such as cohesion and dilation angle, were scaled down from the 
testing results for conservatism, to allow for investigation of impact on increased rockfall. 
A range of joint properties, as shown in Table 5-6, was selected for the sensitivity study. The 
values were established based on the residual friction angle of 30 and three tiers of dilation 
angles. The dilation angles were selected within the range of reported test results presented in 
Drift Degradation Analysis (BSC 2004a, Attachment V) and discussed in Section 3.2.4.2. 
Cohesion is conservatively set to be 0.1 MPa. 
The results of these sensitivity studies show that the variation of joint mechanical properties is a 
secondary effect compared with the variation of fracture geometrical properties (i.e., fracture 
pattern). Results for the three categories are quite similar, irrespective of the variation of the 
mechanical properties used for each category. 
Table 5-6. Three Categories of Joint Properties Used in the Sensitivity Study 
June 2004 
41 
5-28 No. 4: Mechanical Degradation 
Revision 1 
Joint Normal 
Stiffness (Pa/m) Stiffness (Pa/m) 
7.2  1010 
9.4  1010 
1.2  1011 Revision 1 
time-dependency is the result of long-term shear failure along the preexisting fracture planes. A 
potential mechanism for time-dependent yield along rock fractures is the concentration of stress 
on joint asperities (i.e., roughness along the joint surfaces) with associated static fatigue failure 
when subjected to long-term constant shear stress. Static fatigue of hard rocks is typically 
associated with stress levels on the order of 60% to 80% of its unconfined compressive strength 
(see Section 5.3.2.2.4 for greater detail on time-dependent properties of welded tuff). Fatigue 
failure would presumably initiate along asperities on fracture surfaces, with the ultimate effect of 
reducing the fracture surface roughness. From a mechanical perspective, this failure would 
result in reducing or eliminating cohesion and dilation on the joint surface, as well as reducing 
the friction angle to its residual value. 
The impact on drift stability due to the effect of rock joint degradation is assessed here based on 
a conservative estimate of the reduction of joint cohesion and friction angle. The reduced joint 
strength parameters are estimated to be in the range of the residual, postpeak shear displacement 
state with joint cohesion reduced to 0 and the joint friction angle reduced to 30. The reduced 
friction angle is a typical value for a smooth joint reported by Goodman (1980, p. 158) and is 
consistent with the direct shear test results described in Section 3.2.4.1 (DTN: 
GS030283114222.001). Dilation angle is also conservatively presumed to be 0, considering that 
the asperities on fracture surfaces had been sheared off. The net result of these conservative 
assumptions is the potential for greater rockfall. The degraded joint strength and dilational 
properties were applied in the three 10-5 seismic motion cases that represent the case with 
greatest amount of rockfall, the median case, and the case producing no rockfall. The predicted 
number of detached rock blocks and the total rockfall volume show only a slight increase in 
rockfall is predicted for the degraded state. Thus, potential time-related joint strength 
degradation has a minor impact on drift stability in nonlithophysal rock. 
Rock Bridge Strength Solid rock bridges between fractures were automatically generated as 
the extension of finite trace length fractures to form the distinct blocks in the 3DEC model. A 
range of bridge strength parameters, in terms of cohesion, friction angle, and tensile strength, 
was selected for the sensitivity study. This range of intact rock properties was derived from the 
results of triaxial testing of rock cores from the Tptpmn (see Section 3.2.1). A total of 
3 categories were included to cover the possible range of variation for bridge strength parameters 
and was subject to 10-5 motions. The base case joint strength parameters are used for Category 1 
to represent the extreme case where all the bridges are sheared off to become fractures. The 
mean values of the intact Tptpmn strength parameters are used for Category 2. The mean plus 
one standard deviation values determined from triaxial testing are assigned as the strength 
parameters for Category 3. This category represents the upper bound for the rock bridge 
strength. The results show that within the range of variation for the intact strength parameters 
(Categories 2 and 3), rock bridge strength parameters have insignificant impact on rockfall 
prediction. However, if all rock bridges are sheared off, as represented by Category 1, a 
significant increase of rockfall volume occurs with smaller rock block size. 
Intact Rock Block StrengthThe base case parametric analyses assumed that all blocks are 
elastic and, therefore, do not yield. Thus, all rockfall is due to slip and separation along 
preexisting fracture surfaces. Further dynamic analyses were conducted to examine the impact 
of the postclosure ground motion on intact rock spalling mechanisms. In these analyses, the rock 
block strength, estimated to be approximately 70 MPa from back-analysis of the roof spalling in 
June 2004 5-29 No. 4: Mechanical Degradation Revision 1 
the heated overdrive portion of the heated drift test (see Section 4.2.7.1; BSC 2004a), was used 
to limit the peak stress in the periphery of the excavations. The 10-7 motion and some of the 
10-6 motions are capable of inducing significant spalling of intact rock, which could result in 
rockfall in the drift with small-scale spalled rock. This is very similar to the calculated 
postclosure response for the lithophysal rock, as given in the next section. The peak ground 
velocities associated with these low-probability events is very conservative (BSC 2004c). 
Thermal Stress Effects on Seismically Induced RockfallThe initial seismic loading studies 
were conducted assuming in situ stress and seismic loading only. An analysis of the 10-5 motion 
was conducted in which the base case thermal condition (see Figure 5-3, Case 1) stress state was 
added to the in situ stress conditions prior to ground motion application. Instead of conducting 
all 50 runs, three example cases were chosen: the case with the greatest amount of rockfall, a 
case showing the median amount of rockfall, and one case that showed no rockfall. The thermal 
state with the highest level of rock temperature and thermally induced rock stress (80 years after 
waste loading) was chosen. As illustrated in Figures 5-5, 5-6, and 5-7, the initial state (before 
heating began) and the state after 80 years of heating coincide with extreme points on thermally 
induced stress paths at a number of locations around the emplacement drift. The state after 
80 years was felt to be of greatest interest because already-completed nonthermal analyses 
provide the other extreme thermal condition.18 
As shown in Table 5-7, the impact of thermal loading in nonlithophysal rock is to stabilize the 
rock mass and reduce rockfall. The reason for this effect is that the rock mass expansion on 
heating induces tangential compression around the excavations. This compression tends to 
provide increased normal stresses to fractures, thus increasing their shearing resistance as well as 
minimizing joint opening during extensional loading during the seismic event. Thus, the most 
conservative thermal state, from a rockfall standpoint, is actually when the rock is at or near 
ambient temperature. 
18 It was demonstrated in Drift Degradation Analysis (BSC 2004a, Section 6.2.1.3) that little damage occurs in the 
rock mass from in situ stress and heating only; therefore, there is little yield and permanent deformation on 
cooldown. 
June 2004 5-30 No. 4: Mechanical Degradation Revision 1 
Table 5-7. Impact of Thermal Loading on Rockfall for 10-5 Ground Motion, Nonlithophysal Rock 
Thermal Stress Addition Nonthermal Analysis 
Number of 
Blocks 
Dislodged 
Number of 
Blocks 
Dislodged Rockfall 
Volume (m3) 
13.59 
Time of 
Event (years) 
0 
(Nonthermal) 
Rockfall 
Volume 
(m3)* 
42.03 
Case 
Greatest 
Rockfall 
Case 
1.07 2.49 0 
(Nonthermal) 
Median 
Rockfall 
Case 
(Nonthermal) 
173 
14 
0 
(Thermal) 
56 
5 
2 
Time of 
Event 
(years) 
80 
80 
80 5.93 0.00 0 
(Nonthermal) 
No Rockfall 
Case 
Source: BSC 2004a, Table 6-19. 
NOTE: *Volume per 25-m modeled drift length. 
5.3.2.1.9 
5.3.2.2 Lithophysal Rocks 
5.3.2.2.1 
Summary of Data Output Feeds from Nonlithophysal Analyses to Drip Shield 
Analysis and Design 
The results of the nonlithophysal analyses provide primary input to the detailed structural 
analysis of the stability and damage assessment of the drip shield. In particular, the following 
rockfall block information for each analysis at each annual exceedance frequency is fed to the 
drip shield engineering design analysis: (1) mass and shape, (2) velocity components (relative to 
the drip shield), (3) impact location and energy, and (4) timing of impact. This information is 
used for assessment of stability (buckling) and damage (both breaching of the skin of the drip 
shield as well as yield) and forms the basis for the abstraction of damage as a function of peak 
ground velocity in Seismic Consequence Abstraction (BSC 2004e). Structural analysis of the 
effect of rockfall on the drip shield can be found in Drip Shield Structural Response to Rock Fall 
(BSC 2004m). 
June 2004 
Two-Dimensional Discontinuum Analysis of Lithophysal Rock Mass 
The UDEC model, whose development was described in Section 4, was used to perform the rock 
mass degradation analyses discussed in this section. In the UDEC model, the rock mass is 
represented as an assembly of polygonal, elastic blocks. The entire domain is discretized into 
blocks using Voronoi tessellations (Itasca Consulting Group 2002). The joints between blocks 
are considered to be linearly elastic-brittle. The elastic behavior of joints is controlled by normal 
and shear stiffness (joint stiffness is constant). The joint stiffnesses produce a rock mass 
deformation modulus that is calibrated to reproduce the strength envelope of that given by 
lithophysal rock quality categories 1 to 5. Joints can sustain finite tensile stress as prescribed by 
the rock mass tensile strength. The Coulomb slip condition governs the onset of slip as a 
function of joint cohesion and friction angle, which have been calibrated to reproduce the 
5-31 No. 4: Mechanical Degradation Revision 1 
lithophysal rock mass in categories 1 to 5. If a joint fails either in tension or shear, tensile 
strength, friction and cohesion are reset to residual values. 
This model allows for the formation of fractures between blocks, separation and instability 
(under action of gravity and seismic accelerations) of portions of the rock mass around a drift. 
No ground support was considered in the analyses. All cases of thermal and seismic loading 
considered in this section were also analyzed using a continuum, linearly elastic approximation 
by the finite difference code FLAC (Itasca Consulting Group 2002). The results of the 
continuum model were used as a reference for additional interpretation of the results from the 
more complex UDEC model. 
The geometry of the UDEC model is shown in Figure 5-20. As indicated, only the region around 
the drift where inelastic deformation is expected is discretized into Voronoi blocks. The rest of 
the model is composed of a few large, elastic blocks that simply transmit far field stresses and 
ground motion from the boundary to the interior region. 
Each of the individual UDEC models presented in this section considers the rock mass to be 
characterized by a single rock property category. For example, a UDEC Category 5 model will 
use Category 5 rock properties throughout the model region. In reality, the lithophysal rock 
mass can contain multiple rock property categories within a 5- to 10-m zone. The consideration 
of spatial variability of the rock mass porosity, is provided in Drift Degradation Analysis (BSC 
2004a, Appendix S, Section S.4). The impact of spatial variability of porosity and the ultimate 
conservatism associated with the homogeneous properties approach used as a base condition is 
described later. 
June 2004 5-32 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 6-118. 
Figure 5-20. Geometry and Initial Conditions of the UDEC Model 
5.3.2.2.2 
NOTE: Voronoi blocks represent near-field nonlinear behavior. Elastic blocks represent far-field behavior. 
June 2004 
Seismic Consideration in Lithophysal Units 
Drift stability was analyzed for two preclosure and two postclosure ground motion data sets 
(5  10-4, 10-4, 10-5, and 10-6 annual exceedance frequencies). One ground motion, (consisting 
of the three component motions) was supplied for each of the preclosure cases, whereas 15 sets 
of ground motion data were supplied for the postclosure case. For each of these ground motion 
data, the response for the range of rock mass properties categories was investigated. 
An in situ (before excavation) stress state, defined by 7 MPa vertical and 3.5 MPa horizontal 
stresses, is used throughout the simulations, which is consistent with the 3DEC modeling 
discussed previously. The equilibrium state of the model after excavation of a drift represents 
the initial condition for the dynamic analysis. This equilibrium state is achieved by performing a 
quasi-static simulation whose geometry, static boundary, and initial conditions are illustrated in 
Figure 5-21. 
5-33 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 6-119. 
Figure 5-21. Dynamic Model, Initial and Boundary Conditions: Initial Static Simulation 
The boundary conditions used in the dynamic analysis are illustrated in Figure 5-22. Quiet 
boundaries (indicated in Figure 5-22 as viscous boundaries) were used on all models outside 
boundaries. These boundaries prevent reflection of outgoing seismic waves back into the model. 
Quiet boundaries were combined with free-field boundaries on the vertical outside boundaries 
that prevent distortion of vertically propagating plane waves along the boundaries. Dynamic 
loading was applied at the bottom of the model, as propagating vertically upwards. 
5-34 June 2004 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 6-120. 
Figure 5-22. Dynamic Model Boundary Conditions for Dynamic Simulation 
Preclosure CaseThe preclosure ground motions, defined here as annual exceedance 
probabilities of 10-4 or greater, were conservatively examined for unsupported tunnels. The 
analyses indicate that ground motion with a probability of an annual occurrence of 5  10-4 and 
10-4 will not induce any rockfall for rock mass categories 2 through 5. As discussed in 
Section 4.2.2.1, the estimated mean rock mass condition is Category 3, with about 90% of the 
Tptpll having Category 3 or greater properties. A relatively small amount of rockfall from the 
drift walls (shown in Figure 5-23) is expected, even for the lowest quality (less than 10% of all 
lithophysal rock) Category 1. In reality, no rockfall would occur in preclosure since the 
emplacement drifts are fully supported with rock bolts and surface steel sheeting (see 
Figure 1-4). 
Stress monitoring locations within the rock mass surrounding the tunnel show that dynamic 
stress changes induced by this ground motion are small, and the rock mass remains well within 
the elastic range. The observed rockfall in the Category 1 case is simply the small springline 
zone of yielded rock from in situ stresses being shaken down. The conclusion is that only minor 
sidewall sloughing would be expected from this ground motion, and only in the poorest quality 
rock. Light bolting and surface support would easily contain this material. 
5-35 June 2004 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 6-127. 
NOTE: Rock mass categories 2 to 5 show no damage from the 1  10-4 ground motion. Red lines indicate block 
bonds that have failed in shear or tension. 
Rock Mass Category 1 
5.3.2.2.2.1 Postclosure Analyses 
10-5 Ground Motion Parametric AnalysesAll 15 sets of ground motion time histories were 
applied to unsupported base case rock mass strength Categories 1, 3, and 5 to provide a crosssection 
of response across the entire range of rock qualities. The level of damage induced by the 
ground motion is quantified here by the area (volume) of the rock that yields and is detached 
from the surrounding rock mass and falls into the tunnel. Since the model is two dimensional, it 
essentially represents a 1-m-thick slice of material parallel to the axis of the drift. The damage is 
this thus given in terms of volume per meter of emplacement drift, or cubic meters per meter. 
This could be converted to metric tons by multiplying by the density, which for the lithophysal 
rock is approximately 2 tons/m3 (assuming a matrix density of approximately 2.5 tons/m3). 
10-5 Ground Motion Damage AssessmentFigures 5-24 to 5-26 show representative examples 
of the mechanism and level of damage induced for rock strength Categories 1, 3, and 5 for 
ground motion 12 (1.04 m/s peak ground velocity), 4 (1.52 m/s peak ground velocity), and 
7 (3.33 m/s peak ground velocity). These ground motion data cover the approximate range of 
peak ground velocities for all 15 motions. The Seismic Consequence Abstraction (BSC 2004e) 
relates both rockfall and vibratory damage to the drip shield directly to the peak ground velocity 
of all three components of any particular ground motion. Since peak ground velocity and its 
variability are directly related to annual exceedance frequency, a correlation of damage to 
probability can be developed. 
The drift damage levels are plotted for each rock mass strength category as a function of peak 
ground velocity in Figure 5-27. The analyses show that the damage is related to the magnitude 
of peak ground velocity, with significant variability at large peak ground velocity. Approximate 
linear upper and lower bounds are shown for each rock strength category, with a damage band 
shown that covers the range of variability. The variability in the response is a function of the 
Figure 5-23. Geometry of the Model after Simulation for Preclosure Ground Motion (Probability 10-4): 
June 2004 5-36 No. 4: Mechanical Degradation Revision 1 
ground motion, not the rock properties, since constant modulus and strength are assumed for a 
given category. This variability at large peak ground velocity can be explained by examining the 
damage as a function of the total energy in the waveform. A Fast Fourier Transform (FFT) was 
used to perform an integration of each of the 15 10-5 velocity time histories (in terms of the 
velocity squared). The result of this integration is a spectral power density number that is 
proportional to the total kinetic energy (kinetic energy = (mv2)) of the time history. 
The damage as a function of the velocity power spectral density for the ground motion 
component (either H1 or V, whichever has the maximum value of peak ground velocity) for each 
of the 45 analyses is plotted in Figure 5-28. This plot shows that the damage is linearly related to 
the kinetic energy associated with the velocity time history. Therefore, although the peak ground 
velocity is, in general, related to drift damage, the variability in that correlation is related to the 
total kinetic energy in the time history. 
June 2004 5-37 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 6-128. 
NOTE: (upper left) Rock Strength Category 1 (5.6 m3/m drift length), (upper right) Category 3 (0.3 m3/m), (lower left) 
Category 5 (0.2 m3/m), and (lower right) Ground Motion History 12, peak ground velocity = 104 cm/s. 
Figure 5-24. Example of Comparison of Damage Levels for Lower End of Peak Ground Velocity 
(104 cm/s) for 10-5 Annual Exceedance Level 
June 2004 5-38 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 6-129. 
NOTE: (upper left) Rock Strength Category 1 (3.6 m3/m drift length), (upper right) Category 3 (2.3 m3/m), (lower left) 
Category 5 (0.3 m3/m), and (lower right) Ground Motion History 4, peak ground velocity = 152 cm/s. 
Figure 5-25. Example of Comparison of Damage Levels for Peak Ground Velocity of 152 cm/s for 10-5 
Annual Exceedance Level 
June 2004 5-39 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 6-130. 
NOTE: (upper left) Rock Strength Category 1 (49.7 m3/m drift length), (upper right) Category 3 (16.1 m3/m), (lower 
left) Category 5 (21.1 m3/m), and (lower right) Ground Motion History 7, peak ground velocity = 333 cm/s. 
Figure 5-26. Example of Comparison of Damage Levels for Upper End of Peak Ground Velocity 
(333 cm/s) for 10-5 m/s 
June 2004 5-40 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 6-131. 
NOTE: Upper and lower linear bounds are drawn empirically to cover entire range of results for all 15ground 
motion data and are given for visual reference only. Black data points are results for simulations in which 
rock mass has spatially-variable lithophysal porosity. 
Figure 5-27. Estimate 10-5 Damage Level, Expressed as m3/m of Emplacement Drift Length for Rock 
Strength Categories 1 (Top), 3 (Center), and 5 (Bottom) for All 15 Ground Motion Data. 
June 2004 5-41 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 6-132. 
NOTE: The power spectral density is obtained by integrating the square of the velocity time history, producing a 
value proportional to the kinetic energy. 
Figure 5-28. Rockfall Damage as a Function of the Energy Associated with the Vertical Velocity Time 
History 
Damage MechanismThe drift damage mechanism consists primarily of stress-induced failure 
of the rock mass resulting from the stress change associated with the velocity time history. The 
in situ stress field has major vertical and minor horizontal stress components. The vertical 
compression or horizontal shear wave essentially results in a free field dynamic stress increase 
equivalent to (Itasca Consulting Group 2002): 
n,s = Cp,sVn,s 
where: n,s = dynamic induced normal (n) or shear (s) stress component 
 = rock density 
Cp,s = speed of p-wave (p) or s-wave (s) propagation in rock medium 
Vn,s = peak normal (n) or shear (s) ground velocity 
These dynamic components are superimposed on the existing in situ stress field to cause 
additional stressing or relaxation of the rock mass surrounding the drift. The end result of this 
superposition is that the stress tensor changes, both in magnitude and orientation of the principal 
stresses, as the ground velocities oscillate over the duration of the strong ground motion. 
Superposition of in situ and dynamic stresses can cause tensile or shear failure of rock mass, 
resulting in development of an elliptic shape of the opening as the rock mass yields and rockfall 
occurs and falls along the sides of the drip shield. Shear failure is most likely to start at the 
June 2004 5-42 No. 4: Mechanical Degradation Revision 1 
springline, which is the location of the largest stresses under in situ stress conditions. The extent 
of shear failure and rockfall around the circumference of the tunnel, up and down from the 
springline, is due to both the general ratio of rock mass strength to stress, but also to the ratio of 
the vertical to horizontal peak ground velocity. The greater the horizontal component, the 
greater the rotation of the stress tensor, which results in greater inclination of the major principal 
stress. Generally, this shear failure mechanism occurs with the arrival of the peak ground 
velocities. Compressive stresses also appear responsible for some cases in which roof slabbing 
is observed where the rock mass strength and stiffness are larger (i.e., Category 5). A second 
failure mechanism observed includes tensile failure of the rock mass resulting from the reversal 
of the ground motion and inducement of dynamic tensile straining in the rock mass (i.e., when 
seismically-induced tensile stress exceeds in situ compressive stress and tensile strength, 
combined). 
In general, it appears that the rock mass failure and thus rockfall occurs simultaneously with the 
arrival of velocity peaks in the time histories. In a similar fashion, the peak ground velocity has 
been recognized by many authors as the primary contributing factor to dynamic rock mass failure 
in mine tunnels, slopes and dams (e.g., Newmark 1965). In the present case, the shear or tensile 
failure mechanism results in a predicted creation of blocks resulting from fracture of the rock 
mass along the ubiquitous fracture network of the material. The blocks that contact the drip 
shield are relatively small in size and governed by the inherent fracture network and lithophysae 
spacing. The UDEC model studies have used primarily a 0.3 m average fracture spacing in 
developing the block structure, although simulations with 0.2 m spacing have also been 
conducted. In all cases, the transient dynamic stress changes, in addition to preexisting in situ 
stresses, results in breakage of the bonds between the blocks in the rockfall zone around the 
tunnel. In other words, the stressing does not create large blocks that impact the drip shield, but 
small component blocks defined by the inherent fracturing of the Tptpll. 
Although the damage levels appear to correlate somewhat better to the energy content of the time 
history, correlation to peak ground velocity provides a simpler method for interpretation of the 
results. An approximate relationship of damage level to peak ground velocity in terms of both 
m3/m of drift length and physical interpretation can be roughly approximated as follows: 
 Damage level below 5 m3/m results in minor damage to rock particles filling the invert 
along the sides of the drip shield  a peak ground velocity below about 1.5 m/s (see 
Figure 5-24). 
 Damage level from 5 to 15 m3/m results in rock particles covering the sides of the drip 
shield approximately to the height of the drip shield and may cover the top of the drip 
shield  this corresponds to peak ground velocity values of approximately 1.5 to 2 m/s 
(see Figure 5-25). 
 Damage level above 15 m3/m causes complete collapse of the tunnel, this corresponds to 
peak ground velocity values of approximately 2 to 3 m/s (see Figure 5-26). 
Impact of Spatial VariabilityIn the previous analyses, the rock mass properties were 
considered homogeneous for a given drift cross section. Here, the impact of considering actual 
spatial variability of lithophysal porosity on damage from 10-5 ground motion is examined. 
June 2004 5-43 No. 4: Mechanical Degradation Revision 1 
A representative section of the Tptpll was extracted from the upper portion of the lithophysal 
porosity model as described in Section 2.3.2. This model contains a range of lithophysal 
porosity from less than 10% to greater than 25%, averaging approximately 17.5%. The resulting 
UDEC model showing spatially variable porosity is given in Figure 5-29. 
Source: BSC 2004a, Figure 6-133. 
NOTE: Lithophysal porosity variability derived from Figure 2-11. 
Figure 5-29. 
(Bottom) 
Rock mass strength properties for Categories 1 to 5 were interpolated in the model based on the 
lithophysal porosity levels, achieving spatial variability in strength and moduli. This model was 
subjected by all 15 10-5 ground motion data, and damage levels determined. The results of the 
dynamic simulations, in terms of damage versus peak ground velocity, are shown in Figure 5-27. 
The damage levels are approximately in the range of the Category 3 rock mass as expected since 
the mean lithophysal porosity of the model falls within the range of the Category 3 levels. This 
analysis indicates that the use of homogenous rock properties that span the range of strength 
categories does, indeed, span the range of expected response, including conservative damage for 
the low strength categories. 
Drip Shield Loading from Seismic DamageThe loading of the broken rock mass after drift 
collapse for the 10-5 case is examined for the assumption of rigid and deformable drip shields. 
Two cases of drip shield deformability are examined: (1) rigid, rectangular geometry, and 
(2) deformable, arched roof shape (Figure 5-30). The rigid assumption, used due to its 
simplicity, is the base case assumption. The deformable case provides a two-dimensional model 
Contours of Lithophysal Porosity Contoured on the UDEC Spatial Variability Model (Top) 
and a Histogram Showing the Percentage of Lithophysal Porosity within the Model 
June 2004 5-44 No. 4: Mechanical Degradation Revision 1 
of the actual shape and deformability of the drip shield using the standard UDEC element 
structure. The drip shield is subdivided into 30 total segments of equal length. In the 
deformable case, the elements are assumed to be elastic. Since the UDEC model is 
two-dimensional and does not attempt to model the detailed structure of the drip shield, the 
stiffness of the elements must be adjusted to provide an equivalent overall deformability to the 
actual three-dimensional structure. The stiffness of the UDEC elements has been calibrated 
against the three-dimensional LS-DYNA model to reproduce the deformability of the actual drip 
shield design (BSC 2004a). The dynamic loading from detached rock particles during the 
seismic shaking (includes the effect of relative velocity between the drip shield and the falling 
rock) is examined here. The quasi-steady load on the drip shield once the seismic shaking stops 
and the system comes to force equilibrium, and, the dynamic loading of the drip shield when it is 
covered by broken rock and subjected to additional seismic events, are discussed in 
Section 5.3.2.3. 
NOTE: Simple rigid, rectangular structure with footings rigidly fixed to the invert (left) and a deformable, actual 
geometry with footings pinned (i.e., translational motion slaved to the invert) (right). Drip shield footings can 
also slide on and separate from the invert. 
5.3.2.2.2.2 
Figure 5-30. Schematic Representation of Two Cases of Drip Shield Loading Showing (a) Rigid 
Rectangular Geometry and (b) Deformable Arched Geometry 
June 2004 
Dynamic Impact Loading of the Drip Shield 
An example of the dynamic impact loading from side and top impacts to the deformable drip 
shield resulting from the 10-5 ground motion (peak ground velocity = 3.33 m/s case) is given in 
Figure 5-31. In this figure, the average pressure across an element (in Pa) is given as a function 
of time (in seconds). In general, the transient pressure at any given element is approximately one 
order of magnitude greater than the eventual dead weight load of bulked rock for the case of a 
completely collapsed tunnel at equilibrium. However, due to the relatively small size of the rock 
particles developed in the lithophysal rock (in comparison to the much larger blocks in the 
nonlithophysal material), the nonlithophysal dynamic impact loading to the drip shield provides 
a more conservative, bounding case. Therefore, the nonlithophysal rockfall case is used to 
provide dynamic impact loading for the damage analysis and design of the drip shield. 
5-45 No. 4: Mechanical Degradation Revision 1 
Quasi-static loading of the dead weight of the rock on the drip shield is described in 
Section 5.3.2.3. 
Source: BSC 2004a, Figure 6-135. 
NOTE: Segments are numbered counterclockwise from the right side footing. Segments 1 to 10 are found along 
the right wall, 11 to 20 along the roof. 
Figure 5-31. Example of Dynamic Impact Loading to the Right Wall (Top) and Roof (Bottom) of 
Deformable Drip Shield with Arched Roof 
June 2004 5-46 No. 4: Mechanical Degradation Revision 1 
10-6 Ground Motion Parametric AnalysesFifteen 10-6 ground motion time histories are 
characterized by the peak ground velocity of the H1 component of 2.44 m/s, resulting in a 
vertical component with mean of 1.11 m/s (Table 5-1 summarizes peak ground motion values). 
The analyses of the 10-5 ground motion showed extensive damage for peak ground velocity in 
excess of approximately 2 m/s, and therefore, similar damage is expected for all the 10-6 and 
10-7 ground motion cases. A typical model geometry after simulation of the 10-6 motion is 
shown in Figure 5-32 showing complete collapse of tunnels in lithophysal rock. The extent of 
failure and detached rock mass forms a roughly elliptical shape that extends approximately one 
tunnel diameter into the roof of the drift. The rock mass failure mechanism, a combination of 
shear and tension failure, is similar to that described for the 10-5 case. As was discussed in the 
previous section, tunnel collapse occurs for peak ground velocity values in excess of 
approximately 2 to 3 m/s. 
Source: BSC 2004a, Figure 6-133e. 
NOTE: Blocks are colored by magnitude of displacement. 
Figure 5-32 Typical Geometry of the Model after Simulations for Postclosure Ground Motion (Probability 
10-6) 
June 2004 5-47 No. 4: Mechanical Degradation Revision 1 
5.3.2.2.3 Combined Seismic, Thermal, and Time-Dependent Effect in Lithophysal 
Units 
The initial condition for the seismic analysis discussed in the previous section was the in situ 
stress state perturbed only by excavation of the drifts. Throughout the postclosure phase, the 
rock mass temperature, and the associated thermally induced stress state will vary as discussed in 
Section 5.3.1. These stresses, particularly in the first one thousand years of postclosure time, are 
potentially significant as initial condition for to the transient seismic stressing. The impact of the 
thermal stresses is examined here. 
Preclosure temperature and stress increase from heating is small and has no appreciable impact 
on drift degradation over and above preclosure seismic loading under in situ temperature 
conditions. Therefore, the maximum thermal condition (i.e., approximately 30 years after 
cessation of forced ventilation) is used as an initial condition for the seismic analysis 
(Figure 5-4). 
Because the ground motion with 10-6 probability of annual occurrence results in complete drift 
collapse, it was not of particular interest to investigate the effect of that level of ground motion 
combined with thermally induced initial stresses. Instead, a ground motion of 5  10-4 
probabilities of annual occurrence was considered. Rock mass categories 1 and 5 were 
considered in this analysis. 
Seismic analysis using 1  10-4 ground motion after 80 years of heating for rock mass 
Category 1 (lowest strength) resulted in an increased rockfall compared to rockfall from the 
seismic shaking of the rock mass at the in situ stress state only (Figure 5-33, Category 1 (a)). 
Therefore, the heating induces additional damage (compared to damage caused by drift 
excavation), which does not necessarily result in a rockfall under static loading conditions but is 
shaken down by the 1  10-4 ground motion. No rockfall is induced in rock mass Category 5 in 
the case of 90% ventilation efficiency (Figure 5-33, Category 5 (b)). 
June 2004 5-48 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figures R-31 and R-32. 
NOTE: Yield at the in situ stress condition for Strength Categories 1 and 5 can be found in Figure 4-17. 
Figure 5-33. Rockfall and Fractures Induced around a Drift by 1  10-4 Preclosure Ground Motion after 
5.3.2.2.4 
the Peak Thermal Condition Occurring at 80 Years of Heating (30 Years after Closure) in 
Rock Mass Categories 1 and 5 
June 2004 
Time-Dependent Consideration in Lithophysal Units 
Underground and surface excavations, which are designed to be stable after excavation, degrade 
with time, and some eventually collapse completely. The degradation of excavations in hard 
rocks has not been studied extensively because most underground excavations have service lives 
of 100 years or less and are maintained as required. However, there are many examples of 
stable, unsupported excavations associated with mining, and civil construction or naturally 
occurring caves in numerous rock types that are stable, or have suffered only minor instability 
for hundreds or even thousands of years (BSC 2002f, Section 3). Thus, there is no certainty that 
5-49 No. 4: Mechanical Degradation Revision 1 
collapse of unsupported excavations, particularly those subjected to low stress, relatively dry 
conditions, is inevitable. 
The primary reason for eventual yield and collapse is that a hard rock mass, exposed to humidity 
and temperature of the open atmosphere, may undergo strength decay with time when it is loaded 
to stress levels higher than about 50% to 60% of its short-term strength. The rate of strength 
decay depends on, among other parameters, rock type (particularly the mineralogy and grain 
structure), stress state, relative humidity, and temperature. Stress corrosion is considered the 
primary mechanism causing strength degradation of hard rocks (Potyondy and Cundall 2001, 
Section 3). 
The emplacement drifts at Yucca Mountain will be stable under currently existing conditions 
(in situ stresses and rock mass strength) with ground support as demonstrated by the calculations 
presented here. However, it is expected that the ground support will completely lose its integrity 
during the 10,000-year regulatory period, and drift degradation, to some extent, will occur due to 
strength decay of the rock mass. Drift degradation is an important issue for the repository design 
and performance as drifts have to remain open during the preclosure period; and eventually the 
caved rock resulting from degradation will load the drip shields possibly affecting their integrity 
and performance. Prediction of the rate of drift degradation for the duration of the 10,000-year 
regulatory period is a highly approximate task (it requires extrapolation of testing results, which 
have been done for a period of months, to a period of 10,000 years). Uncertainty in such 
predictions is quite large. The rate of drift degradation has been analyzed and is documented in 
detail in Drift Degradation Analysis (BSC 2004a, Appendix S). 
Time-Dependent Drift Degradation Analysis MethodologyThere is currently no accepted 
methodology for estimating the time-dependent degradation behavior of hard rocks. The 
empirical term stand-up time is often used in the mining industry to refer to an estimate of the 
amount of time that personnel can safely be sent beneath unsupported drill-and-blast tunnel 
headings without fear of loose material creating a working hazard (Bieniawski 1989). This 
stand-up time is typically on the order of hours or days, even for good quality rock masses, and 
is often confused with actual drift instability and collapse mechanisms. 
Unsupported or lightly supported tunnels (although perhaps not safe from a personnel 
standpoint) can stand in stable condition for long time periods, particularly in good quality rock 
masses. For example, the ESF and ECRB Cross-Drift tunnels were constructed in 1995 to 1997 
and in 1998, respectively. Although the ESF main loop is located largely in the Tptpmn, the 
ECRB Cross-Drift cuts through and exposes all of the repository host horizon units. The tunnels 
are, in general, lightly supported with friction rock bolts and light wire mesh in the tunnel roof, 
with occasional friction bolts in the tunnel walls. There is no evidence of significant 
deterioration or degradation of the rock mass in these tunnels since excavation, and no 
significant rockfall has occurred. Tunnel steel set load (CRWMS M&O 1998b) and deformation 
measurements (CRWMS M&O 1998c) have been regularly monitored since excavation, showing 
stable conditions as well. The rock mass is obviously in a stable and self-supporting mode with 
no obvious deterioration in 5 to 8 years. 
A key mechanism of failure in low porosity, brittle, crystalline rocks is development in the 
propagation of cracks parallel to the greatest principal stress direction. Extensive studies over 
June 2004 5-50 No. 4: Mechanical Degradation Revision 1 
the last 40 years have demonstrated that the critical parameters that control crack growth are 
stress, temperature, and the partial pressure of water at the crack tip. A crack grows due to the 
hydration and breaking of siliconoxygen bonds at the tip of the crack. The rate at which the 
crack grows is controlled by the diffusion of water to the crack tip. This mechanism is 
commonly termed stress corrosion cracking. Experimental data on single crystals of quartz, as 
well as in rocks, have validated these mechanisms (e.g., Martin 1972; Kranz 1979). 
In brittle rocks subjected to a constant stress, a crack will propagate in a time-dependent way. 
The diffusion of water to the crack tip controls the rate. Since the physical mechanism is 
hydration, there is a volume increase on the surface of the crack. Consequently, the propagation 
of water to the crack tip is controlled by diffusion along a crack surface where the aperture 
decreases and diffusivity decreases. This gives rise to a logarithmic time-dependent crack 
growth. The rate at which water diffuses to the crack tip decreases with time, the rate of crack 
growth slows down, and the observed deformation rate decreases. When the crack reaches a 
critical length, the rock fails. Furthermore, the lower the stress for a constant temperature, the 
longer the time to failure (Kranz 1980; Martin et al. 1997a). 
With the understanding of the physical mechanisms of crack growth, it is reasonable to 
understand how the time dependence works. The rate dependence is a physical process that is 
scalable. The common experimental testing technique used to determine the time to failure for 
brittle rocks is the static fatigue compression test, which is essentially the same as a creep 
experiment. The rock sample is loaded in uniaxial or confined compression to a load that is a 
given percentage of its strength. This load is held constant until the sample fails, and the time to 
failure is recorded. Since the rate dependence is scalable, static fatigue experiments conducted 
for time frames from several minutes to several months can be extrapolated to much longer times 
with confidence. A limited amount of static fatigue compression test data on the Tptpmn is used 
as the basis for extrapolation of the time-to-failure as a function of applied stress conditions. 
This data is supplemented by the more extensive literature data on static fatigue of granite as a 
means of comparison of failure time-rock strength relationships. 
The PFC stress corrosion model by Potyondy and Cundall (2001), developed for similar timedependent 
predictions for the Canadian waste disposal research program, has been extensively 
documented and validated against time-dependent tunnel breakout in granite at the Underground 
Research Laboratory in Manitoba, Canada. The PFC stress corrosion model is first calibrated to 
reproduce the static fatigue response of LdB granite and interpolated for nonlithophysal tuff 
(i.e., the Tptpmn). The model is then used to investigate the impact of lithophysal porosity on 
the rate of time-dependence, resulting in the generation of a set of curves of time-to-failure 
versus applied stress level for various levels of lithophysal porosity. 
A rock mass damage-time relationship, developed from the PFC simulations, is embedded in the 
same UDEC model used earlier to perform lithophysal stability calculations. This model is used 
to generate estimates of drift degradation as a function of time for a range of rock mass strength 
categories. The model is validated for short time periods against observations of the existing 
ESF and ECRB drifts, and for long time periods against observations of lack of observed 
yielding in the Tptpll. This work is reviewed here; details of the methodology and analysis 
methods are referred to Drift Degradation Analysis (BSC 2004a, Appendix S). Currently, a 
significant program in static fatigue measurement of Tptpmn cores is being conducted to 
June 2004 5-51 No. 4: Mechanical Degradation Revision 1 
supplement the limited existing data base. This present work will be updated as this long-term 
testing data is developed. 
Static-Fatigue Behavior of Granite and TuffCMartin et al. (1997a) present static-fatigue 
results for a total of 16 specimens of welded (lithophysae poor) tuff from borehole NRG-7/7A at 
Yucca Mountain and from Busted Butte boulders taken from the same block of rock. The 
specimens were 2:1 aspect-ratio right circular cylinders with a diameter of 50.8 mm. Load 
application was rapid, with full load being reached in less than 10 seconds. 
The seven borehole specimens were tested, drained, and vented to the atmosphere at a 
temperature of 225C and a confining pressure of 10 MPa at differential stresses ranging from 
40 to 130 MPa. None of these specimens had failed after loading for times ranging from 
2.5  106 to 5.9  106 seconds. The nine Busted Butte specimens were tested at a pore water 
pressure of 4.5 MPa, a temperature of 150C, and a confining pressure of 5 MPa at differential 
stresses ranging from 115 to 150 MPa. The high pore pressure (i.e., 4.5 MPa) ensured that pore 
water remained in the liquid state even though the temperature is above the boiling point. The 
test results are summarized in Table 5-8. The applied load in the axial direction and the effective 
confining pressure are denoted by 1 and Pc, respectively. The axial load at failure during a 
short-term test is denoted by f. The stress difference maintained during a static-fatigue test 
conducted at a confining pressure of Pc is  = 1 C Pc. The stress difference at failure during a 
short-term test is c = f C Pc. 
To facilitate comparison between different data sets, static-fatigue curves were generated by 
plotting the logarithm of time-to-failure, tf, versus the driving-stress ratio given by  /c = 
(1 C Pc) / (f C Pc). Six of these specimens failed at times less than 2  106 seconds, while the 
remaining three specimens (BB-9392-H, -G, and -J) did not fail during the testing period. The 
times-to-failure for these six tests can be plotted versus applied load (Figure 5-34); however, the 
peak strength must be estimated in order to plot them versus driving-stress ratio for comparison 
with data from the Lac du Bonnet granite (Schmidtke and Lajtai 1985; Lau et al. 2000). For 
these purposes, the peak strength of the tuff samples at an effective confinement of 0.5 MPa is 
estimated to be 151 MPa, to give a failure time of one second for a driving-stress ratio of unity.19 
Approximate linear relationships in semi log space have been fit to the unconfined and confined 
granite and tuff data. The granite data shows a flatter slope, or faster time-to-failure than the tuff 
for similar ratios of the applied stress to unconfined compressive strength in this plot. This 
would be expected as the tuff is a fine-grained volcanic that shows very little hysteresis on 
unloading until brittle failure occurs. The granite, on the other hand, is composed of coarser 
mineral grain structure of several different minerals and exhibits permanent deformation at lower 
strain levels. 
19 The unconfined compressive strength values by Martin et al. (1997a) for six saturated 50.8-mm-diameter Busted 
Butte specimens tested at a strain rate of 10.5/s ranged from approximately 105 to 200 MPa, with a mean of 
approximately 128 MPa for the five weakest specimens. 
June 2004 5-52 No. 4: Mechanical Degradation Revision 1 
Table 5-8. Static-Fatigue Data for Busted Butte Specimens 
 tf f Specimen 
BB-9392-K 
BB-9392-N 
BB-9392-E 
BB-9392-C 
BB-9392-F 
BB-9392-B 
BB-9392-H 
BB-9392-G 
BB-9392-J 
/c 
0.99 
0.94 
0.89 
0.89 
0.85 
0.85 
0.87 
0.87 
0.76 
log(tf) 
(sec) 
0.08 
0.60 
2.40 
2.80 
3.77 
6.29 
6.07 
5.86 
6.30 
(sec) 
1.2 
4 
250 
636 
5,848 
1,960,000 
1,180,000 
732,000 
2,000,000 
(MPa) 
149.0 
141.0 
134.6 
134.2 
132.8 
127.8 
131.4 
131.3 
115.0 
Source: Martin et al. 1997a. 
BB-9392-G, and BB-9392-J did not fail during the test. 
Pc 
(MPa) 
0.5 
0.5 
0.5 
0.5 
0.5 
0.5 
0.5 
0.5 
0.5 
Source: BSC 2004a, Figure 6-148. 
(MPa) 
151 
151 
151 
151 
151 
151 
151 
151 
151 
NOTE: Specimens were saturated and tested at a pore water pressure of 4.5 MPa and 
temperature of 150C. Specimens were loaded directly to creep stress (1) in 
less than 10 seconds. Specimen diameter is 50.8 mm. Specimens BB-9392-H, 
NOTE: Tests of Lac du Bonnet granite conducted at 25C. Tuff tests conducted at 150C. LdB = Lac du Bonnet. 
Figure 5-34. Static-Fatigue Data for Welded Tuff and Lac du Bonnet Granite 
June 2004 5-53 No. 4: Mechanical Degradation Revision 1 
PFC Corrosion Model Calibration and Examination of Impact of Lithophysal VoidsCThe 
static fatigue data for welded tuff shown in Figure 5-34 is for nonlithophysal samples. Since the 
majority of the repository lies within lithophysal rock, it is necessary to determine the impact of 
lithophysal voids on the static fatigue response (essentially, the slope of the best-fit lines in this 
figure). 
Since the matrix material of the nonlithophysal and lithophysal tuff is mineralogically and 
mechanically similar, an approach to estimate the impact of lithophysae on the time-dependency 
can be developed based on the existing static fatigue data. The PFC program is first calibrated to 
reproduce the static fatigue response of the nonlithophysal rock. Using the same matrix 
properties, lithophysal porosity is added and time-dependency extrapolated. This is a reasonable 
approach because the primary effect of lithophysae is to adjust the internal stress condition of the 
sample, which, in turn, impacts the time-to-failure. PFC automatically determines the internal 
stress redistribution and concentration in the sample and thus accounts for this primary effect. 
The PFC model for lithophysal tuff was described previously in Section 4. This PFC model 
(consisting of circular voids within a well-connected base material) has the same matrix material 
for which the stress-corrosion behavior is measured. The long-term behavior of the PFC 
material is characterized by performing a series of static-fatigue tests on the PFC lithophysal-tuff 
model. Recall that the PFC model consists of rigid, circular particles interconnected at their 
contact points. The macroscopic material behavior of the PFC model is governed by the strength 
and stiffness properties of the contact points. The PFC stress corrosion model uses the same 
construction and bonding characteristics as the model presented earlier, with the difference being 
that the bond strengths include time-related parameters consistent with a stress corrosion 
cracking mechanism. These time-dependent bond parameters (1, 2, and ҡ a) are determined via 
calibration of the PFC model to static fatigue time-to-failure test data. These parameters do not 
affect the short-term behavior. The properties of the PFC2D material are obtained by 
numerically testing 1:1 aspect-ratio specimens (Figure 5-35) of one-meter diameter with void 
porosities of 0, 0.1, and 0.2 under static-fatigue conditions at confinements of 0.1 and 5 MPa. 
June 2004 5-54 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 5-5. 
NOTE: Examples have void porosities of 0.107 (a) and 0.204 (b). 
Figure 5-35. PFC2D Specimens of Tuff Material for Static Fatigue Testing 
Figure 5-36 shows a typical creep curve and damage plot from a PFC nonlithophysal rock test 
in which the sample uniaxial load is held constant at 80% of the unconfined compressive 
strength. Damage occurs in the form of formation of cracks as a function of time. As seen, the 
sample undergoes time-dependent deformation until approximately 5,000 seconds, when creep 
rupture occurs and the sample fails. Numerical tests such as these are used first to calibrate the 
PFC model against the laboratory data for nonlithophysal tuff, but then to extrapolate the 
time-to-failure data to lithophysal rock with varying levels of lithophysal porosity. 
Figure 5-37 shows a plot of the time-to-failure for different levels of sample porosity as a 
function of the ratio of applied stress to unconfined compression strength for lithophysal tuff 
samples. As one might expect, as the lithophysal porosity increases, the time-to-failure of the 
rock mass decreases, presumably due to the presence of increased level of tensile stresses in the 
material. 
Inclusion of Time-Dependency Into an Engineering Approximation of Drift Degradation 
To utilize the time-to-failure data for lithophysal tuff, it must be generalized into an 
engineering-based model that can predict the evolving stress state around excavations resulting 
from time-dependent strength degradation as well as from in situ, thermal, and seismic loading. 
The same UDEC approach that has been used for the previous calculations is used for this 
purpose. Because the UDEC model is used in a quasi-static mode in this regard, it is most 
convenient to frame the time-to-failure data in the form of a simple damage coefficient that 
represents the time-evolution of strength degradation. 
June 2004 5-55 No. 4: Mechanical Degradation Revision 1 
Damage Coefficient DerivationCThe axial load at failure (peak strength) during a short-term 
test performed at an elapsed time, t, since the start of a static-fatigue test is denoted by *
f = *
f 
(t). The values of *
f are bounded by: 
*
f (0)  f 
(Eq. 5-1) *
f (tf)  (/c ) (f . Pc) + Pc 
where (/c ) is the ratio of the driving stress to the unconfined compressive strength of the 
material, and Pc is the confining pressure of the test. The values of *
f for times 0 < t < tf are 
found by stopping the static-fatigue test at the desired time and measuring the peak strength. 
The strength degradation is quantified by means of a damage coefficient: 
(Eq. 5-2) 
(Eq. 5-3) 
D = 1 . *
c /c 
where *
c = *
f . Pc is the principal stress difference at failure. Substituting values from 
Equation 5-1 into this expression provides the following bounds for the damage coefficient: 
D(0) = 0 
D(tf) = 1 . (/c ) 
The time evolution of the damage coefficient for the nonlithophysal tuff material tested at a 
confinement of 0.1 MPa is shown in Figure 5-38. These results were produced by performing 
10 numerical compression tests conducted on damaged samples derived from four 
static-fatigue tests at driving stress ratios of 0.8, 0.6, 0.4, and 0.2. In other words, the PFC stress 
corrosion model spontaneously produces stress corrosion fracturing as the simulated static 
fatigue test is run. To determine the impact of this damage on the strength of the sample, the 
static fatigue simulations are stopped at various elapsed times and the damaged sample 
compressed to determine its peak strength. The loss of cohesive and tensile strength, expressed 
as a ratio of the original strength, is termed the damage coefficient. Most damage occurs during 
the final stages of a static-fatigue test; tests performed at lower driving-stress ratios produce an 
earlier (in terms of normalized times-to-failure) onset of damage. Damage evolution for 
lithophysal tuff follows a similar form, with the time factor determined from the time-to-failure 
behavior illustrated in Figure 5-37. 
5-56 June 2004 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 5-12. 
Figure 5-36. Creep Curve and Damage in mS50 Material for Static-Fatigue Test (0.1 MPa Confinement) 
at Driving-Stress Ratio of 0.8 
Source: BSC 2004a, Figure 5-16. 
NOTE: The time-dependency has approximately the same slope for all void porosities. Straight-line fit. 
Figure 5-37. Effect of Void Porosity on Static-Fatigue Curves (0.1 MPa Confinement) for Lithophysal 
Tuff Material (0% to 20% Void Porosity) 
June 2004 5-57 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 5-21. 
Figure 5-38. Time Evolution of Damage Due to Strength Degradation Coefficient for Nonlithophysal Tuff 
Material during Static-Fatigue Tests (0.1 MPa Confinement) at Driving-Stress Ratios 
Ranging from 0.2 to 0.8 
Static-Fatigue Curves and the Evolution of Damage Due to Strength DegradationThe 
static-fatigue behavior of Lac du Bonnet granite and welded lithophysal tuff forms the basis of 
the UDEC model for stress corrosion around a drift. The static-fatigue curves provide the 
time-to-failure (tf ) of the material at a particular driving-stress ratio ( / c 
). 
The static-fatigue data for Lac du Bonnet granite at 0 and 5 MPa confinement and tuff at 5 MPa 
confinement and 4.5 MPa pore pressure are shown in Figure 5-34. Each data set was fit with a 
straight line, and the line was extrapolated to encompass driving-stress ratios ranging from 0 
to 1. This is a conservative assumption, because the curves most likely approach infinity at a 
driving-stress ratio less than about 0.5. Three lines for the tuff data are shown in Figure 5-39. 
The blue line in the figure is the least-square linear fit through the tuff data, while the purple line 
(best fit through origin curve) is the best-fit for the tuff data for unconfined compressive strength 
(i.e., failure at 1 s at the unit driving ratio of 1.0). Since the tuff data are very limited and are for 
confined conditions (effective confining stress of 0.5 MPa) only, the data input to UDEC uses a 
simplified best fit curve (red line) termed the tuff best fit curve. The tuff best fit curve will be 
updated as the ongoing test data become available. 
Using the time-to-failure versus stress-to-strength ratio (Figure 5-39) and the damage coefficient 
evolution plot, a simplified, general damage evolution curve in terms of time can be developed 
and used directly in the UDEC program for drift stability modeling. Figure 5-40 presents these 
relations for the tuff best-fit curve. This family of curves can be used to directly relate reduction 
in cohesive and tensile strength of the rock mass as a function of stress state and elapsed time 
since excavation. 
June 2004 5-58 No. 4: Mechanical Degradation Source: BSC 2004a, Figure 5-25. 
NOTE: LdB = Lac du Bonnet granite. Tuff tests were conducted at 150C. 
Figure 5-39. Static-Fatigue Curves Used as Input to the UDEC Analyses 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 5-59 Revision 1 
Source: BSC 2004a, Figure 5-28. 
NOTE: Simplified curves are conservative in that complete damage (i.e., failure and loss of strength) is assumed to 
occur once time-to-failure is achieved. 
Figure 5-40. Damage Curves Used as Input to the UDEC Tuff Best-Fit Analyses 
Stress Corrosion Modeling with UDEC Drift Scale ModelThe damagetime response 
illustrated in Figure 5-40 was embedded into the strength of the block contact surfaces of the 
UDEC lithophysal rock model used previously for drift scale stability calculations. Recall that 
the small, irregular-shaped blocks that compose the UDEC lithophysal mass are interconnected 
by shear (cohesive and frictional) and tensile strength components. The strength and stiffness 
parameters of these contacts are calibrated to the lithophysal strength categories. Here, it is 
assumed that the damage coefficients are related directly to the percentage loss of cohesion and 
tensile strength of these contacts; therefore, a damage coefficient of zero is no strength loss and a 
damage coefficient of 1 is 100% strength loss. The details of the generalization of the damage 
criteria within the contact strength logic in the UDEC program can be found in Drift 
Degradation Analysis (BSC 2004a, Appendix S). 
UDEC drift stability analyses were initially conducted for in situ stress only. The time 
increments for solution of drift degradation were 1, 5, 10, 50, 100, 500, 1,000, 5,000, and 
10,000 years. The damage states at 1, 5, and 10 years provide a means of back-analysis against 
existing observations in the ESF and ECRB Cross-Drift. Only the results for the best-fit tuff 
response are given here; additional analyses for the comparative granite time-to-failure data can 
be found in Drift Degradation Analysis (BSC 2004a). 
Time-Dependent Effect from In Situ Stress Only in Lithophysal UnitsResults for the 
predicted drift degradation for the selected time periods for the best fit tuff time-strength 
June 2004 5-60 No. 4: Mechanical Degradation Revision 1 
relations for strength categories 1, 2, 3, and 5 are shown in Figures 5-41 to 5-44. Strength 
Category 1 (the poorest quality tuff) predicts immediate damage in the springline area of the drift 
characterized by extensive breakout which undergoes some, but not significant change over time. 
This means that the stress/strength ratio at the springline area is exceeded from in situ stress 
alone. This response is not observed in the existing ECRB Cross-Drift or ESF tunnels, and, as 
discussed in Chapter 5, the Category 1 strength level, corresponding to the poorest rock 
conditions, represents localized, poor rock conditions and is viewed as very conservative. It is 
noted that the springlines of the ECRB Cross-Drift are, in general, unsupported, and thus ground 
support is not preventing breakouts as predicted by this model. In a similar fashion, Category 2, 
representing about 7% of the Tptpll, shows springline damage levels immediately that are not 
observed in situ. 
As discussed previously, back-analysis of current tunnel observations suggests that strength 
Category 3, which corresponds to approximate lithophysal porosity 15% to 20%, represents 
about 25% of the Tptpll, whereas Category 5 is indicative of the highest quality of lithophysal 
tuff and about 30% of the Tptpll. Little, if any, damage occurs for either of these rock strength 
categories due to in situ stress alone. 
June 2004 5-61 No. 4: Mechanical Degradation Source: BSC 2004a, Figure 5-36. 
Figure 5-41. Predicted Evolution of Damage Due to Strength Degradation for Category 1Tuff Best-Fit 
Static-Fatigue Curve, Applied In Situ Stress Only 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 5-62 Source: BSC 2004a, Figure 5-37. 
Figure 5-42. Predicted Evolution of Damage Due to Strength Degradation for Category 2Tuff Best-Fit 
Static-Fatigue Curve, Applied In Situ Stress Only 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 5-63 Source: BSC 2004a, Figure 5-38. 
Figure 5-43. Predicted Evolution of Damage Due to Strength Degradation for Category 3Tuff Best-Fit 
Static-Fatigue Curve, Applied In Situ Stress Only 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 5-64 Source: BSC 2004a, Figure 5-39. 
Figure 5-44. Predicted Evolution of Damage Due to Strength Degradation for Category 5Tuff Best-Fit 
Static-Fatigue Curve, Applied In Situ Stress Only 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 5-65 Revision 1 
Combined Thermal and Time-Dependent Effects in Lithophysal UnitsThroughout the 
10,000-year regulatory period, the emplacement drifts and surrounding rock mass will be subject 
to a cycle of heating and cooling. The time-dependent strength degradation will be a function of 
the in situ stress concentrations around the drifts as well as the transient, thermally induced stress 
changes. The addition of thermal stresses around the excavation will accelerate the process of 
strength degradation and potential drift instability. The results of numerical simulation of drift 
degradation as a result of these two stress states are shown in Figures 5-45 to 5-47 for Categories 
2,20 3, and 5 rock strength conditions. Time-dependent strength degradation is assessed using the 
tuff best-fit static-fatigue testing data least-squares fit shown in Figure 5-39. 
As expected, the greatest level of in situ, thermal, or time-dependent rockfall occurs for the 
Category 2 rock mass, as shown in Figure 5-45. Initially, most of rockfall comes from the walls, 
which are loaded to a near-yielding state for this rock mass category under in situ stress 
conditions alone. Strength degradation, combined with the increase in rock wall temperature, 
increases the tangential stress component in the walls, resulting in a small amount of rockfall 
from the walls within 5 to 10 years after emplacement of the waste. This localized spalling 
response would, in reality, be contained by ground support. The large increase in the 
temperature, and consequently in the immediate drift wall stresses, after the forced ventilation 
stops causes additional rockfall at the 80 years time frame when peak stress conditions are 
reached. Little additional rockfall occurs after that time as the rock mass slowly cools. 
It might be expected that less rockfall would be predicted in Category 5 (Figure 5-47) than in 
Category 3 (Figure 5-46) due to the higher quality and lower porosity of the Category 5 rock 
mass. However, the greater modulus of the Category 5 lithophysal rock mass causes an increase 
in the tangential compressive stress and greater depth of yielding in the roof, in a fashion similar 
to that observed in the Tptpmn, Drift Scale Heater Test. This phenomenon is predicted to occur 
even for the case when time-dependency is not considered. In any case, Categories 3 and 5 show 
only small amounts of predicted rock degradation due to combined in situ and thermal stressing 
and time-dependent strength property change. 
20 Category 2 is used to represent the lower 10% of the rock mass quality. The response for Category 1 is similar. 
June 2004 5-66 No. 4: Mechanical Degradation Source: BSC 2004a, Figure 5-41. 
Figure 5-45. Predicted Evolution of Damage Due to Strength Degradation for Category 2Tuff Best-Fit 
Static-Fatigue Fit, Combined In Situ and Thermal Stresses 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 5-67 Source: BSC 2004a, Figure 5-42. 
Figure 5-46. Predicted Evolution of Damage Due to Strength Degradation for Category 3Tuff Best-Fit 
Static-Fatigue Curve, Combined In Situ and Thermal Stresses 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 5-68 Source: BSC 2004a, Figure 5-43. 
Figure 5-47. Predicted Evolution of Damage Due to Strength Degradation for Category 5Tuff Best-Fit 
Static-Fatigue Curve, Combined In Situ and Thermal Stresses 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 5-69 Revision 1 
Combined Seismic, Thermal, and Time-Dependent Effects in Lithophysal UnitsThe effect 
of a seismic event characterized by the 10-4 probability of annual recurrence, combined with in 
situ and thermal stressing, was investigated for time-dependent degradation modes for 
Categories 2 and 5 rock masses. The 10-4 seismic loading was chosen for this analysis as it may 
be considered that multiple events of this annual exceedance frequency could occur during the 
postclosure period. This work supplements previous seismic analyses in which thermal load was 
considered, but time-dependent strength loss was not. 
The time frame when the combined in situ and thermally induced stress states reach their peak is 
at approximately 30 years after cessation of forced ventilation, or about 80 years after waste 
emplacement. However, when the time-dependent strength degradation is considered, the state 
when the maximum stresses are generated around the drift is not necessarily the critical state. 
The largest stresses occur relatively early during the regulatory period, and subsequently, the 
stresses decay gradually, returning to the state that existed prior to heating. At the same time the 
strength of the rock mass monotonically decreases as a function of time. 
In order to investigate the extreme effects of the combination of in situ and thermal stress and 
seismic ground motion on drift stability, dynamic analyses were carried out for the model states 
at 80 and 10,000 years after waste emplacement. 
The model geometry before and after the dynamic simulation is shown in Figures 5-48 and 5-49 
for seismic ground motion 80 years after emplacement, and in Figures 5-50 and 5-51 for seismic 
ground motion 10,000 years after emplacement. In all the cases, additional rockfall is predicted 
due to the ground motion, which essentially shakes down the already-damaged rock mass 
resulting from the in situ and thermal stressing. This increase in the rockfall due to the 10-4 
seismic loading is somewhat greater than that predicted for in situ and thermal stressing alone, 
particularly for the Category 2 case. There is also somewhat greater rockfall in the case of an 
earthquake 10,000 years, as opposed to 80 years, after waste emplacement. 
Summary of Time-Dependent Degradation Results: Estimation of Drift Profile for Feeds to 
the Abstraction of Drift SeepageThe analyses presented here have provided an estimate of the 
extent of drift degradation due to in situ and thermally induced stresses as a function of time, 
combined with time-dependent rock mass strength change. Additionally, the effect of 
higher-probability preclosure ground motions on shaking down any damaged and loosened 
material from the degradation process was examined. The analyses show that a spectrum of drift 
damage is predicted ranging from wall breakouts with plan view diameter change of up to about 
2 drift diameters (e.g., Figures 5-48 and 5-50) in the highest porosity rock (Categories 1 and 2: 
about 10% of the Tptpll) to minor breakouts along the walls and crown of the drifts in the 
average to lowest porosity conditions (e.g., Figures 5-49 and 5-51 for Categories 3 to 5: about 
90% of the Tptpll). The modeling predicts that the majority of the damage is due to thermal 
stressing effects and occurs simultaneously with the peak thermal stressing about 30 years after 
ventilation shutdown. Time-dependency results in some small additional damage over time. 
However, the overall conclusion is that the drift profile change and drip shield loading (discussed 
in next section) from nonseismic degradation is minor in comparison to damage from seismic 
stressing. 
June 2004 5-70 No. 4: Mechanical Degradation Revision 1 
The predicted drift profile change as a function of time is required as input for the estimates of 
time-related seepage flux into the emplacement drifts. Among other factors, the seepage flux 
depends on both the size and geometry of the drift and the capillary strength of the fractured rock 
surrounding the drift opening (BSC 2004o, Section 6.4.2.4). In partially or fully collapsed drifts, 
the larger size and potentially different crown shape after collapse will reduce the potential for 
flow diversion compared to the initial drift geometry; furthermore, the larger footprint of the 
collapsed drift leads to an increase in the total amount of percolation flux arriving at the drifts, 
which, in turn, can affect the total amount of seepage. In addition, the capillary-barrier behavior 
at the drift wall can be affected by the rubble rock particles filling the opening, as the capillary 
strength inside the opening will be different from the zero capillarity condition in the 
initially-open drift. Thus, the geometry of the degraded drift and the capillary strength of the 
rubble material inside the drift are of importance in this abstraction. 
Source: BSC 2004a, Figure 5-44. 
NOTE: Displacement magnitudes are in meters. 
Figure 5-48. Effect of 10-4 Ground Motion after 80 Years of Heating in Category 2 before Seismic 
Shaking (Left) and after Seismic Shaking (Right): Blocks Colored by Contours of 
Displacement Magnitude 
June 2004 5-71 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 5-45. 
Figure 5-49. Effect of 10-4 Ground Motion after 80 Years of Heating in Category 5 before Seismic 
Shaking (Left) and after Seismic Shaking (Right): Blocks Colored by Contours of 
Displacement Magnitude 
Source: BSC 2004a, Figure 5-46. 
Figure 5-50. Effect of 10-4 Ground Motion after 10,000 Years of Heating in Category 2 before Seismic 
Shaking (Left) and after Seismic Shaking (Right) Contours of Displacement Magnitude 
June 2004 5-72 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 5-47. 
Figure 5-51. Effect of 10-4 Ground Motion after 10,000 Years of Heating in Category 5, before Seismic 
5.3.2.3 
5.3.2.3.1 Introduction 
The previous calculations of drift degradation from all sources of loading or strength (i.e., in situ 
stress, thermal stress, seismic stressing and shaking, and time-dependent strength degradation) 
have resulted in (1) dynamic loading of the drip shield from falling rock and (2) quasi-static load 
on the drip shield once the fallen rubble has come to rest. The primary concern for dynamic 
impact loading on the drip shield occurs in nonlithophysal rocks due to the larger size of rock 
wedges that can form in these units. Quasi-static loading is a potentially more important issue in 
the lithophysal rocks where potential collapse heights are greater and particle sizes are smaller. 
These factors can lead to larger quasi-static loads applied to the drip shield. 
This section is concerned with prediction of the magnitude and distribution of the quasi-static 
loading. The time-dependency analyses presented in the previous section show that nonseismic 
drift degradation alone is not expected to result in complete tunnel collapse during the regulatory 
period. However, seismic loading and the accompanying rock mass failure can lead to collapse 
of the tunnels for postclosure ground motions of 10-5, 10-6 or 10-7 annual exceedance frequency. 
In these cases, which are highly conservative due to the large and potentially physically 
unrealizable ground motions (BSC 2004c), the rock mass surrounding the drift fails and breaks 
into relatively small particles that fall under gravity to rest beside and on top of the drip shield. 
As shown in Figure 5-31(b), tunnel collapse during large (e.g., PGV greater than about 2 m/sec) 
seismic events is predicted to occur rapidly, within approximately 2 seconds. The analyses 
indicate that the drip shield is rapidly encased by the broken rubble, preventing further large 
translational motions. 
The seismic calculations further provided an estimate of the maximum extent of caving that 
would occur in the sidewalls and roof accompanying these events. The drifts, which become 
Shaking (Left) and after Seismic Shaking (Right): Blocks Colored by Contours of 
Displacement Magnitude 
June 2004 
Quasi-Static Loading to the Drip Shield from Caved Lithophysal Rock 
5-73 No. 4: Mechanical Degradation Revision 1 
filled with rubble, attain an elliptical shape with maximum increase in dimension of about 
1.5 times the original drift diameter. Similar estimates of the maximum collapse height can be 
determined by conducting discontinuum simulations in which the rock mass, under in situ 
loading is assumed to undergo a loss of cohesive (shear) and tensile strength such that it fails to 
its maximum caved height. This height is governed by the in situ stress conditions and the 
bulking of the broken and detached rock that fills the drift and provides a stabilizing back 
pressure to the excavation surfaces. Such an estimate represents a conservative upper bound of 
the load of the caved rock on the drip shield irrespective of the rate of strength decay and the 
residual strength of the rock mass. The quasi-static loads generated from these analyses are used 
as a design basis (along with the dynamic rockfall loads) for the drip shield stability calculations. 
Predictions of ultimate drift collapse height and shape and the subsequent load of the broken 
rock on the drip shield were performed using three different approaches: analytically-based 
approximations, numerical continuum modeling approximations, and numerical discontinuum 
modeling. Three methods, with varying degrees of complexity, were used to provide load 
calculations from alternative conceptual models with the objective of providing an assessment of 
conservatism. Each of the methods uses certain assumptions regarding caving of the rock above 
the drifts and the transfer of the stresses within the broken rock mass. As discussed below, the 
level of conservatism is the greatest in the analytically-based approach and the smallest in the 
approach that represents the rock mass as a discontinuum. The analytical and numerical 
continuum models require assumption of the ultimate shape of the collapsed zone and the 
bulking factor of the rubble and do not take into account rock mass properties or in situ stresses. 
Load on the drip shield is estimated simply from the height of broken rock above it. 
The discontinuum numerical modeling technique is judged to most closely represent the actual 
caving process. The only assumption made in this method is the size and shape of rock particles 
produced as the rock mass fails. The ultimate shape and size of the collapsed drift and the 
bulking factor of the caved rock are determined from the model. Load on the drip shield is 
determined directly from the interaction of the rubble and the drip shield. The drip shield is 
modeled using a two-dimensional representation in which the deformability and shape are 
correctly represented. The invert is modeled as a rigid boundary and the footings of the drip 
shield are free to slide or separate as forces dictate. The analyses here are aimed at determining 
the drip shield loading provided by the broken rock, and not to model the stability or perform 
design calculations of the drip shield. However, to provide an adequate determination of the 
load, the interaction of the broken rock and the structure needs to be modeled. These analyses 
provide a model of this interaction and the resulting stresses for the case of complete collapse. 
The resulting applied stresses are fed to the analysis of the drip shield in which the geometry and 
yield properties of the structure is modeled in detail. 
June 2004 5-74 No. 4: Mechanical Degradation Revision 1 
5.3.2.3.2 
(Eq. 5-4) 
Bulking of the Collapsed Rock Mass 
When the rock mass above underground openings collapses it increases volume (i.e., it bulks). 
During the collapse, either sudden or gradual, the rock mass breaks into a number of pieces 
(blocks) that fall separately, and rotate as it falls and contacts the muck pile. When blocks 
equilibrate after caving, they do not fit together resulting in increased porosity and overall 
volume. The rock mass of volume V in the in situ conditions has volume VB after caving, 
where: 
VB = (1 + B)V 
where B is termed the bulking factor. 
The amount of bulking (i.e., the bulking factor, B) depends, among other things, on the lithology, 
preexisting internal structure (lithophysae, jointing, bedding), and the mechanism of collapse. 
For example, the density of crushed limestone is in the range between 1,360 kg/m3 and 
1,440 kg/m3; while density of crushed dolomite is in the range between 1,280 kg/m3 and 
1,600 kg/m3 (Fruchtbaum 1988). Considering that the specific gravity of limestone and dolomite 
is approximately 2.6 (Bauer et al. 1991), and using an in situ porosity of 20% (Goodman 1980), 
the in situ density of limestone and dolomite is approximately 2,200 kg/m3. Consequently 
bulking of these rocks from in situ state to a crushed state is between 37.5% and 72%. Duncan 
et al. (1980) reported that porosity of rockfill for dams is between 23% and 36%. The rockfill 
used for dams is crushed to satisfy a certain size requirement and is compacted during 
construction, which leads to reduction of its porosity. It appears from this discussion that the 
bulking factor for caved rock can be conservatively selected to be in the range between 0.2 and 
0.4, even for lithophysal rock, which has initial porosities ranging from about 20% to 35%. 
Caving of the underground excavations is a self-limiting process in situations where the 
excavation diameter is much less than the depth below ground surface (i.e., where the depth is 
approximately 3 times or more of the tunnel radius, the depth is over 100 times the radius at 
Yucca Mountain). At a certain stage of caving, due to bulking, the volume of the caved rock 
completely fills the volume of the original excavation and the volume occupied by the collapsed 
rock before onset of collapse. When the cave is completely filled, the broken rock provides a 
backpressure, which prevents further collapse of the rock mass. This self-limiting concept of 
complete filling is a simplistic and conservative mechanism that provides the basis for the 
analytical approaches below. 
In addition to the potential for complete filling as a cave-stabilizing mechanism, caving often 
results in formation of elliptical-shaped excavations, which, in turn, results in tangential stress 
concentrations at the cave crown where the radius of curvature is the greatest. These higher 
stresses can confine the rock, arresting cave development without complete filling of the tunnel. 
This is the case in many caves or other natural excavations that may form naturally arched, stable 
excavations. Analytical approaches do not examine the impact of in situ stress conditions, but 
this is accounted for when using the discontinuum modeling approaches that are used in the 
following analysis. 
5-75 June 2004 No. 4: Mechanical Degradation Revision 1 
Analytical Models of Collapse Height and ShapeIt is considered in this approach that the 
cave above the emplacement drift grows until it becomes filled with the broken rock, irrespective 
of the rock strength or rock mass stress state. The extent of the caved rock is calculated as a 
function of the bulking factor, B, considering that the cave stabilizes when it is completely filled 
with the broken rock. An additional unknown in this approach requiring assumption is the shape 
of the cave. In all cases, the entire dead weight of the material vertically overlying the drip 
shield is assumed to act on the drip shield. The frictional forces acting between the caved 
material and the stable sidewalls and within the caved particles themselves, is not taken into 
account, resulting in a lack of any portion of the vertical load being transferred to the 
surrounding rock mass. Two extreme conditions illustrated in Figures 5-52 and 5-53 and were 
considered. 
The piping mode of roof collapse (Figure 5-52) may occur for conditions where the rock mass 
is horizontally bedded and there is a relatively large ratio of the span of the excavation to its 
depth. This type of roof collapse is sometimes induced to occur in coal mines (with a 
thinly-bedded overburden), using the longwall mining method, and often occurs suddenly. Roof 
piping collapse is not a likely mode of drift collapse at Yucca Mountain for the following 
reasons: 
 None of the rock mass units are thinly stratified 
 The ratio of depth to drift diameter is very large, and therefore the rock mass may be 
considered infinite in extent from a mechanical standpoint. 
 Drift collapse due to time-dependent strength decay will evolve gradually over a long 
period of time. 
The piping mechanism is considered here as a conservative extreme condition and is a 
mechanism that results in the largest vertical extent of the cave, H. 
The other extreme conservative condition of the rock mass collapse around the underground 
opening (shown in Figure 5-53) corresponds to the limit equilibrium conditions around a shallow 
tunnel, which Terzaghi (1943) used to calculate the load on the tunnel support. Slip lines extend 
from the drift walls at an angle of 45  /2 from the vertical direction, where  is the friction 
angle. 
June 2004 5-76 No. 4: Mechanical Degradation Source: BSC 2004a, Figure 6-156. 
Figure 5-52. Piping Type of Caving Mechanism 
Source: BSC 2004a, Figure 6-157. 
Figure 5-53. Terzaghi-Type Caving Mechanism 
No. 4: Mechanical Degradation 
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June 2004 5-77 Revision 1 
In both analytical representations, the cave height, and thus the dead weight of material that lies 
on the drip shield, is a function of the assumed bulking factor (BSC 2004a). 
Numerical Continuum Approach to Simulate Piping and Terzaghi-Type MechanismsThe 
next degree of greater detail in estimating cave formation and rubble loading is to use numerical 
modeling approaches. The simplest numerical model is one in which the rock mass and rubble 
are considered to be continuum materials. The analytical approaches do not take into account the 
effect of rock mass stress conditions and material properties. The use of numerical methods 
allows testing of the proposed piping and Terzaghi assumptions when the rock mass stresses and 
material properties are taken into account. The numerical models can also be used to examine 
the importance of potential stress arching within the broken rock that lies within the drift. 
However, continuum-based analyses also represent conservative, but more accurate, simulations 
of the collapse height than either the piping or Terzaghi-type mechanisms. A simplistic 
methodology was used in which the rock mass was represented as a Mohr-Coulomb material 
within FLAC, a continuum numerical code. 
The estimated strength of the lowest quality lithophysal rock mass was used in the initial drift 
equilibrium state at excavation (Category 1). Subsequently, cohesion and tensile strength of the 
rock mass were reduced gradually, in steps to induce collapse. At each stage of strength 
reduction the model was run until either equilibrium was achieved, or there was clear indication 
that equilibrium could not be achieved (i.e., the rock mass around the drift was collapsing). 
Once the collapse was detected the model simulation was interrupted, and the cave height was 
calculated based on the bulking factor and the volume of the rock mass within the destabilized 
region. An example of the numerical prediction of collapse for the Terzaghi mechanism is 
shown in Figure 5-54. 
It is noted that the bulking factor required for this analysis controls the ultimate cave propagation 
and is an uncertain parameter (as is the ultimate shape of the cave zone). Again, two limiting 
mechanisms were considered and generated using the numerical model: (1) the piping 
mechanism (where the caved rubble is assigned zero cohesive strength), in which the cave width 
was limited to the drift width and (2) the Terzaghi mechanism, in which cave width coincides 
with the width of the destabilized region of the rock mass under in situ stress conditions. 
Subsequently the drift and the caved region were filled with zones (caved rock selected to have 
no cohesive or tensile strength, and density accounting for the bulking factor), and the model was 
run to the equilibrium state to determine the load on the drip shield. 
June 2004 5-78 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 6-160. 
Figure 5-54. Failure Mechanism of a Deep Tunnel in Cohesionless Material 
Discontinuum Numerical Approach to Estimating Drip Shield LoadThe most detailed and 
representative approach to representing the actual caving behavior of a rock mass is the use of 
discontinuum numerical modeling. The collapse problem was simulated using the UDEC 
program to provide the most realistic prediction of caving by including actual site-specific 
material strength and in situ stress conditions. After initial stress equilibrium of the excavation 
was achieved for a circular 5.5 m diameter drift, the cohesion and tensile strength of the rock 
mass were subsequently reduced in a series of steps to provide the ultimate collapse height and 
shape of the excavation when the rubble eventually provides sufficient back pressure to the 
excavation to suppress further failure. The loading on the sides and top of the drip shield is then 
determined by the model. 
The cohesive and tensile strengths are reduced in 5 increments from full strength properties (for 
a given rock strength category) to zero, allowing the rock to fail along its internal fracture 
boundaries, and to fall by gravity onto the invert and drip shield. In this model, the bulking of 
the caved rock is not an assumed model parameter but is the calculated result of the modeling. 
The bulking of the rubble depends on the size and shape of the falling blocks, which are 
predetermined by the size and the shape of the UDEC blocks in the model. The UDEC analyses 
use irregular-shaped blocks with a mean side length of 0.3 m. To ensure that the model 
estimates produce values of bulking factor that are considered to be low from common practice, 
additional analyses were performed with a 0.2 m block size so that the resulting bulking factor 
was equal or less than 0.2, the lower bound of the bulking factor expected in rocks gained from 
construction experience (described previously). 
June 2004 5-79 No. 4: Mechanical Degradation Revision 1 
After the model has come to equilibrium under complete collapse, the broken rock blocks will 
rest on the drip shield and the invert of the drift (Figures 5-55 to 5-57). Figure 5-55 shows the 
geometry of the fully collapsed tunnel and rockfall in contact with the deformable drip shield. 
The contact of the footings with the drip shield is free to slide or separate on the invert. The 
exertion of the side-loading from rockfall results in lateral translation of the drip shield and more 
uniform loading distributions on the sides, with larger vertical loading to the crown. The 
ultimate collapsed height and shape of the cave is illustrated in Figure 5-56, which shows 
contours of displacement of the rock mass and rubble arising from the yield and collapse process. 
The ultimate collapse zone grows to approximately twice the original diameter of the tunnel. 
The rubble attains a bulking factor of approximately 19% in the case of 0.2 m block size. A total 
of 6 realizations of block shape (all assuming an average dimension of 0.2 m) were simulated, 
resulting in a range of potential loading conditions on the drip shield. The drip shield is 
subdivided into 30 segments, and average loads on each segment are defined. Figure 5-57 shows 
a histogram of the average pressure on each of the 30 elements for all six realizations. The mean 
of all the realizations is also plotted. As seen in this plot, the average loads from all realizations 
on the right and left hand sides are approximately the same, while the average pressure on the top 
of the drip shield is about 50% higher. Locally larger or lower point loads may exist, depending 
on the how the rubble falls and compacts around the drip shield. 
5.3.2.4 Comparison of the Quasi-Static Drip Shield Loading from Various Methods 
The predictions of average vertical pressure of the caved rock exerted on the top of the drip 
shield by all three approaches are summarized in Figure 5-58. This figure shows the pressure as 
a function of bulking factor. In the case of the analytical and continuum numerical model, the 
bulking factor is an assumed parameter. In the case of the discontinuum modeling, the bulking 
factor is determined directly from the results of the modeling and is the reason why the results 
are concentrated around a bulking factor of approximately 0.2. Results for the discontinuum 
modeling are given for assumptions of a deformable drip shield, as well as for a rectangular, 
rigid drip shield as a means of examining the impact of drip shield shape and deformability on 
stress arching in the rubble. 
As expected, the analytical models yield the largest loads resulting from the overly conservative 
assumptions inherent in each approach. The continuum numerical model accounts more 
accurately for transfer of load by friction from the caved rock to the surrounding stable rock 
mass. Consequently predicted loads for small bulking factors and large cavity size are much 
smaller than analytical predictions. When the bulking factor is large, the height of the cave 
becomes small. Stress arching cannot be realized within the small column of the cave rock, and 
consequently the predictions from the analytical and continuum models are identical. The most 
realistic approach, using the discontinuum model, does not use an imposed (assumed) condition 
about the shape or extent of the caved region. It also accounts for load transfer through the 
caved rock in a reasonable fashion via frictional contact between rock blocks. A complete 
discussion of the various analyses leading to estimates of drip shield pressure distributions can be 
found in Drift Degradation Analysis (BSC 2004a). The predictions of the pressures on the drip 
shield using this approach are smaller than the predictions of the analytical and continuum 
models for all values of the bulking factor, but diverge significantly for bulking factors below 
approximately 0.2. 
June 2004 5-80 No. 4: Mechanical Degradation Revision 1 
A number of analyses were also run to test the impact on quasi-static load due to seismic shaking 
on a previously collapsed drift. Figure 5-58 also shows the final loading on the drip shield when 
a previously collapsed drift (e.g., Figure 5-55) is subjected to shaking from subsequent ground 
motions with annual exceedance frequencies of 10-4, 10-5, and 10-6. As seen, the effect of the 
additional seismic shaking on the loose rubble (approximately 20% initial bulking factor) is to 
cause additional compaction of the rubble, with a reduction in bulking factor. For example, the 
10-4 motion has little effect on rubble compaction; however, the bulking factor decreases to 
approximately 14% and 10% for the 10-5 and 10-6 motions, respectively. However, the 
increased compaction does not lead to dramatically increased loads on the drip shield due to 
arching in the rubble. 
Source: BSC 2004a, Figure 6-164 
NOTE: Stresses are given in Pascal. 
Figure 5-55. Quasi-Static Drift Degradation, 0.2 m Block Size: Equilibrium State for Deformable Drip 
Shield with Arched Top, Footings Free to Slide or Detach from the Invert 
June 2004 5-81 No. 4: Mechanical Degradation Source: BSC 2004a, Figure 6-165. 
Figure 5-56. Quasi-Static Drift Degradation, 0.2 m Block Size: Contours of Displacement Magnitude for 
Deformable Drip Shield Showing Approximate Collapse Height of 7 m (about 6 m above 
Top of Drip Shield Crown) 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 5-82 Revision 1 
Source: BSC 2004a, Figure 6-166. 
NOTE: Average pressure on each segment is shown for all six realizations. Segment numbering starts at 1 at right 
footing and continues counterclockwise to the left footing. Those elements on the right, top, and left sides of 
the drip shield are shown. 
Figure 5-57. Quasi-Static Pressure on Drip Shield Segments for Six Realizations for Random, 0.2 m 
Block Geometries 
June 2004 5-83 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 6-170. 
NOTE: Solid lines show the analytical methods. Symbols are numerical solutions. 
Figure 5-58. Summary of Average Vertical Pressure on the Drip Shield as a Function of Bulking Factor 
for Analytical and Numerical Analysis 
5.3.2.5 Rubble Load Feeds to Engineering Structural Analysis of the Drip Shield 
The rubble loading estimates described above are ultimately used as input to the structural design 
analysis of the drip shield and are not part of this technical basis document. Three-dimensional 
finite element structural analyses are performed to examine stress and deflection of the drip 
shield and, in particular, buckling modes or excessive deformation that may result in contact with 
the waste package. The estimated rubble quasi-static loads derived from discontinuum modeling 
as depicted in Figure 5-57 are used as direct input to the structural analysis. 
5.4 SUMMARY OF OUTPUT FROM THE DRIFT DEGRADATION ANALYSIS TO 
OTHER DISCIPLINES 
The results of the drift degradation modeling provide feeds to several engineering calculations 
and abstraction models that eventually feed the TSPA model. This information is summarized 
here: 
 Preclosure Analysis of Rockfall and Waste Package BreachAn analysis of potential 
rockfall due to preclosure in situ, thermal, and seismic loading was conducted for 
June 2004 5-84 No. 4: Mechanical Degradation Revision 1 
unsupported21 emplacement drifts and other repository excavations (BSC 2004k). This 
work provides an estimate of the maximum credible rock masses during the preclosure 
period and is used to show that damage of a waste package due to rockfall for 
unsupported openings is not a credible event. 
 Engineering Calculation of Drip Shield Stability and Damage from RockfallThis 
includes dynamic and quasi-static loading. 
- Dynamic LoadingNonlithophysal rockfall masses, relative velocity components, 
impact location, and rockfall shapes are supplied to three-dimensional engineering 
structural analysis of the drip shield for each set of analyses at 10-5, 10-6, and 10-7 
annual exceedance frequencies. In lithophysal rock, where particle sizes are 
relatively small, the dynamic stress time-history on each segment of the deformable 
drip shield is supplied to three-dimensional engineering structural analysis of the 
drip shield. Estimates of drip shield stability and yielded surface area are determined 
by the calculations in Drip Shield Structural Response to Rock Fall (BSC 2004m). 
- Quasi-Static LoadingA set of realizations of drift collapse were derived from 
discontinuum modeling. These realizations allowed complete collapse of the drift 
with filling of rubble until further collapse was arrested due to the back pressure to 
the excavation from the rubble. The stresses applied by the rubble to individual 
elements comprising the deformable drip shield at equilibrium were supplied to 
engineering for three-dimensional structural analysis of the drip shield. The 
engineering analysis examines buckling and deformation potential of the drip shield. 
The time-relation of the collapse does not enter into these analyses because the 
conservative condition of complete collapse of the excavations is being taken into 
account in the drip shield structural design. 
 Drift SeepageThis includes the drift profile and drift periphery damage assessment. 
- Drift ProfileThe cross-sectional profile of the emplacement drifts impacts the 
amount of seepage water that enters the tunnels. Drift collapse profiles for the range 
of rock strength categories for non-seismic and seismic loading conditions are fed to 
the drift seepage abstraction (BSC 2004o) for estimation of seepage inflow to the 
drifts. The drift is assumed to be filled with rubble, which is taken into account in 
the determination of the capillary barrier (as opposed to an empty drift). 
- Drift Periphery Damage AssessmentThe rock mass strain and estimated change 
in fracture aperture around the periphery of the drift and the bulking factor and 
porosity of the collapsed rubble are provided for use in analysis of the drift capillary 
barrier for a collapsed drift. 
21 The excavations are assumed to be unsupported to support the case that ground support during preclosure is not 
important to safety. 
June 2004 5-85 No. 4: Mechanical Degradation Revision 1 
 Drift EnvironmentThis includes the drift collapse. 
- Drift CollapseDrift collapse could have an impact on heat transfer from the waste 
package to the surrounding rock mass and, thus, affect waste package and drip shield 
temperature. This is particularly true if significant degradation were to occur during 
the time period in which temperatures are in excess of the boiling point (i.e., about 
1,000 years). Time-dependent analyses show that, in the absence of seismic events 
with peak ground velocity greater than about 2 m/s, the drifts in either lithophysal or 
nonlithophysal rock remain largely intact during this time period. 
June 2004 5-86 No. 4: Mechanical Degradation Revision 1 
6. SUMMARY AND INTERACTIONS WITH ENGINEERED AND 
NATURAL SYSTEMS 
This document presented a description of the postclosure performance issues involving 
mechanical degradation of excavations. 
6-1 
6.1 SUMMARY 
This report reviews the geologic, laboratory, field, and numerical analysis work that has been 
performed to document the degradation of the rock mass surrounding the emplacement drifts of a 
geologic repository at Yucca Mountain. The factors leading to drift degradation include the 
stresses from overburden (in situ) stresses, stresses induced by the heat released by the emplaced 
waste, the stresses due to seismically related ground motion, and the time-dependent strength 
degradation. 
The strategy for resolution of these issues was described, and the studies performed to support 
this resolution were discussed in detail. The strategy was originally presented in March 2002 
and documented by Board (2003) and reviewed in DOE-NRC Appendix 7 and Technical 
Exchange (April 2003 (Stablein and Gil 2003)) meetings. The studies included: (1) performance 
of detailed geologic and geotechnical mapping studies aimed at describing those geologic 
features relevant to geomechanical behavior, (2) development of data of mechanical and thermal 
rock properties for the primary repository host horizons, and (3) development and calibration of 
material and numerical models for prediction of rock mass response to loading. 
These factors have been modeled and analyzed, resulting in the prediction of the general stability 
of the emplacement drifts, and, specifically, the amount and size distribution of rockfall in the 
repository drifts during both the preclosure and postclosure regulatory periods. 
The following statements summarize the results from this drift degradation modeling and 
analysis activity and present the key conclusions: 
 The rock mass at the repository host horizon has been geologically characterized to 
support the stability and rockfall modeling activities presented in this report. Drift 
degradation models have been developed for both nonlithophysal and lithophysal rock. 
A detailed description of the rock mass characteristics of the repository host horizon was 
performed. The available rock mass geotechnical data, including fracture geometry, 
lithophysal abundance and geometric characteristics, and geotechnical rock properties, 
are sufficient to support detailed drift degradation analyses using both continuum and 
discontinuum approaches. 
 The drift-scale temperature history was calculated throughout the preclosure and 
postclosure periods of the repository. The temperature history was used to calculate the 
thermal stress state that develops within the rock mass due to the heat energy released 
from the stored nuclear waste, and appropriate thermal properties and boundary 
conditions for thermal loading have been applied. Three-dimensional, mountain-scale 
models of heat transfer in the rock have been conducted to determine topographical, 
repository shape, and drift location effects. 
June 2004 No. 4: Mechanical Degradation Revision 1 
 A nonlithophysal rockfall model was developed using the three-dimensional 
discontinuum code, 3DEC, with the following features: 
- Appropriate boundary conditions are provided for thermal and seismic loading. 
- Critical fracture patterns are included from multiple sampling from a synthetic rock 
mass volume that contains a realistic fracture population based on field mapping data. 
- Appropriate thermal and mechanical properties of rock blocks and joints are used. 
- Long-term degradation of joint strength parameters is considered. 
- Site-specific ground motion time histories appropriate for both the preclosure 
(5  10-4, 10-4 annual exceedance frequencies) and postclosure (10-5, 10-6, and 10-7 
annual exceedance frequencies) time periods are included in the model. 
 A lithophysal rockfall model was developed using the two-dimensional discontinuum 
code, UDEC, with the following features: 
- Appropriate thermal and mechanical properties of the lithophysal rock have been 
determined through a combination of large diameter laboratory testing, field-scale 
testing, and extrapolation using calibrated numerical models. The model 
extrapolation is used to examine the variability of mechanical properties with 
lithophysal porosity, shape, size, and distribution. The rock mechanical property 
range from testing has been subdivided into five quality categories that are 
functions of rock mass porosity. Categories 1 and 2 represent the lower end of the 
observed rock quality within the lithophysal rock, whereas Categories 3 and 4 
represent an approximate average condition. Category 5 represents the higher 
quality, lower porosity portions of the Tptpll. These categories provide a basis for 
parametric calculations. A material model that accounts for the lithophysal rock mass 
behavior is developed. The majority of analyses were conducted for these bounding 
ranges of properties; however, a lithophysal porosity spatial variability model was 
developed to examine directly the impact of in situ variability on seismic and 
time-dependent drift stability. 
- Appropriate boundary conditions are provided for thermal and seismic loading. 
- The rock mass is represented as an assembly of polygonal, elastic blocks in which the 
bond strength of the blocks is calibrated such that the overall mechanical behavior of 
the mass is consistent with the material model developed for the lithophysal rock. 
- The discontinuum lithophysal rockfall model allows for the formation of 
stress-induced fractures between blocks (i.e., the formation of internal fracturing) 
separation and instability (under the action of gravity or seismic shaking) of the rock 
mass around the drift. 
- The effect of long-term degradation of rock mass strength is considered, assuming a 
stress corrosion mechanism that is dependent on time and applied stress level. An 
June 2004 6-2 No. 4: Mechanical Degradation Revision 1 
timate of the degradation of emplacement drifts was developed by incorporating a 
time-related rock mass shear and tensile strength reduction factor into the drift 
degradation numerical model. 
- Site-specific ground motion time histories appropriate for both the preclosure 
(5  10-4, 10-4 annual exceedance frequencies) and postclosure (10-5 and 10-6 annual 
exceedance frequencies) time periods are included in the model. 
 Model validation activities include: (1) validating the mechanical material models or 
representations for the two specific repository host rock types (i.e., lithophysal and 
nonlithophysal rocks), and (2) validating the implementation of these material models in 
general numerical modeling schemes. 
 The results for the nonlithophysal units are summarized as follows: 
- Preclosure ground motion results in minor drift damage due to rockfall. 
- Postclosure ground motion results in a range of drift damage as a result of wedge-type 
rockfall (i.e., controlled by geologic structure). The mean block size is less than 
approximately 0.2 MT for all preclosure and postclosure motions. The large and 
highly conservative postclosure ground motions (particularly at the 10-7 level) can 
result, for some cases, in spalling of the intact rock blocks surrounding the 
excavations. The great conservatism in this result is questionable, as these large 
amplitude motions may not be physically realizable. 
- Thermal effects have a minor impact on rockfall. 
- Time-dependent strength degradation has a minor impact on rockfall. 
 The results for the lithophysal units are summarized as follows: 
- Degradation is primarily controlled by stress conditions. 
- Rock block size produced by degradation is controlled by the intense fracturing due 
to natural cooling of the lithophysal rock matrix as well as the spacing of lithophysae, 
and is unrelated to the more widely spaced, longer cooling fractures. The spacing of 
fractures in the lithophysal rock averages approximately 0.05 m, with a resulting 
small block side length when the rock is overstressed. 
- Preclosure ground motion results in minor drift damage due to rock failure. 
- The 10-5 ground motion results in a range of response from no damage to collapse 
based on the highly variable amplitude of the individual time histories. A reasonable 
relationship between damage level and peak ground velocity or kinetic energy in the 
motion has been established. General collapse of the drifts occurs for peak ground 
velocity greater than about 2 m/s. The highly conservative 10-6 postclosure ground 
motion result in predicted collapse of the drift, with fragmented rock particle sizes on 
June 2004 6-3 No. 4: Mechanical Degradation Revision 1 
the order of centimeters to decimeters. It is questionable whether these conservative 
motions are physically realizable. 
- Thermal and time-dependent effects alone are expected to result in relatively small 
amounts of rockfall. Only in the poorest quality of rock are significant 
time-dependent breakouts expected. Time-dependent degradation is not expected to 
result in total collapse of the emplacement drifts. Nonetheless, a conservative 
approach to drip shield quasi-static loading has been assumed for drip shield 
structural design. The discontinuum model has been used to simulate complete 
collapse of the tunnels, and to estimate the subsequent vertical and lateral loads 
applied to the drip shield. These loads provide input loading conditions for drip 
shield structural analysis and design. 
- Drift profiles based on complete collapse and rubble-filled drifts are used as input to 
the abstraction of drift seepage. These collapsed drift profiles are assumed at 
repository closure. 
The drift degradation models and analyses documented in this report address the requirements of 
NRC/DOE agreements regarding rockfall and related issues to support the resolution of NRC 
KTI on Repository Design and Thermal-Mechanical Effects. 
6.2 CLOSURE 
The drift degradation models have adequately captured the physical phenomena associated with 
the various components of rock mass behavior anticipated within the repository horizon. 
Appropriate boundary and initial conditions have been applied to the models, and the technical 
bases for the development of these rockfall models have been adequately documented. 
Sufficient data have been collected to adequately model the drift degradation processes. The 
technical bases and ranges of data used in the rockfall models are documented. 
Data uncertainty has been characterized through parameter sensitivity studies in the rockfall 
models. Model uncertainty has been characterized through an evaluation of alternative 
conceptual models, and the model results have been validated by comparison to field and 
laboratory data, alternative numerical approaches, and industry experience through external 
technical review. 
The most significant uncertainties impacting the results of the rockfall models are those 
associated with the postclosure ground motion and time-dependent degradation. Some of the 
ground motion data provided are larger than the largest ground motion observed and may not be 
physically realizable. Therefore, predictions of complete drift collapse with postclosure ground 
motion may be unrealistic. Currently lacking a technical basis to limit such ground motion to 
smaller values, these inputs represent the best available information to support this work. 
Prediction of the time-dependent degradation for the duration of the regulatory period of 
10,000 years is a highly approximate task. Empirical data to provide evidence for tunnel 
collapse over hundreds to thousands of years are not available. A mechanics-based approach is 
used in which a stress corrosion crack model, based on static fatigue test data, is used for 
prediction of long-term stability of excavations. A small data base of time-dependent testing of 
June 2004 6-4 No. 4: Mechanical Degradation Revision 1 
tuff is currently available to support this work. A significant program in static-fatigue 
measurement of Tptpmn cores is being conducted. This present work will be updated as the 
long-term testing data are developed. 
The drift degradation and modeling activities summarized in this report are sufficient to support 
a license application. The data on rockfall size and amount are sufficient to provide input to drip 
shield design calculations, consequence models for the seismic scenario for the total system 
performance assessment for the license application, and seepage abstraction models for the 
nominal scenario for the total system performance assessment for the license application. 
June 2004 6-5 No. 4: Mechanical Degradation INTENTIONALLY LEFT BLANK 
6-6 No. 4: Mechanical Degradation 
Revision 1 
June 2004 Revision 1 
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7.2 CODES, STANDARDS, AND REGULATIONS 
10 CFR 63. Energy: Disposal of High-Level Radioactive Wastes in a Geologic Repository at 
Yucca Mountain, Nevada. Readily available. 
7.3 DATA, LISTED BY DATA TRACKING NUMBER 
GS000608314224.004. Provisional Results: Geotechnical Data for Station 35+00 to Station 
40+00, Main Drift of the ESF. Submittal date: 06/20/2000. 
GS000608314224.006. Provisional Results: Geotechnical Data for Station 26+00 to 30+00, 
North Ramp and Main Drift of the ESF, Full-Periphery Geotechnical Maps (Drawings OA-46- 
222 through OA-46-226) and Rock Mass Quality Ratings Report. Submittal date: 06/28/2000. 
GS030283114222.001. Direct Shear Data from Selected Samples of the Topopah Spring Tuff. 
Submittal date: 02/20/2003. 
GS960708314224.008. Provisional Results: Geotechnical Data for Station 30 + 00 to Station 35 
+ 00, Main Drift of the ESF. Submittal date: 08/05/1996. 
GS960708314224.010. Provisional Results: Geotechnical Data for Station 40+00 to Station 
45+00, Main Drift of the ESF. Submittal date: 08/05/1996. 
7-7 June 2004 No. 4: Mechanical Degradation Revision 1 
GS960908314224.015. Provisional Results: Geotechnical Data for Stations 30+00 to 40+00, 
Main Drift of the ESF, Full-Periphery Geotechnical Maps and Rock Mass Quality Ratings 
Report. Submittal date: 09/09/1996. 
GS960908314224.016. Provisional Results: Geotechnical Data for Station 40+00 to 50+00, 
Main Drift of the ESF, Full-Periphery Geotechnical Maps and Rock Mass Quality Ratings 
Report. Submittal date: 09/09/1996. 
GS960908314224.020. Analysis Report: Geology of the North Ramp - Stations 4+00 to 28+00 
and Data: Detailed Line Survey and Full-Periphery Geotechnical Map - Alcoves 3 (UPCA) and 4 
(LPCA), and Comparative Geologic Cross Section - Stations 0+60 to 28+00. Submittal date: 
09/09/1996. 
GS971108314224.025. Revision 1 of Detailed Line Survey Data, Station 26+00 to Station 
30+00, North Ramp and Main Drift, Exploratory Studies Facility. Submittal date: 12/03/1997. 
GS990408314224.004. Full-Periphery Geologic Maps for Station 10+00 to 15+00, ECRB Cross 
Drift. Submittal date: 09/09/1999. 
LL980411104244.061. DST Baseline REKA Probe Measurements for Thermal Conductivity 
and Diffusivity. Submittal date: 04/24/1998. 
LL980902104244.070. DST Baseline REKA Probe Measurements for Thermal Conductivity 
and Diffusivity. Submittal date: 09/03/1998. 
MO0001SEPDSTPC.000. Drift Scale Test (DST) Temperature, Power, Current, and Voltage 
Data for June 1, 1999 through October 31, 1999. Submittal date: 01/12/2000. 
MO0002ABBLSLDS.000. As-Built Borehole Locations and Sensor Locations for the Drift 
Scale Test Given in Local (DST) Coordinates. Submittal date: 02/01/2000. 
MO0007SEPDSTPC.001. Drift Scale Test (DST) Temperature, Power, Current, and Voltage 
Data for November 1, 1999 through May 31, 2000. Submittal date: 07/13/2000. 
MO0107SEPDSTPC.003. Drift Scale Test (DST) Temperature, Power, Current, and Voltage 
Data for December 1, 2000 through May 31, 2001. Submittal date: 07/06/2001. 
MO0202SEPDSTTV.001. Drift Scale Test (DST) Temperature, Power, Current, and Voltage 
Data for June 1, 2001 through January 14, 2002. Submittal date: 02/28/2002. 
MO0211TMHIS104.002. Acceleration, Velocity, and Displacement Time Histories for the 
Repository Level at 5X10-4 Annual Exceedance Frequency. Submittal date: 11/14/2002. 
MO0301SPASIP27.004. Sampling of Stochastic Input Parameters for Rockfall Calculations and 
for Structural Response Calculations Under Vibratory Ground Motions. Submittal date: 
01/15/2003. 
June 2004 7-8 No. 4: Mechanical Degradation Revision 1 
MO0301TMHIS106.001. Acceleration, Velocity, and Displacement Time Histories for the 
Repository Level at 10-6 Annual Exceedance Frequency. Submittal date: 01/28/2003. 
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. 
MO0402AVDTM105.001. Acceleration, Velocity, and Displacement Time Histories for the 
Repository Level at 10-5 Annual Exceedance Frequency. Submittal date: 02/09/2004. 
MO0403AVTMH107.003. Acceleration, Velocity, and Displacement Time Histories for the 
Repository Level at 10-7 Annual Exceedance Frequency. Submittal date: 03/22/2004. 
MO9807DSTSET01.000. Drift Scale Test (DST) Temperature, Power, Current, Voltage Data 
for November 7, 1997 through May 31, 1998. Submittal date: 07/09/1998. 
MO9906DSTSET03.000. Drift Scale Test (DST) Temperature, Power, Current, and Voltage 
Data for September 1, 1998 through May 31, 1999. Submittal date: 06/08/1999. 
SN0206F3504502.012. Revised Thermal Conductivity, Volumetric Heat Capacity and Thermal 
Diffusivity Data for ECRB Thermal K Test 1 (Two-Hole Test). Submittal date: 06/07/2002. 
SN0206F3504502.013. Revised Thermal Conductivity, Volumetric Heat Capacity and Thermal 
Diffusivity Data for ECRB Thermal K Test 3 (Three-Hole Test, with Results from 1/22/2002 
through 4/9/2002). Submittal date: 06/07/2002. 
SN0208F3504502.019. Thermal Conductivity, Volumetric Heat Capacity and Thermal 
Diffusivity Data for ECRB Thermal K Test 2 (Six-Hole Test). Submittal date: 08/30/2002. 
SN0208L01B8102.001. Thermal Expansion Properties of Lithophysal Tuff, Batch #1 (Test 
Dates: August 3, 2002 through August 16, 2002). Submittal date: 08/28/2002. 
SN0208L0207502.001. Mechanical Properties of Lithophysal Tuff, Batch #1 (Test Dates: July 
31, 2002 through August 16, 2002). Submittal date: 08/20/2002. 
SN0208T0503102.007. Thermal Conductivity of the Potential Repository Horizon Rev 3. 
Submittal date: 08/26/2002. 
SN0211L01B8102.002. Thermal Expansion Properties of Lithophysal Tuff, Batch #2 (Test 
Dates: October 20, 2002 through October 25, 2002). Submittal date: 11/13/2002. 
SN0211L0207502.002. Mechanical Properties of Lithophysal Tuff, Batch #2 (Test Dates: 
October 22, 2002 through October 25, 2002). Submittal date: 11/13/2002. 
SN0305L0207502.006. Porosity of Laboratory Mechanical Properties Test Specimens for Batch 
#1 and Batch #2. Submittal date: 05/20/2003. 
June 2004 7-9 No. 4: Mechanical Degradation Revision 1 
SN0306L0207502.008. Revised Mechanical Properties Of Welded Tuff From The Lower 
Lithophysal Zone Of The Topopah Spring Tuff, Batch #3 (Test Dates: March 6, 2003 Through 
April 18, 2003). Submittal date: 06/20/2003. 
SN0307T0510902.003. Updated Heat Capacity of Yucca Mountain Stratigraphic Units. 
Submittal date: 07/15/2003. 
SNF37100195002.001. Hydraulic Fracturing Stress Measurements in Test Hole: ESF-AODHDFR1, 
Thermal Test Facility, Exploratory Studies Facility at Yucca Mountain. Submittal date: 
12/18/1996. 
SNL01A05059301.005. Laboratory Thermal Conductivity Data for Boreholes UE25 NRG-4, 
NRG-5; USW NRG-6 and NRG-7/7A. Submittal date: 02/07/1996. 
SNL01B05059301.006. Laboratory Thermal Expansion Data for Boreholes UE25 NRG-4, 
NRG-5; USW NRG-6 and NRG-7/7A. Submittal date: 02/07/1996. 
SNL02112293001.001. Results from Shear Stress Experiments on Natural Fractures on Samples 
from NRG-4 and NRG-6 Drillholes. Submittal date: 08/18/1994. 
SNL02112293001.003. Results from Shear Stress Experiments on Natural Fractures from NRG4 
& NRG-6. Submittal date: 03/13/1995. 
SNL02112293001.005. Mechanical Properties of Fractures in Specimens from Drillhole USW 
SD-9. Submittal date: 07/15/1996. 
SNL02112293001.006. Mechanical Properties of Fractures in Specimens from Drillhole USW 
SD-7 and ESF-TMA-MPBX-3 at Elevated Temperature. Submittal date: 07/30/1996. 
SNL02112293001.007. Mechanical Properties of Fractures in Specimens from Drillholes USW 
NRG-7/7A and USW SD-12. Submittal date: 08/08/1996. 
UN0106SPA013GD.004. Drift Scale Thermal Test (DST) REKA Probe Developed Data for 
Thermal Conductivity and Diffusivity for the Period 05/01/1998 to 04/30/2001 (Heated 
Measurements for Boreholes 151, 152, and 153). Submittal date: 06/28/2001. 
UN0201SPA013GD.007. DST REKA Probe Developed Data for Thermal Conductivity and 
Diffusivity for the Period 05/01/2001 to 12/31/2001 (Heated Measurements for Boreholes 151 
and 152). Submittal date: 01/07/2002. 
June 2004 7-10 No. 4: Mechanical Degradation ROCK-MASS CLASSIFICATION FOR LITHOPHYSAL ROCK 
(RESPONSE TO RDTME 3.05) 
No. 4: Mechanical Degradation 
APPENDIX A 
Revision 1 
June 2004 Revision 1 
Note Regarding the Status of Supporting Technical Information 
This document was prepared using the most current information available at the time of its development. This 
Technical Basis Document and its appendices providing Key Technical Issue Agreement responses that were 
prepared using preliminary or draft information reflect the status of the Yucca Mountain Projects scientific 
and design bases at the time of submittal. In some cases this involved the use of draft Analysis and Model 
Reports (AMRs) and other draft references whose contents may change with time. Information that evolves 
through subsequent revisions of the AMRs and other references will be reflected in the License Application 
(LA) as the approved analyses of record at the time of LA submittal. Consequently, the Project will not 
routinely update either this Technical Basis Document or its Key Technical Issue Agreement appendices to 
reflect changes in the supporting references prior to submittal of the LA. 
June 2004 No. 4: Mechanical Degradation Revision 1 
APPENDIX A 
ROCK-MASS CLASSIFICATION FOR LITHOPHYSAL ROCK 
(RESPONSE TO RDTME 3.05) 
This appendix provides a response to Key Technical Issue (KTI) agreement Repository Design 
and Thermal-Mechanical Effects (RDTME) 3.05. The agreement relates to concerns regarding 
the effects of lithophysal cavities on mechanical properties of the rock mass. 
A.1.1 RDTME 3.05 
Agreement RDTME 3.05 was reached during the U.S. Nuclear Regulatory Commission 
(NRC)/U.S. Department of Energy (DOE) Technical Exchange and Management Meeting on 
Repository and Design Thermal-Mechanical Effects held February 6 to 8, 2001, in Las Vegas, 
Nevada (Reamer and Williams 2001). There has been no submittal related to this KTI agreement 
to the NRC. 
The wording of the agreement is as follows: 
RDTME 3.05 
Provide the Rock Mass Classification Analysis (or some other document) 
including the technical basis for accounting for the effects of lithophysae. The 
DOE will provide a rock mass classification analysis (or other document), 
including the technical basis for accounting for the effects of lithophysae, 
expected to be available to NRC in FY 2002. 
The agreement focuses on a concern regarding the effects of lithophysal cavities (lithophysae) on 
rock-mass mechanical properties and on whether conventional rock-mass classification systems 
that are widely used in mining and rock engineering for jointed rock masses are suitable for 
estimating rock-mass mechanical properties of lithophysal rocks. The concern is further 
elaborated in Integrated Issue Resolution Status Report (NRC 2002, Section 2.1.7.3.3.2, 
pp. 2.1.7-14 and 2.1.7-15). The NRC concerns are as follows: 
 Use of empirical correlations between rock-mass mechanical properties and rock-mass 
quality indices, such as the Bartons Q index or the Bieniawskis rock-mass rating, to 
account for the effects of lithophysae is unprecedented and not supported by any data on 
or model investigation of the effects of lithophysae on the mechanical characteristics of 
rock. 
 Bartons Q index or Bieniawskis rock-mass rating index may be appropriate for 
accounting for the effects of fractures. Some modification of these index values would 
be necessary if the DOE uses the indexes to account for the effects of lithophysae. The 
A.1 KEY TECHNICAL ISSUE AGREEMENT 
A-1 June 2004 No. 4: Mechanical Degradation Revision 1 
technical basis for such modification is all the more important because about 75%1 of 
the emplacement area may lie within the lithophysal rock units. 
A.1.2 Related Key Technical Issue Agreements 
RDTME 3.02This KTI agreement requires that drift degradation and ground support analyses 
be conducted for critical combinations of in situ, thermal, and seismic stresses. These load 
combinations are addressed in Section 5. 
RDTME 3.04This KTI agreement involves providing a geotechnical parameters report that 
includes rock-mass property estimates of the lithophysal rocks of the Topopah Spring formation. 
RDTME 3.05 deals specifically with providing estimates of the lithophysal rock-mass properties. 
Thus, the work performed to resolve RDTME 3.05 also forms a portion of the geotechnical 
parameters report that is used to resolve RDTME 3.04. Sections 3 and 4 examine the rock-mass 
properties estimates. 
RDTME 3.10This KTI agreement requires verification of the adequacy of the use of 
two-dimensional models for analysis of drift degradation. Sections 4.2.2.2 and 5.3.3.1.2 discuss 
the use of two- and three-dimensional models for drift degradation analysis. 
RDTME 3.11This KTI agreement requires examination of the long-term degradation of the 
rock mass in lithophysal and nonlithophysal rocks. This appendix discusses the methodology for 
accounting for lithophysae in the mechanical rock-mass properties and material models. 
Section 5.3.3.2.4 summarizes the specific approach to accounting for long-term strength 
degradation of material properties via use of static fatigue testing of tuffs. 
RDTME 3.12This KTI agreement requires a dynamic analysis of ground support systems 
during the preclosure phase using site-specific ground motions and discontinuum numerical 
modeling. The technical basis document centers on the postclosure dynamic analysis of 
lithophysal and nonlithophysal rocks using discontinuum methods. The preclosure dynamic 
analysis of ground support is provided in Scoping Analysis on Sensitivity and Uncertainty of 
Emplacement Drift Stability (BSC 2003a). 
RDTME 3.13This KTI agreement requires technical justification for boundary conditions for 
models used in drift degradation and ground support analyses. Section 5 discusses static and 
dynamic mechanical and thermal boundary and initial conditions. 
TSPAI 2.02, Items 58 and 62TSPAI 2.02, Items 58 and 62 related to the inclusion of rockfall 
and its potential mechanical impacts on engineered barriers and on the thermal-mechanical 
impacts of long-term rock-mass degradation on engineered barriers and potential hydrologic 
changes in the rock mass. RDTME 3.05 deals specifically with the estimation of rock-mass 
properties of the lithophysal rocks of the Topopah Spring formation and the development of 
rock-mass material and numerical models for representing fracture, rockfall, and long-term 
degradation under in situ, thermal, and seismic loading. The estimates made for rockfall and 
1 Although 75% is used in Integrated Issue Resolution Status Report (NRC 2002), the actual percentage is 
approximately 85%. 
A-2 June 2004 No. 4: Mechanical Degradation Revision 1 
long-term degradation and change of opening shape feed performance assessment studies of the 
engineered barriers. Section 1 describes this integration in more detail. 
A.3 RESPONSE 
The strategy for developing estimates of rock-mass properties and material models for 
lithophysal rocks was presented by Board (2003). The strategy does not rely on rock-mass 
classification methods for empirically deriving the rock-mass mechanical properties for the 
lithophysal rock. Instead, the approach for determining the effects of lithophysae involves: 
1. Conducting a detailed site-specific geologic description of the lithophysal rock mass in 
the Exploratory Studies Facility (ESF) and Enhanced Characterization of the 
Repository Block (ECRB) Cross-Drift. 
2. Testing of large-diameter core in the laboratory and in situ slot testing to establish 
properties of rock material as functions of specimen size and lithophysal porosity for a 
limited data set. 
3. Calibrating the PFC (i.e., collectively, the PFC2D (BSC 2002a) and PFC3D 
(BSC 2002b) and UDEC (BSC 2002c) discontinuum numerical models of unfractured 
lithophysal rock to provide a mechanics-based predictive tool of lithophysal behavior. 
This is described in Section 4.2.3. 
4. Utilizing the models to extrapolate rock-mass mechanical behavior to a wide range of 
rock-mass lithophysal conditions, including the size, shape, porosity, and distribution 
of lithophysae to establish their impact on variability of the rock-mass properties. 
5. Developing bounding rock-mass mechanical properties ranges based on the laboratory 
testing, numerical extrapolation, and examination of the spatial variability of 
lithophysal porosity. 
A.2 RELEVANCE TO REPOSITORY PERFORMANCE 
Agreement RDTME 3.05 deals with demonstrating an understanding of the material properties 
of the lithophysal rock mass. The material properties of importance to repository design and 
performance are the thermal, deformability, and strength properties and their potential changes 
with time, as well as variability within the repository host horizon. A description of geotechnical 
properties is given in Section 3, and a detailed characterization of all properties is given in 
Subsurface Geotechnical Parameters Report (BSC 2003b). These properties are used as direct 
input to numerical models that predict the drift degradation behavior of the emplacement drifts 
when subjected to in situ, thermal, and seismic loading, as well as time-dependent changes in 
mechanical properties. The results of these calculations are estimates of the volume and size of 
potential rockfall and the change in the shape of the tunnels as functions of time. This 
information directly feeds calculations of the structural stability of waste packages and drip 
shields due to rockfall in the preclosure and postclosure time frames. The rockfall also 
potentially causes an insulating effect around the drip shields, which could impact in-drift 
temperatures as a function of time. The change in the shape of the tunnels as a function of time 
feeds the analysis of seepage to emplacement drifts. 
A-3 June 2004 No. 4: Mechanical Degradation Revision 1 
The development and technical basis for the above strategy is documented in Lithophysal Rock 
Mass Mechanical Properties of the Repository Host Horizon (BSC 2004a) and the Subsurface 
Geotechnical Parameters Report (BSC 2003b). The results from these assessments are 
implemented in design calculations and sensitivity analyses presented in detail in Drift 
Degradation Analysis (BSC 2004b) and Evaluation of Emplacement Drift Stability for KTI 
Resolutions (BSC 2004c) and summarized in Sections 4 and 5. 
The strategy is summarized below. 
Geologic Description of the Physical and Mineralogical Characteristics of Lithophysae 
within the Lower Lithophysal Unit (Tptpll)A mapping program was undertaken in the ECRB 
Cross-Drift to describe the statistical variability of the size, shape, porosity, and distribution of 
lithophysae, as well as interlithophysal fracturing across the entire Tptpll. This work provides 
the basis for understanding the physical variability of the lithophysal rock mass and is discussed 
in Section 2.3.2. 
Laboratory and Field Testing of Lithophysal RockLarge-scale testing is necessary to 
determine the mechanical properties of lithophysal rock due to the presence of the lithophysal 
voids. A laboratory (approximately 0.3 m diameter samples) and in situ (approximately 1.1 m 
samples) testing program was used to obtain and mechanically test large specimens of the upper 
lithophysal unit (Tptpul) and the lower lithophysal unit (Tptpll). These laboratory data provide a 
base (in addition to the small core testing database) for relating strength and modulus to 
lithophysal porosity as well as determining the relationship between modulus and strength. 
Additionally, the field-scale testing provides information on the effects of scale involving a 
greater specimen size and variability of shape, size, and distribution of lithophysal porosity. 
Section 3.2.1 discusses the mechanical properties of intact rock, including correlation of these 
properties with rock porosity. Section 3.2.3 discusses the in situ mechanical testing of 
lithophysal rock. 
Development of the Numerical Modeling Approach for the Representation of Lithophysal 
RockThe rock sample size-dependency and variability in lithophysal rock-mass properties 
introduced by the lithophysae make a conventional, statistically based laboratory testing program 
impractical. An essential part of the resolution strategy is the development of a numerical 
modeling approach that is capable of simulating the mechanics of deformability and yield of the 
lithophysal rock. Different discontinuum numerical modeling approaches, the PFC and UDEC 
models, are used to examine the basic mechanisms of how lithophysae affect the failure 
characteristics and moduli of the Tptpul and Tptpll. These models were chosen because of their 
ability to represent physical voids in a material and for their capability to model complex failure 
mechanisms, such as fracture initiation and propagation between voids. PFC3D was used to 
explore the suitability of the two-dimensional modeling. 
The model matrix (containing no voids or fractures) strength and elastic moduli were first 
calibrated through comparison of the model response to the results from laboratory testing of 
nonlithophysal rocks. The models were then validated for lithophysal rocks by assuming the 
same matrix as the nonlithophysal case, and then adding voids of varying size to replicate the 
lithophysal porosity of the rock mass. Simulated uniaxial compression tests were then compared 
to laboratory and field testing results to verify the general predictability of the approach. An 
June 2004 A-4 No. 4: Mechanical Degradation Revision 1 
outcome of this process was an explanation for the mechanisms of strength and modulus 
reduction that accompanies additional porosity. Triaxial compression experiments were used to 
develop estimated yield criteria and dilation angles for lithophysal rocks as a function of 
porosity. Sections 4.2.1 and 4.2.3 and this appendix discuss the numerical modeling approach, 
material model requirements, and validation of the adopted models for predicting lithophysal 
rock behavior. 
Extrapolation of the Variability of Lithophysal Rock-Mass Properties Using Numerical 
ModelsThe calibrated PFC model is used for extrapolation purposes to supplement the 
laboratory and field testing database. The variability of rock mechanical properties due to 
lithophysae shape and spatial distribution is studied by randomly creating voids of simple shape 
(e.g., circle, triangle, or star) in the matrix material and by modeling realistic void shapes and 
distributions corresponding to lithophysal cavities identified in ECRB Cross-Drift panel maps 
(1  3 m). Variability in the rock response and size effect appears to be a function primarily of 
the distribution of the lithophysae (i.e., percentage porosity and how evenly distributed it is 
through the rock mass), and, to a lesser extent, the deviation of the shape of true voids from 
circular voids. The property scatter apparent at given values of lithophysal porosity reduces 
significantly when the data are plotted in terms of strength versus Youngs modulus, suggesting 
that this correlation is independent of void geometry. Section 4.2.2 and this appendix discuss 
how the numerical models for lithophysal rock were calibrated and used to model lithophysal 
rock behavior. 
Establishing the Range of Material Properties of Lithophysal Rocks for Design and 
Performance AssessmentThe laboratory and field data are integrated with the computational 
property variability estimates to establish the range of strengths and moduli that represent the 
rock-mass properties in the ECRB Cross-Drift and, especially, the Tptpll. A bounding approach 
based on parametric modeling over the entire range of estimated rock-mass properties is used for 
conducting drift degradation analyses. These rock parameters ranges are then used as a basis for 
excavation stability calculations. Sections 4.2.3 and 4.2.4 and this appendix describe how 
lithophysal rock mass properties estimates were developed and bounded. 
Development of a Drift-Scale Emplacement Drift Degradation Modeling ToolThe impact 
of modeled spatial variability of lithophysal porosity in the Tptpll and establishment of the 
conservatism of use of a bounding range approach is examined through numerical analysis of 
nonhomogeneous rock-mass strength and deformability. A discontinuum model based on the 
UDEC program was developed and calibrated to represent the range of strength and moduli 
(discussed above). The range of rock mass properties that were used to calibrate the UDEC 
model are presented in Sections 4.2.3 and 4.2.4. Validation of the UDEC model was performed 
against laboratory and field observations, as described in Section 4.2.7. The UDEC model was 
then used to estimate the stability and rockfall for in situ, thermal, and seismic loading, as well as 
time-dependent strength degradation. A parametric study was conducted using the range of 
strength and moduli to provide lower-bound, mean, and upper-bound estimates of rock-mass 
yield and rockfall for the Tptpll. Section 4.2 and this appendix provide further explanation of 
how lithophysal rock mass properties were calibrated, developed, and bounded. 
June 2004 A-5 No. 4: Mechanical Degradation Revision 1 
The information in this appendix is responsive to KTI agreement RDTME 3.05 made between 
the DOE and NRC for the following reasons: 
 The impact of lithophysae on mechanical properties is accounted for based on laboratory 
and field testing at various physical scales. 
 Numerical analyses of rock containing lithophysae have been used to enhance 
understanding and provide evidence of the mechanisms impacting mechanical properties 
of rock. 
 The spatial variability of lithophysae and bulk porosity in the Tptpll is accounted for 
based on geologic field mapping and numerical extrapolation. 
 This variability is also accounted for in design analyses through parametric studies that 
cover a bounding range of rock parameters. 
The approach of applying numerical modeling to the rock containing lithophysae is corroborated 
with laboratory and field test data and observation of excavation effects in the ECRB Cross-Drift 
in the Tptpll unit. 
The information in this report is responsive to agreement RDTME 3.05 made between the DOE 
and NRC. The report contains the information that DOE considers necessary for NRC review 
for closure of this agreement. 
A.4 BASIS FOR THE RESPONSE 
The Yucca Mountain exploratory excavations for the repository are located within a portion of 
the Topopah Spring formation, a volcanic welded tuff subdivided into four subunits 
characterized by their geologic features and structure. These four subunits are the upper 
lithophysal unit (Tptpul), the middle nonlithophysal unit (Tptpmn), the lower lithophysal unit 
(Tptpll), and the lower nonlithophysal unit (Tptpln). The nonlithophysal units (Tptpmn and 
Tptpln) are fine-grained, low porosity (i.e., approximately 11%), strong volcanic rocks that 
contain abundant but relatively nonpersistent cooling fractures. The lithophysal units (Tptpul 
and Tptpll) are composed of the same fine-grained matrix material but have additional porosity 
contributed by lithophysae (i.e., open voids that are a result of gas localization during the cooling 
process) and by rims and spots formed from the crystallization of vitric rock in the presence of 
vapor. These lithophysae, which vary in size from the millimeter to meter scale, make up about 
10% to 30% of the volume of the Tptpll subunit, averaging approximately 15%. Rim and spot 
material has a porosity averaging 30% and makes up about 4% of the volume of the Tptpll, 
ranging from about 1% to 12%. 
Agreement RDTME 3.05 addresses the means by which the presence of lithophysae will be 
accounted for in thermal and mechanical analysis models. At the time of the writing of that 
agreement (Reamer and Williams 2001), the use of modified empirical classification and design 
methods typically used in rock mechanics (e.g., Bartons Q index and Bieniawskis rock-mass 
rating classification schemes) to account for lithophysae was being considered. In response to 
June 2004 A-6 No. 4: Mechanical Degradation Revision 1 
NRC concerns, it was determined that an alternative approach based on development of 
site-specific rock-mass properties data for the lithophysal rocks was warranted. 
The present approach being used in estimating lithophysal rock-mass properties relies on a 
program incorporating site-specific laboratory and field testing supplemented by numerical 
modeling of lithophysal rock. As a result, the DOE has conducted and analyzed tests on 
lithophysal rock in the Tptpul and Tptpll units, consisting of large-diameter core (up to 0.3 m 
diameter) and in situ compression tests (slot tests up to 1.1 m across). Additionally, the Tptpll 
portion of the ECRB Cross-Drift has been systematically mapped to identify lithophysal rock 
characteristics (abundance, shape, and size variability), and work is underway to similarly map 
portions of the Tptpul unit in the ECRB Cross-Drift as well. Lastly, two- and three-dimensional 
numerical modeling of larger-scale samples (1 m scale) with realistic lithophysal voids have 
been carried out using the PFC and UDEC discontinuum programs. These numerical simulations 
have been utilized to further develop the stress-strain response of lithophysal rocks and to 
establish rock parameter ranges necessary to extrapolate the possible behavior and material 
properties of lithophysal bulk rock. This appendix deals specifically with the geotechnical 
characterization of the lithophysal rock mass and the impact of lithophysal porosity on 
mechanical properties. The impact of lithophysal porosity on thermal properties is derived from 
laboratory and field thermal testing and is discussed in Section 3.2.5. 
Figure A-1. Example Tptpll Panel Map and Showing Lithophysae, Rims, Spots, and Lithic Clasts 
A.4.1 Geologic Mapping of Physical and Mineralogical Characteristics of Lithophysae 
Documentation of the variability of physical and mineralogical characteristics of lithophysae 
forms the basis for further analytical modeling. These field characteristics are documented via 
the use of several techniques using tape and survey measurements of lithophysae sizes taken at 
specified intervals across the unit thickness of the Tptpll exposed within the ECRB Cross-Drift. 
Detailed mapping and photography of 1  3 m panels maps along the walls of the ECRB 
Cross-Drift serves as a core of site-specific documentation (Figure A-1). 
The results of this work are summarized in this appendix and described in Subsurface 
Geotechnical Parameters Report (BSC 2003b, Section 8.8.4). Each of these techniques has 
resulted in an estimate of the lithophysal porosity along the tunnel, as well as estimates of the 
abundance of the lithophysae rims and spots and the shape and size distribution of the 
June 2004 A-7 No. 4: Mechanical Degradation Revision 1 
lithophysal porosity. The distribution of lithophysal porosity and other features obtained from 
the 15 m averaged tape traverse data corrected to panel map and angular traverse data is 
presented in Figure A-2. This figure shows that average lithophysal porosities greater than 20% 
makes up only a small (about 10%) of the Tptpll. The greatest abundance of porosity levels is 
found in the range of approximately 10% to 20%. Since lithophysal porosity is the greatest 
controlling factor in the mechanical properties of the rock mass, the abundance of lithophysal 
porosity can be directly related to the abundance of mechanical properties, as discussed 
Section A.4.5. 
Source: BSC 2003b, Attachment VII, Figure VII-15. 
NOTE: Relatively thin bands of highest lithophysal porosity (greater than 20%) occur near the top of the Tptpll, with 
decreasing values near the bottom of the subunit. Higher and lower localized volumes of lithophysal 
porosity may be present. The 5 m averaged large-lithophysae inventory (brown) is not included as part of 
the lithophysal or total porosity curves. 
ECRB Cross-Drift 
Field mapping (an example of the fitted lithophysal cavity porosity is shown in Figure A-2) has 
been used to develop a model for simulating the spatial variability of lithophysal porosity within 
the Tptpll (see Section 2.3.2). The observed variations in lithophysal porosity within the Tptpll 
are assumed to be laterally continuous across the volume of rock sampled and to vertically show 
a statistically distinguishable sequence of stratiform subzones. The ECRB Cross-Drift transects 
the entire Tptpll unit allowing determination of the stratiform nature of laterally continuous 
subzones of lithophysal porosity within the Tptpll. The lateral continuity of lithostratigraphic 
features and the projection of these features along the apparent dip of the ECRB Cross-Drift are 
used to develop a vertical cross section of the distribution of lithophysal porosity within the 
Tptpll. 
Figure A-3 presents two simulated (overlapping) vertical projections of lithophysal porosity for 
50-m-tall vertical cross sections extending from the top and bottom of the Tptpll. Each of the 
cells of these cross sections represent a volume of the Tptpll and have a lithophysal porosity 
associated with it. As seen in these cross sections, the lithophysal porosity occurs as stratiform 
subzones, with the highest values (i.e., 20% or higher) occurring in thin bands near the top of the 
Figure A-2. Example of Calculated Porosity of Lithophysal Cavities (red), Rims (violet), Spots (blue), 
Matrix-Groundmass (green), and the Total Porosity (black) in the Tptpll Exposed along the 
June 2004 A-8 No. 4: Mechanical Degradation Revision 1 
subzone. The lowest lithophysal porosities occur in the lower portions of the subzone near the 
contact with the Tptpln. These simulated cross sections are used as a basis for examining the 
impact of spatial variability on rock-mass mechanical response and comparing assumptions of 
homogenous rock-mass properties. 
Source: BSC 2003b, Section 9.4, Figure 9-46. 
NOTE: Cross section A is a 50  200 cell table representing a 1  1 m grid and cross section B is a 20  80 cell 
table representing a 2.5  2.5 m grid for the simulated section at 17 + 56 in the ECRB. Cross sections C 
and D represent simulated section at 20 + 14 in the ECRB. 
Figure A-3. Illustration of the Process of Sampling and Modeling Spatial Variability Using Lithophysal 
PorosityTwo 50  200 m Simulated Cross Sections at Overlapping Upper (A/B) and 
Lower (C/D) Sections of the Tptpll 
Results of Geological MappingGeological mapping studies in the Tptpll have provided the 
baseline information required for inclusion in the thermal and mechanical analyses. This 
information includes a determination of the following: 
 Abundance of lithophysal cavities, rims, and spots in the Tptpll 
 The size distribution of lithophysae 
 An inventory of lithophysae shapes 
June 2004 A-9 No. 4: Mechanical Degradation Revision 1 
 Panel map sampling of the two-dimensional spatial variability of lithophysal porosity 
within the Tptpll (i.e., the distribution of lithophysal cavity space versus solid rock 
bridge length). 
A.4.2 Laboratory and In Situ Thermal and Mechanical Testing of Lithophysal Rocks 
A series of laboratory and field tests of mechanical and thermal properties of Tptpul and Tptpll 
rock have been completed and reviewed in Section 3.2.1 and have been documented in detail in 
Subsurface Geotechnical Parameters Report (BSC 2003b), which reviews the available 
geotechnical database. A review of the results presented in Section 4.2 follows. Lithophysal 
Rock Mass Mechanical Properties of the Repository Host Horizon (BSC 2004a) provides a 
detailed assessment of the relationship between porosity and mechanical rock properties. 
Laboratory TestingDue to the variable size of lithophysae, only large-diameter core samples 
(e.g., ideally about several lithophysae diameters in diameter) can be used for estimating rock 
strength values. In 1985, compression tests were conducted on cores with a diameter of 0.27 m 
(10.5 in.) taken from the Tptpul at Busted Butte. An extensive drilling program was undertaken 
in the ESF and ECRB in 2002 to provide representative large diameter (0.29 m, 11.5 in.) core 
samples of lithophysal rock from the Tptpul and Tptpll zones within the repository host horizon. 
Fourteen suitable specimens were recovered and tested in uniaxial compression at dry, saturated, 
and roommoisture conditions. The lithophysal porosity of these core samples, estimated from 
the surface line counting of features, ranges from approximately 10% to 30%. The results of 
these tests, in terms of unconfined compressive strength versus Youngs modulus, are shown in 
Figure A-4. 
June 2004 A-10 No. 4: Mechanical Degradation Revision 1 
NOTE: Variation in range of strength and modulus is primarily a function of lithophysal porosity and level of 
saturation. Linear fits to data presented for 0.29 and 0.27 m cores separately. 
Figure A-4. Unconfined Compressive Strength versus Youngs Modulus for 0.29 m (11.5 in.) and 
0.27 m (10.5 in.) Diameter Cores of Tptpul and Tptpll 
In Situ Mechanical TestingThree in situ flatjack slot tests were conducted on 1.1 m span 
samples in the Tptpll and Tptpul (see Section 3.2.3). The tests involve cutting two thin, parallel 
slots, separated by approximately 1.1 m, in the sidewall or floor of the tunnel. The lithophysal 
content from the face and slots of the blocks were mapped to define the size, shape, and 
percentage of lithophysae and spots. Steel flatjack bladders were inserted into the slots and 
pressurized to load the sample in a state of near-uniaxial compression (the ends of the sample are 
not freed, thus there is some confinement from the sample ends). Deformations of the block 
were determined in a central, internal borehole to allow calculation of the stress-strain response 
of the rock mass. 
These tests have shown that the in situ deformation modulus is approximately 1.0 to 3.0 GPa 
(one test was located in the poorest quality failed rock in the tunnel springline area, yielding a 
modulus of 0.5 GPa). In general, it has been concluded that these tests illustrate the impact of 
mining and stress-related damage in the immediate sidewall of the tunnels and are not 
necessarily indicative of rock-mass strength and moduli. However, the tests illustrate some 
component of the size effect present in sample size increased from 0.3 m to approximately 1 m. 
June 2004 A-11 No. 4: Mechanical Degradation Revision 1 
A.4.3 Development of Basic Mechanical Constitutive Models for Lithophysal Rock 
through Calibration of Numerical and Analytical Models 
Conducting a statistically significant number of large core mechanical tests and in situ scale tests 
on lithophysal rocks as a means of determining the variability of rock-mass properties is not 
practical, due to sampling difficulties and laboratory testing limitations. Instead, the strategy 
described by Board (2003) suggested that calibration and validation of a suitable mechanical 
numerical modeling approach could be used to supplement available laboratory and field testing 
and to clarify understanding of the mechanical response of lithophysal rock. This approach 
allows investigation of the effects of lithophysae size, shape, and porosity distribution on 
mechanical constitutive behavior of lithophysae-containing rock. The PFC program employs a 
micromechanical modeling approach, which represents the rock matrix as a large number of rigid 
circular or spherical particles that are bonded together at their contact points with simple shear 
and tensile bonds and normal and shear stiffness. The UDEC program, also based on a 
discontinuum approach, is used as an alternative computational model to simulate compressive 
load testing on lithophysal rock specimens. 
As a simulated PFC rock specimen is stressed, the inter-particle bonds can fail, leading to 
frictional sliding between particles. This conceptually simple model can exhibit complex 
constitutive behavior including realistic growth of stress-induced fracturing. Of importance to 
the present problem is that holes of arbitrary shape, size, and distribution can be represented in 
the model as physical entities. Thus, stress concentrations and fracturing between holes can be 
represented in a realistic way, allowing detailed examination of deformation and failure 
mechanisms. 
The PFC model has been calibrated against the laboratory testing data discussed in 
Section 4.2.3.2. Numerical compressive strength experiments were conducted first for 
nonlithophysal welded tuff to derive strength properties of the matrix (Figure A-5). In 
subsequent analyses, simple uniformly distributed circular (and later triangular) holes were 
added to the model, and the relationship between laboratory-derived strengths, moduli, and 
porosities was examined. The modeling results showed good agreement with laboratory data 
(Figure A-6). Results indicated that the primary strength-decreasing effect of the lithophysae is 
due to initiation and propagation of tensile splitting between lithophysae under compressive load. 
As porosity increases, the spacing between lithophysae decreases, and rock acquires a greater 
propensity for tensile splitting at the lower applied stresses. 
June 2004 A-12 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Section 6.5.6, Figure 6.5-20. 
NOTE: Nonlithophysal welded tuff (upper left) is calibrated to provide matrix properties. Circular holes (upper right) 
provide a simple model of lithophysae, whereas lower models provide more realistic (hand-stenciled and 
digitized) shapes from Tptpll panel maps. Nonlithophysal rock fails in a brittle fashion through propagation 
of a major shear fracture (composed of small tensile fractures) through the sample. Lithophysal samples fail 
due to tensile splitting (each red line is a bond breakage between small particles) between holes. Variability 
in lithophysal strength arises due to abundance, shape, and distribution of the holes throughout the sample. 
Figure A-5. PFC Calibration Experiment Samples and Their Respective Unconfined StressStrain 
Curves for Cases of Circular and Stenciled Lithophysae Shapes 
Further PFC numerical testing was performed during which lithophysal panel maps (see 
Figure A-1) were discretized and used as 1.0  1.0 m compressive test specimens (Figure A-5) 
for the model. As shown in Figure A-6, the simple models containing circular holes display less 
variability in results than similar analyses that utilize actual lithophysae shapes. Actual shapes 
result in greater variability in test results (particularly at low porosities) and a lower estimate of 
mean strength and modulus. 
The conclusion drawn from the PFC model calibration is that this approach provides a 
reasonable methodology for simulating the mechanical response of lithophysal welded tuff to 
stressing and that this tool can be used to study variability in material response in addition to 
laboratory and field testing. 
June 2004 A-13 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Section 6.5.2, Figures 6.5-13 and 6.5-14. 
NOTE: For a given porosity, there is a greater variability in the strength of the PFC samples when true lithophysae 
shapes are introduced. This effect is particularly significant at low porosities due to the greater variability of 
distribution and solid bridge lengths between lithophysae in a given sample volume. 
Figure A-6. Plots Showing Data from Large Core (0.29 diameter) Compression Testing of Tptpul and 
Tptpll Compared to PFC Simulations Using Circular and Triangular Shaped Lithophysae as 
well as Actual Stenciled Shapes from Enhanced Characterization of the Repository Block 
Panel Maps 
A.4.4 Extrapolation of Mechanical Response Using the PFC ProgramEstablishing 
Bounding Ranges for Rock-Mass Properties 
The calibrated PFC and UDEC computational models are used as an extrapolation tool to 
examine the impact of lithophysae size, shape, porosity, and distribution on rock-mass 
June 2004 A-14 No. 4: Mechanical Degradation Revision 1 
mechanical properties. Parametric studies have been conducted with simple (e.g., circular and 
triangular) shapes of lithophysae as well as actual, complex shapes and distributions digitized 
from ECRB Cross-Drift panel maps. The results of these studies (shown in Figure A-6) are used 
as a guide to understand the variability of mechanical properties for a given level of lithophysal 
porosity. Figure A-7 shows the large-core laboratory mechanical test data given in Figure A-4 
plotted with the results of the PFC modeling studies assuming realistic lithophysae shapes and 
distributions from panel maps. As seen in this plot, approximate bounding estimates of the 
unconfined compressive strength and Youngs Modulus have been drawn for the range of 
saturated and unsaturated laboratory data as well as the PFC panel map extrapolations of 
lithophysal shape and porosity variability. The linear fit to the 0.29 m core data, which is used as 
a base-case measure of the mechanical properties for room dry conditions, is shown as is the 
mean of the two bounding curves. The plots are dashed for those points that fall outside the 
range of the data. The laboratory data shows that the saturated environmental conditions form a 
lower bound to the property ranges and that the PFC extrapolations of shape variability fall 
within the bounding curves. 
Source: BSC 2004a, Section 6.6.3, Figures 6.6-5 and 6.6-7. 
NOTE: The bounding curves are empirically derived to include all large core laboratory data, including 0.267 m 
diameter saturated cores from Busted Butte. Characteristic base-case properties, derived from 0.29 m 
diameter cores, are also shown. Extrapolated (PFC) strength values below 10 MPa represent small 
samples with large lithophysae and inconsistent with observations of lack of yielding in ECRB and ESF. 
Figure A-7. Estimated Upper and Lower Bounds of Unconfined Compressive Strength and Youngs 
Modulus for Lithophysal Rock. 
June 2004 A-15 No. 4: Mechanical Degradation Revision 1 
The PFC models estimate the impact of lithophysae shape and distribution on the strength of 
small (meter scale) samples of rock. When the lithophysae size is roughly proportional to the 
sample size (the case in some of the PFC panel map models), it is possible to produce strengths 
that are unrealistically low. This sample size effect can be seen in Figure A-7 in the shaded 
region in which no laboratory data are available, and only PFC-extrapolated values are seen. To 
test whether these lower strengths can actually occur in the field or are due to the extrapolation 
procedure, they were used as input to a number of drift-scale stability models of the ECRB 
subjected to in situ stress conditions. The results of these simulations, in terms of expected 
stability of the tunnel, were compared to field observations to assess if such low strength values 
are possible. The UDEC discontinuum program was run for the range of strengths shown in 
Figure A-7. It was found that unconfined compression strengths with values below about 
10 MPa predict that significant sidewall spalling of the ECRB Cross-Drift and the ESF would be 
observed in the lithophysal rock. Significant sidewall spalling, as predicted by the model, would 
involve obvious shear failure of the tunnel springlines and breakout of the sidewall. On the 
contrary, these drifts are stable and in good structural condition despite the fact that only light 
bolting (none in the sidewalls of most of the drift) is used. Therefore, a lower-bound strength 
cutoff of 10 MPa, which coincides with the lowest strength properties measured from large core 
compression, is assumed. 
A.4.5 Abstraction of the Thermal and Mechanical Constitutive Response into Numerical 
Predictive Models 
Estimate of Rock Strength Categories for Bounding Property Range AnalysisThe basic 
constitutive behavior defined by testing and PFC extrapolations must be abstracted into a design 
assessment tool capable of representing the lithophysal rock-mass response as well as allowing 
for modeling of the drift degradation process. To accomplish these objectives, a similar 
two-dimensional discontinuum approach is used and implemented in the UDEC program, which 
is similar to PFC but more efficiently examines large-scale mechanical problems. The process of 
implementation, calibration, and validation of this approach for representation of lithophysal 
rock is reviewed here and described in detail in Drift Degradation Analysis (BSC 2004b, 
Section 7). 
A UDEC model is developed with a finely discretized block structure in which the block sizes 
are approximately equal to the spacing of the ubiquitous fracture fabric of the lithophysal rock 
mass (Figure A-8). The fracture fabric within the Tptpll was mapped in a series of small-scale 
fracture panel maps as discussed in Drift Degradation Analysis (BSC 2004b, Section 6.1.4.1). 
The 18 m of traverse shows multiple fracture sets with an average spacing of 0.05 m and an 
average trace length of 0.29 m. The blocks themselves are elastic but bonded at their contacts 
with a Mohr-Coulomb material capable of simulating shear and tensile failure. Thus, the overall 
response of the rock mass is that of an equivalent Mohr-Coulomb material but with the ability to 
fracture between blocks in shear or tension as the stresses dictate. This fracturing, along with 
gravitational or seismic accelerations, may cause detachment of rock fragments and rockfall. 
June 2004 A-16 No. 4: Mechanical Degradation Revision 1 
NOTE: Overall, the model behaves as an equivalent Mohr-Coulomb rock mass until fracture occurs between blocks, 
allowing detachment of rock fragments and rockfall. 
Figure A-8. Discretization of the UDEC Discontinuum Model into Small and Irregular Blocks to 
Represent the Lithophysal Rock Mass 
For the model to represent the lithophysal rock mass, it must be calibrated to reproduce the 
modulus and compressive strength ranges that have been determined from the laboratory testing 
and PFC extrapolations (see Figure A-7). The approach to modeling used here is to assume the 
rock-mass properties are constant and homogeneous for any given tunnel cross section. This 
assumption is justified for two reasons: 
 The lithophysae are, in general, much smaller than the diameter of the tunnel and, thus, 
do not need to be modeled explicitly. Instead, the effect of the porosity is accounted for 
in the overall strength and modulus of the rock mass 
June 2004 A-17 No. 4: Mechanical Degradation Revision 1 
 Although the rock mass porosity is variable for any given cross section, use of a 
bounding range approach assuming constant properties is conservative. This assumption 
was examined by representing spatial variability of lithophysal porosity (and the 
associated strength and modulus spatial variability) and is presented in 
Section 5.4.2.2.2.1 and discussed below. The results show that accounting for spatial 
variability produces rock mass strength and moduli that tend toward the mean of the 
lithophysal porosities within a given cross section. Thus, use of parametric analysis 
utilizing constant properties that encompass the range of material properties values will 
produce conservative estimates of degradation. 
The impact of variability of the mechanical properties of the Tptpll on drift stability is accounted 
for by conducting parametric stability analyses that cover the entire range of probable rock-mass 
properties (i.e., from the highest porosity and lowest strength and modulus to the lowest porosity 
and highest strength and modulus). To accomplish this task, the strength and modulus range of 
the large-scale laboratory testing results is divided into five equally spaced categories 
(Figure A-9). By conducting numerical analyses over this range of data, the levels of rock 
quality and rock-mass response (from lowest to highest porosity ranges and size effects) can be 
covered. 
June 2004 A-18 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Section 6.6.3, Figure 6.6-7. 
NOTE: Base-case average properties defined for each category are the mean and the upper and lower bounds of 
each range. Category 1 is highest porosity, lowest quality rock and category 5 is lowest porosity, highest 
quality rock. Results of rock-mass strength estimates from spatial variability studies are shown assuming 
mean and lower bound properties. 
Figure A-9. Lithophysal Rock Strength and Modulus Range Divided into Five Rock Strength Categories 
Covering the Large Core Laboratory Testing and PFC Extrapolation Lithophysal Shape 
Extrapolation Studies 
The abundance of occurrence of these rock-mass categories in the Tptpll, and thus their 
importance from a stability standpoint, is demonstrated by the mapped lithophysal cavities 
shown in Figure A-2. The data in Figure A-9 are divided into five lithophysal porosity 
categories that roughly correspond to the five rock-mass strength categories shown in the same 
figure. Table A-1 presents an approximate relationship between lithophysal porosity abundance 
and strength categories. 
June 2004 A-19 No. 4: Mechanical Degradation Revision 1 
Table A-1. Approximate Relationship of Rock Strength Category to Lithophysal Porosity Abundance in 
the Tptpll 
Approximate Lithophysal 
Porosity Range (%) 
Rock Strength 
Category 
Abundance in Tptpll 
(%) 
20 to 25 
15 to 20 
10 to 15 
<10 
1 >25 3 
2 7 
3 26 
4 34 
5 30 
Source: BSC 2004a, Sections 6.6.1 and 6.6.2, Table 6.6-1. 
Impact of Spatial Variability of Lithophysal Rock PropertiesThe approach described above 
assumes that the rock properties are constant for a given tunnel cross section. Field mapping 
shows that lithophysal porosity is not uniform and homogeneous but varies over a distance scale 
of perhaps a few meters (see Figure A-2). Therefore, in reality, the rock mass at the drift scale 
has spatially variable lithophysal porosity, and thus spatially variable mechanical properties. The 
question arises as to whether the use of ranges of constant properties in drift stability modeling is 
a conservative approach or whether there needs to be an explicit accounting for spatial variability 
in properties. 
To investigate the impact of lithophysal spatial variability on mechanical properties ranges, the 
simple model of Tptpll lithophysal variability (see Figure A-3) was used as a basis for a 
numerical examination of its impact on rock-mass mechanical properties. Numerical 
compression experiments, similar to the PFC modeling studies, were conducted on 5  10 m 
rock-mass samples selected randomly from the Tptpll spatial variability model. The selected 
samples were of a drift scale to sample the range of lithophysal variability seen from the top to 
bottom of the Tptpll. Each sample was subdivided into meter-scale elements which were 
assigned (from the spatial variability model) a lithophysal porosity and, in turn, unconfined 
compression strength and Youngs modulus obtained from the mean or lower bound values in 
Figure A-6. 
Typical samples may have zones of lithophysal porosity ranging from less than 10% to greater 
than 20%. These spatially variable samples were then subjected to compression testing to define 
their mechanical properties. Two sets of tests were run using mean and lower bound properties 
assumptions (given in Figure A-7). The results of these experiments (see Figure A-9) show that 
the rock-mass compressive strength and moduli of spatially variable samples tends toward the 
mid-point of the properties ranges roughly corresponding to the mean porosity of the samples. 
The average rock strength condition tends toward an equivalent strength category of 3 to 4, 
which is also the most prevalent lithophysal porosity condition mapped underground. This 
illustrates, on a drift scale, the use of constant rock properties in models to represent that strength 
categories 1 and 2 are conservative and represent only localized regions of poor quality rock. 
Calibration of the Drift Scale UDEC Discontinuum Model to Rock Strength Categories 
The model calibration to the rock properties categories is accomplished by subjecting samples of 
the UDEC discontinuum model rock mass to uniaxial and triaxial compressive strength tests. 
During these tests, fitting of the Mohr-Coulomb strength parameters (cohesion and tensile 
June 2004 A-20 No. 4: Mechanical Degradation Revision 1 
strength) and the fracture stiffness values is performed to achieve the measured modulus and 
compressive strength average and lower bound for each rock strength category (Figure A-10). 
As seen in Figure A-10, the calibration involved generation of stress-strain curves for the 
equivalent material. The calibrated UDEC discontinuum model was then validated by 
comparisons to the stability condition of the existing excavations. Field observations show good 
ground conditions (e.g., stable conditions and lack of tunnel deterioration over time even though 
minimal ground support is used) in the Tptpll or Tptpul. There are a small number of locations 
where minor springline surface fracturing can be observed in large diameter boreholes in the 
Tptpll in the ECRB Cross-Drift. Additionally, minor roof spalling was observed under thermal 
overdrive conditions in the Tptpmn heated drift test. Both of these observations, as well as 
comparisons to laboratory failure modes, were used as validation examples in Drift Degradation 
Analysis (BSC 2004b) and are reviewed in Section 4.2.7.2. 
Source: DTN: MO0306MWDDDMIO.001. 
NOTE: Model fracture stiffness, shear, and tensile strength calibrated to reproduce laboratory values of Youngs 
Modulus and unconfined compressive strength. Samples show fracturing and failure modes in tension and 
unconfined and confined compression. 
Figure A-10 Example of the Calibration of the UDEC Discontinuum Block Model to Laboratory Testing 
Results in Unconfined and Confined Compression and Tension, Strength Category 5 
Consideration of Long Trace Length Fracturing in the TptpllThe discussion heretofore has 
centered on the role of lithophysal porosity and the ubiquitous short-length fractures as the 
primary controllers of lithophysal rock-mass behavior. There are also longer trace length 
June 2004 A-21 No. 4: Mechanical Degradation Revision 1 
fractures (length greater than 1 m) that are fairly widely spaced. Drift Degradation Analysis 
(BSC 2004b, Section 6.4) provides an analysis of potential failure modes involving these longer 
fractures. A simulated network of these fractures was developed using the FracMan program 
(USGS 1999) based on detail line survey mapping as described in this technical basis document. 
A series of 76 three-dimensional 3DEC models of these longer fractures in the Tptpll were 
analyzed to determine if any kinematically admissible wedges were possible. Even with the 
unrealistic use of zero cohesion and friction, there were no moveable wedges formed by these 
structures in the Tptpll. Therefore, the assumption that rock-mass strength controlled by 
lithophysal porosity and the ubiquitous, small scale fracturing is valid. 
A.4.6 Conclusions 
RDTME 3.05 deals with the strategy for developing a methodology to determine rock-mass 
mechanical properties and their variability for lithophysal rocks. It was determined that reliance 
on empirical rock-mass classification methods for estimating lithophysal rock properties was not 
appropriate at this time as insufficient excavation experience in this rock is available upon which 
to develop property correlations. The approach taken to establish lithophysal rock-mass 
properties and constitutive models was based on site-specific geological mapping, laboratory and 
field testing, and numerical model calibration. The testing and material property extrapolations 
accomplished utilizing numerical modeling as a tool were used to provide an estimate of the 
mean and range of rock-mass parameter values as a function of the lithophysal porosity. The 
porosity has been determined to be the primary factor in controlling the rock-mass mechanical 
properties, with significant impact on rock-mass strength and deformation modulus. 
For parametric evaluation of rock-mass behavior, it is appropriate to subdivide the range of 
rock-mass porosities into a number of categories that allow direct relation of in situ geologic 
mapping to the associated rock-mass properties. Estimation of rock-mass properties at the drift 
scale and the impact of spatial variability of lithophysal porosity was performed based on 
geologic mapping and laboratory testing. Inclusion of spatial variability has shown that 
rock-mass properties tend toward the center of the range used in parametric stability calculations. 
This result is expected as the rock-mass properties tend toward that defined by the average 
lithophysal porosity. Thus, the use of parametric stability analysis based on constant rock-mass 
properties allows conservatism in the analyses to be assured. Numerical analysis of drift 
degradation under the influence of in situ, thermal, and seismic loading is accomplished by using 
a parametric approach for all rock-mass quality categories. This strategy and the results obtained 
are considered responsive to this KTI agreement. 
A.5 REFERENCES 
A.5.1 Documents Cited 
Board, M. 2003. Resolution Strategy for Geomechanically Related Repository Design and 
Thermal-Mechanical Effects (RDTME). REV 00. Las Vegas, Nevada: Bechtel SAIC Company. 
ACC: MOL.20030708.0153. 
BSC (Bechtel SAIC Company) 2002a. Software Code: PFC2D. V2.0. PC WINDOWS 
2000/NT 4.0. 10828-2.0-00. 
June 2004 A-22 No. 4: Mechanical Degradation Revision 1 
BSC 2002b. Software Code: PFC3D. V.2.0. PC. 10830 2.0-00. 
BSC 2002c. Software Code: UDEC. V3.1. PC WINDOWS 2000/NT 4.0. 10173-3.1-00. 
BSC 2003a. Scoping Analysis on Sensitivity and Uncertainty of Emplacement Drift Stability. 
800-KOC-TEG0-00600-000-000. Las Vegas, Nevada: Bechtel SAIC Company. 
ACC: ENG.20031125.0002. 
BSC 2003b. Subsurface Geotechnical Parameters Report. 800-K0C-WIS0-00400-000-00A. 
Las Vegas, Nevada: Bechtel SAIC Company. ACC: ENG.20040108.0001. 
BSC 2004a. Lithophysal Rock Mass Mechanical Properties of the Repository Host Horizon. 
800-K0C-SS00-00200-000-00A. Las Vegas, Nevada: Bechtel SAIC Company. 
ACC: MOL.20040510.0200. 
BSC 2004b. Drift Degradation Analysis. ANL-EBS-MD-000027 REV 03A. Las Vegas, NV: 
Bechtel SAIC Company. ACC: MOL.20040513.0081. 
BSC 2004c. Evaluation of Emplacement Drift Stability for KTI Resolutions. 800-KMC-SSE0- 
00200-000-00A. Las Vegas, Nevada: Bechtel SAIC Company. ACC: MOL.20040510.0199. 
NRC (U.S. Nuclear Regulatory Commission) 2002. Integrated Issue Resolution Status Report. 
NUREG-1762. Washington, D.C.: U.S. Nuclear Regulatory Commission, Office of Nuclear 
Material Safety and Safeguards. TIC: 253064. 
Reamer, C.W. and Williams, D.R. 2001. Summary Highlights of NRC/DOE Technical 
Exchange and Management Meeting on Repository Design and Thermal-Mechanical Effects. 
Meeting held February 6-8, 2001, Las Vegas, Nevada. Washington, D.C.: U.S. Nuclear 
Regulatory Commission. ACC: MOL.20010307.0511 through MOL.20010307.0521. 
USGS (U.S. Geological Survey) 1999. Software Code: FracMAN. V.2.511. PC Windows NT. 
10114-2.511-00. 
A.5.2 Data, Listed by Data Tracking Number 
MO0306MWDDDMIO.001. Drift Degradation Model Inputs and Outputs. Submittal date: 
06/19/2003. 
A-23 June 2004 No. 4: Mechanical Degradation INTENTIONALLY LEFT BLANK 
A-24 No. 4: Mechanical Degradation 
Revision 1 
June 2004 APPENDIX B 
SCOPING ANALYSIS OF INPUT DATA AND SENSITIVITY AND UNCERTAINTY 
ANALYSIS OF EMPLACEMENT DRIFT STABILITY 
(RESPONSE TO RDTME 3.06) 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 Revision 1 
Note Regarding the Status of Supporting Technical Information 
This document was prepared using the most current information available at the time of its development. This 
Technical Basis Document and its appendices providing Key Technical Issue Agreement responses that were 
prepared using preliminary or draft information reflect the status of the Yucca Mountain Projects scientific 
and design bases at the time of submittal. In some cases this involved the use of draft Analysis and Model 
Reports (AMRs) and other draft references whose contents may change with time. Information that evolves 
through subsequent revisions of the AMRs and other references will be reflected in the License Application 
(LA) as the approved analyses of record at the time of LA submittal. Consequently, the Project will not 
routinely update either this Technical Basis Document or its Key Technical Issue Agreement appendices to 
reflect changes in the supporting references prior to submittal of the LA. 
June 2004 No. 4: Mechanical Degradation Revision 1 
APPENDIX B 
(RESPONSE TO RDTME 3.06) 
This appendix provides a response for Key Technical Issue (KTI) agreement Repository Design 
and Thermal-Mechanical Effects (RDTME) 3.06. This agreement relates to concerns regarding 
the design sensitivity and uncertainty of input parameters to preclosure drift stability. 
B.1.1 RDTME 3.06 
Agreement RDTME 3.06 was reached during the U.S. Nuclear Regulatory Commission 
(NRC)/U.S. Department of Energy (DOE) Technical Exchange and Management Meeting on 
Repository and Design Thermal-Mechanical Effects held February 68, 2001, in Las Vegas, 
Nevada (Reamer and Williams 2001). There has been no submittal related to this KTI agreement 
to the NRC. 
The wording of this agreement is as follows: 
RDTME 3.06 
Provide the design sensitivity and uncertainty analyses of the rock support system. 
The DOE will prepare a scoping analysis to determine the significance of the 
input parameters for review by NRC staff by August 2002. Once an agreed set of 
significant parameters has been determined by the DOE and NRC staff, the DOE 
will prepare an analysis of the sensitivity and uncertainty of the preclosure rock 
support system to design parameters in a revision to Ground Control for 
Emplacement Drifts for SR, ANL-EBS-GE-000002 (or other document) 
supporting any potential license application. This is expected to be available to 
NRC in FY 2003. 
The agreement focuses on a concern regarding the design sensitivity and uncertainty of input 
parameters to the stability of the drifts and the ground support system. The concern is further 
elaborated in the Integrated Issue Resolution Status Report (NRC 2002, Section 2.1.7). The 
NRC concerns are paraphrased as follows: 
 Mechanical-property uncertainties were not discussed in the DOE analyses of 
ground-support performance for site recommendation (CRWMS M&O 2000). There 
are considerable uncertainties in all the mechanical properties needed for design 
analyses. The influence of such uncertainties on the assessment of the performance of 
the subsurface structures, systems, and components should be clearly identified, and the 
identification should be supported with adequate technical basis (NRC 2002, 
p. 2.1.7-21). 
SCOPING ANALYSIS OF INPUT DATA AND SENSITIVITY AND UNCERTAINTY 
ANALYSIS OF EMPLACEMENT DRIFT STABILITY 
June 2004 
B.1 KEY TECHNICAL ISSUE AGREEMENT 
B-1 No. 4: Mechanical Degradation Revision 1 
 Effect of lithophysae and fractures are not adequately accounted for in rock-mass 
properties estimates used in numerical models (NRC 2002, pp. 2.1.7-14 to 2.1.7-16). 
 Other concerns on specific rock-mass properties including rock-mass Youngs modulus 
(NRC 2002, p. 2.1.7-16), rock-mass strength (NRC 2002, p. 2.1.7-17), rock-mass 
thermal expansivity (NRC 2002, p. 2.1.7-18), rock-mass thermal properties (NRC 
2002, p. 2.1.7-18), and spatial and temporal variation of mechanical properties (NRC 
2002, p. 2.1.7-19). 
B.1.2 Related Key Technical Issue Agreements 
Agreement RDTME 3.05 (Appendix A) is related to RDTME 3.06. RDTME 3.05 addresses 
rock-mass mechanical properties of lithophysal rock. The rock-mass properties estimate, based 
on the approach described in resolution of RDTME 3.05 (Appendix A), is used for resolution of 
RDTME 3.06. 
B.3 RESPONSE 
The response to RDTME 3.06 is based upon a series of parametric analyses prepared to assess 
the sensitivity and uncertainty of input parameters that govern drift stability and the performance 
of the ground support system. Although this document deals largely with postclosure drift 
degradation issues, the rock mass properties and numerical modeling approaches summarized 
here are also used to investigate preclosure drift stability and ground support performance issues. 
A description of the derivation of the basic tools used for resolution of RDTME 3.06 is given in 
B.2 RELEVANCE TO REPOSITORY PERFORMANCE 
The preclosure safety analysis demonstrates the safety of the proposed design and operations in 
the geologic repository operations area with regard to the overall preclosure performance 
objectives. The safety strategy for the preclosure operating period is to demonstrate that the 
ground control system is not required to prevent or mitigate credible rockfall. This 
demonstration relies upon analyses that show that the waste package does not breach when 
impacted by credible rock blocks. The safety analysis comprises a systematic examination of the 
site, design, potential hazards, initiating events and their resulting event sequences, and the 
potential radiological exposures to workers and the public (10 CFR 63.112). 
The emplacement drifts are an array of horizontal tunnels trending at 72 azimuth. Each drift has 
a diameter of 5.5 m and will be separated from the adjacent drifts by a center-to-center distance 
of 81 m (BSC 2003a; Williams 2002). The emplacement drifts provide the subsurface access 
and openings for the structures, systems, and components used for emplacement and retrieval 
operations. The emplacement area host rock provides shielding for the rest of the underground 
facilities from radiation emanating from the waste packages. 
The rock-mass surrounding the emplacement drifts will be subjected to loadings from in situ 
stress, thermal loading, and seismic ground motion. Both the design of the emplacement drifts 
and the surrounding rock mass are relevant for determining emplacement drift stability and 
repository performance. 
B-2 June 2004 No. 4: Mechanical Degradation Revision 1 
Sections 3.2 and 4.3. The use of continuum- and discontinuum-based numerical modeling 
approaches for analysis of lithophysal and nonlithophysal units is described in Section 4.2.2.2 
and in Appendix D. The analyses summarized below are described in detail in Scoping Analysis 
on Sensitivity and Uncertainty of Emplacement Drift Stability (BSC 2003b). 
The items summarized in this appendix for resolution of RDTME 3.06 include: 
 Identification of input parameters 
 Assessment of drift stability, considering the sensitivity and uncertainty of the 
following rock-mass mechanical properties and related parameters 
 The effect of lithophysae on the lithophysal rock-mass properties and the associated 
uncertainties 
 The effect of fractures on the nonlithophysal rock-mass properties and the 
associated uncertainties 
 Spatial variation of rock-mass mechanical properties 
 Consideration of the excavation-disturbed zone 
 Consideration of the range of rock strength within each rock category 
 Conduct analysis to address the sensitivity and uncertainty of thermal modeling-related 
parameters to drift stability 
 Uncertainties associated with thermal properties 
 Off-normal thermal scenarios 
 Waste emplacement sequence and repository edge effect. 
A summary of the scoping analysis in response to these items is provided below: 
 Unsupported emplacement drifts are expected to be stable, with relatively small rock 
displacements and minor yield zones induced by excavation for drifts located in both 
lithophysal units and nonlithophysal units. Thermal and seismic loads during the 
preclosure period are small in comparison to the rock mass strength (see Section 5.3.2). 
With the exception of minor yielding in the immediate periphery of emplacement drifts 
for the poorest lithophysal rock mass strength category, elastic rock mass response is 
predicted for the in situ, thermal, and seismic loading stresses. The resulting safety 
factor against yield for unsupported emplacement drifts in the preclosure period is at 
least 2 (BSC 2003c, Section 6.1.3; BSC 2003b, Sections 6.1.1 and 6.1.2). An example 
of contours of a factor of safety is given in Figure B-1. The ground support system is, 
therefore, not required for overall drift stability; the main function of the ground 
support is to provide retainment of loose and detached rock particles around the 
periphery of the excavation. 
June 2004 B-3 No. 4: Mechanical Degradation Revision 1 
 As discussed in Section 4.2, lithophysal rock mass models assume a range of 
homogeneous (constant) rock mass properties. In reality, the rock mass properties are 
variable, based on the variability of the lithophysal porosity. The suitability of the 
homogeneous property versus spatially variable property approach to drift stability 
assessment is examined. Assessment of the spatial variation of the rock mass 
properties in lithophysal rock is based on the simulated in situ lithophysal porosity 
variability, as discussed in Section 2.3.2. The overall rock mass response, assuming in 
situ variability of strength and modulus around the drifts, is predicted to be 
approximately equivalent to the response predicted assuming a constant, homogeneous 
rock mass characterized by the median rock mass quality for lithophysal rock 
(e.g., rock strength Category 3). This same conclusion was reached in regard to 
postclosure seismic damage assessment (Section 5.3.2.2.2.1). The conclusions from 
these analyses is that the results of homogeneous property parametric analyses for 
bounding rock properties ranges are reasonable and provide conservative results (BSC 
2003b, Section 6.3.1). 
 The analysis considers the effect of an excavation-disturbed zone. This disturbed zone 
is a thin zone of more heavily fractured rock around the excavation periphery due to 
mining or stress-induced yield. Rock mass mechanical properties to describe the 
disturbed zone are derived from in situ slot testing that was conducted in the Tptpll and 
Tptpul disturbed zone in the ESF (see Section B.4.1.4 of Appendix B and Section 3.2.2 
of the technical basis document). The inclusion of this more fractured rock mass 
produces lower stress concentration at the wall and higher deformation around the 
opening, but the unsupported drift remains stable (BSC 2003b, Section 6.3.3.1). This 
conclusion is confirmed by observations in the ESF tunnel of stable 7.62-m-diameter 
tunnels in which walls are only occasionally supported with rock bolts. 
 The effect of strength variation within a fixed strength category in the lithophysal rock 
is analyzed considering the bounding range of rock strength presented in Section A.4.5 
in Appendix A. The results indicate that, even with the lower-bound strength, the 
emplacement drifts are predicted to be stable (BSC 2004a, Section 6.2). 
 Sensitivity of rock mass strength parameters (cohesion and friction angle) for the 
nonlithophysal rock was assessed for estimated rock mass properties and the 
Hoek-Brown failure criterion (Section 4.3) (Hoek et al. 2002). The results show that 
the strength properties have little impact on the outcome of the analysis results (BSC 
2003b, Section 6.3.4). 
 Values of thermal conductivity and specific heat with one standard deviation more and 
less than the mean values are used as an upper bound and a lower bound while 
evaluating thermal property uncertainties. The peak temperature values at the drift 
crown are about 5C different from that of the base case due to the variation of 
thermal conductivity. An additional 1.5C is added to this peak temperature due to 
the heat capacity changes (BSC 2003b, Section 6.4.2). 
 Off-normal thermal scenarios considered various ventilation shutdown durations at 
different preclosure times. The results demonstrated a rapid temperature increase of 
June 2004 B-4 No. 4: Mechanical Degradation Revision 1 
14C in the 1-week shutdown cases and less than 2C increase in the 1-day shutdown 
case. The temperature jump diminished rapidly, with temperature returning to the 
temperature history unaffected by the shutdown, after the ventilation was restored. An 
extreme case with 1-month shutdown was also performed. This case shows a rapid 
temperature increase of 28C and a relatively slow decrease of temperature after the 
ventilation was restored. Analysis of drift stability for a preclosure seismic event 
(10-4 annual exceedance frequency), combined with the off-normal thermal scenario, 
show negligible impact on stability (BSC 2003b, Section 6.4.3). 
 Effects of the waste emplacement sequence were investigated using a two-drift NUFT 
preclosure calculation. The results of the emplacement sequence calculations exhibited 
minor temperature changes in the first and the second drifts from the base case NUFT 
calculation (BSC 2003b, Section 6.4.4). 
The information in this report is responsive to agreement RDTME 3.06 made between the DOE 
and NRC. This report contains the information that DOE considers necessary for NRC review 
for closure of this agreement. 
June 2004 B-5 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2003b, Figure 6-11. 
NOTE: Areas marked with an x in the figure exhibited yielding during the simulation process but equilibrated to an 
elastic state at the end of the simulation. RMC = rock mass category. 
Figure B-1. Yield Zone and Safety Factor Contours after Seismic Shaking, Lithophysal Rock 
June 2004 B-6 No. 4: Mechanical Degradation Revision 1 
B.4 BASIS FOR THE RESPONSE 
A two-dimensional plane-strain thermal-mechanical parameter analysis is used to assess the 
stability of unsupported emplacement drifts. The two-dimensional finite-difference code FLAC 
(Itasca Consulting Group 2002) is used for the analysis. A combination of in situ, thermal, and 
seismic loadings is included in the analysis. The continuum-based analyses assume the rock 
mass conforms to a Mohr-Coulomb failure criteria, with property variations for lithophysal and 
nonlithophysal rock described in Sections 4.2 and 4.3, respectively (BSC 2003b, Section 6.1). 
B.4.1 Analysis to Address the Sensitivity and Uncertainty of Rock-Mass Mechanical 
Property-Related Parameters to Drift Stability 
A number of parametric studies are summarized in the following sections to highlight the 
sensitivity and uncertainty aspects. Detailed results of the analyses are presented in Scoping 
Analysis on Sensitivity and Uncertainty of Emplacement Drift Stability (BSC 2003b). 
B.4.1.1 Effect of Lithophysae on Lithophysal Rock-Mass Properties and the Associated 
Uncertainties 
The effect of lithophysae on lithophysal rock-mass mechanical properties (i.e., strength and 
modulus) has been addressed in the RDTME 3.05 resolution (Appendix A). A range of 
rock-mass mechanical properties has been estimated that represents the range of rock mass 
lithophysal porosity variability present in the Tptpll and Tptpul (e.g., Appendix A, Figure A-9). 
The sensitivity of the parameter ranges was investigated in the scoping analysis with 
consideration of in situ stress, preclosure thermal loading, and seismic ground motion (BSC 
2003b, Section 6.1.1). Figure B-1 showed the yield zone and contours of safety factor against 
yield after seismic shaking with 5  10-4 ground motion for the best quality lithophysal rock 
(Category 5), the median quality lithophysal rock (Category 3), and the poorest quality 
lithophysal rock (Category 1). Minor yielding at the drift perimeter is observed for the poorest 
quality rock only. Stress paths1 for locations around the opening surrounded by the median 
quality rock, as predicted during thermal loading and seismic shaking, are presented in 
Figures B-2 and B-3. It is evident that the stress states are well below the yield surface, and the 
stress-strain response is in the elastic (nonfailed) regime. 
B.4.1.2 Effect of Fractures on Nonlithophysal Rock-Mass Properties and the Associated 
Uncertainties 
The effect of fractures on nonlithophysal rock-mass properties has been included in this analysis 
based on the conventional approach for jointed rock medium by Hoek et al. (2002). A range of 
rock-mass mechanical properties has been estimated based on this conventional approach, which 
correlates the rock-mass mechanical properties to rock quality indices derived from field 
geotechnical characterization and laboratory testing. The sensitivity of the parameter range was 
investigated in the scoping analysis with consideration of in situ stress, preclosure thermal 
1 Stress path refers to a history of the principal stress at a point in the rock mass as it undergoes the transient loading 
and unloading associated with heating and cooling or seismic stressing. Plotting of the transient stress path on a 
standard plot of principal stresses with superimposed Mohr-Coulomb yield criteria (e.g., Figure B-2) allows easy 
identification of the location, timing, and extent of yield of the rock mass. 
June 2004 B-7 No. 4: Mechanical Degradation Revision 1 
loading, and seismic ground motion (BSC 2003b, Sections 6.1.2 and 6.3.4). Stress paths for 
locations around the opening surrounded by median quality rock predicted during thermal 
loading and seismic shaking are presented in Figures B-4 and B-5. The yield criteria shown in 
these plots are based on a conservative estimate of Mohr-Coulomb strength parameters derived 
from a tangent to a nonlinear Hoek-Brown envelope at low (5 MPa) minor principal stress. This 
results in a very conservative estimate of the equivalent rock mass unconfined compressive 
strength (i.e., 21 MPa in Figures B-4 and B-5). The typical but less conservative approach to 
derivation of strength properties is to use a best-fit linear Mohr-Coulomb criteria to the 
near-linear portion of the Hoek-Brown envelope (see Hoek et al. 2002). In that case, the rock 
mass compressive strength for the median quality case is approximately 44 MPa (Section 4.3) 
(BSC 2003b), and the factor of safety against yield would be significantly larger than that 
implied by the stress paths in Figures B-4 and B-5. Even using the conservative estimate of rock 
mass strength properties, the stress states are well below the yield surface for all loading 
conditions, and the predicted stress-strain response is in the elastic (nonfailed) regime. 
Source: BSC 2003b, Figure 6-5. 
NOTE: Regions of plot beneath the strength envelope represent elastic (nonfailed) stress conditions. 
Figure B-2. Stress Path for Selected Locations during Thermal Loading, Lithophysal Rock, Median 
Quality Rock (Compression as Positive) 
June 2004 B-8 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2003b, Figure 6-12. 
NOTE: Regions of plot beneath the strength envelope represent elastic (nonfailed) stress conditions. 
Figure B-3. Stress Path for Selected Locations during Seismic Loading, Lithophysal Rock, Median 
Quality Rock (Compression as Positive) 
June 2004 B-9 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2003b, Figure 6-18. 
NOTE: Regions of plot beneath the strength envelope represent elastic (nonfailed) stress conditions. 
Figure B-4. Stress Path for Selected Locations during Thermal Loading, Nonlithophysal Rock, Median 
Rock Quality (Compression as Positive) 
June 2004 B-10 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2003b, Figure 6-21. 
NOTE: Regions of plot beneath the strength envelope represent elastic (nonfailed) stress conditions. 
Figure B-5. Stress Path for Selected Locations during Seismic Loading, Nonlithophysal Rock, Median 
Rock Quality (Compression as Positive) 
June 2004 
B.4.1.3 Spatial Variation of Rock-Mass Mechanical Properties 
Spatial variation of rock-mass mechanical properties is discussed in Sections 2.3.2 
and 5.3.2.2.2.1 and is considered in the scoping analysis report (BSC 2003b, Section 6.3.1). The 
variation of lithophysal porosities within a model cross-sectional area is simulated based on 
mapped lithophysal porosity data and the assumption that the lithophysal porosity variation 
occurs in a stratiform layering that is conformal to the dip of units. A detailed description of the 
lithophysal porosity simulation is provided in the scoping analysis (BSC 2003b, Attachment I). 
The correlation between lithophysal porosity and the strength and modulus developed from 
large-scale laboratory testing and the PFC model extrapolations (e.g., Table A-1 and Figure A-9 
of Appendix A) are used for estimating the corresponding variation of the strength and modulus 
in the analysis region. Figure B-6 shows the contours of rock mass cohesion (i.e., rock mass 
strength), illustrating the spatial variation of lithophysal porosity estimated from field panel 
mapping in the ECRB Cross-Drift. The magnitude of deformation and stresses predicted for the 
case with spatial parameter variation are approximately equivalent to similar predictions in 
which homogeneous, median rock quality category rock (Category 3) is assumed. Figure B-7 
shows the comparison of the principal stress contours between the case with spatial variation and 
the case with Category 3 rock. The Category 3 rock is, therefore, considered to be a reasonable 
mechanical representation of the in situ rock considering spatial variation. 
B-11 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2003b, Figure 6-46. 
NOTE: Cohesion is proportional to rock mass shear strength, which is, in turn, a function of lithophysal porosity. 
Regions of high cohesion are indicative of low porosity. 
Figure B-6. Contours of Rock Mass Cohesion (Shear Strength) in the Tptpll Surrounding an 
Emplacement Drift for a Representation of Spatial Variability of Lithophysal Porosity 
June 2004 B-12 No. 4: Mechanical Degradation Source: BSC 2003b, Figure 6-51. 
Figure B-7. Maximum Principal Stress Contours Comparison (a) with Assumption of Homogeneous 
Rock Mass, Category 3 Rock, and (b) with Representation of Rock Mass with 
Spatially-Variable Rock Properties 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 B-13 Revision 1 
B.4.1.4 Consideration of Excavation-Disturbed Zone in Lithophysal Rock 
An additional lithophysal rock strength category was created to account for the test results from 
the in situ slot tests, which were conducted in the disturbed zone at the ESF tunnel springline. It 
is believed that the low moduli and strength values measured near the tunnel wall are 
characteristic of the excavation-disturbed zone (BSC 2003b, Section 6.3.3.1). The low values, 
although conservative, are considered to be unrepresentative of the rock mass in the confined 
state away from the immediate tunnel periphery. A sensitivity analysis, including the effect of 
the excavation-disturbed zone represented by a 2-m ring of disturbed rock mass, was conducted. 
The modulus and strength results from the in situ slot tests (see Table 3-3 of the technical basis 
document) were used to represent the disturbed rock zone, while strength Category 1 rock was 
used to represent the surrounding rock mass in the model. The soft inclusion produces lower 
stress concentration and higher deformation. The maximum closure reaches 90 mm with 
consideration of excavation-disturbed zone compared to 55 mm of maximum closure for the case 
without the excavation-disturbed zone during thermal loading. Although the inclusion of the 
excavation-disturbed zone results in a larger yielding area and lower safety factor (BSC 2003b, 
Section 6.3.3.1), the unsupported opening remains stable. 
B.4.1.5 Consideration of the Range of Rock Strength within a Fixed Category 
The effect of strength variation within a fixed category in the lithophysal rock was analyzed 
considering the bounding range presented in Section A.4.5 and Figure A-9 in Appendix A (BSC 
2004a, Section 6.2). The relationship of unconfined compressive strength to elastic modulus for 
lithophysal rock is reproduced in Figure B-8. Both the upper and lower bounds of rock mass 
categories 1, 3, and 5 have been considered. Load combinations include in situ stress, thermal 
loading, and seismic loading. Results indicate that the drift closures and stresses in rock adjacent 
to emplacement drifts are sensitive to the variation in rock mass strength. Coupled ground 
support analyses were also conducted. Increase of drift closures and axial forces in bolts is 
observed for the lower-bound cases, while increase of stresses was found for upper-bound cases. 
The emplacement drifts are predicted to be stable even with the lower-bound strength considered 
(BSC 2004a, Section 6.2). 
B-14 June 2004 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004b, Section 6.6.3, Figures 6.6-5 and 6.6-7. 
NOTE: Base-case average properties defined for each category are the mean and the upper and lower bounds of 
each range. Category 1 is highest porosity, lowest quality rock and category 5 is lowest porosity, highest 
quality rock. Results of rock-mass strength estimates from spatial variability studies are also shown for 
mean and lower bound properties. 
Figure B-8. Lithophysal Rock Strength and Modulus Range Divided into Five Rock Strength Categories 
Covering the Large Core Laboratory Testing and PFC Extrapolation Lithophysal Shape 
Studies 
B.4.2 Analysis to Address the Sensitivity and Uncertainty of Thermal ModelingRelated 
Parameters to Drift Stability 
June 2004 
B.4.2.1 Uncertainties Associated with Thermal Properties 
Uncertainties in thermal properties, such as thermal conductivity, heat capacity, and the thermal 
expansion coefficient are evaluated by examining their effects on the performance of 
emplacement drifts (BSC 2003b, Sections 6.4.2 and 6.6.2). Values of thermal conductivity and 
specific heat with one standard deviation greater and less than the mean values are used as an 
upper bound and a lower bound for the thermal property uncertainties. Results of the thermal 
sensitivity calculations are presented in Figure B-9. The peak temperature values at the drift 
crown are about 5C different from that of the base case due to the variation of thermal 
conductivity. An additional 1.5C is added to the peak temperature due to the heat capacity 
changes. 
B-15 No. 4: Mechanical Degradation Revision 1 
In thermal-mechanical analyses with consideration of lower thermal conductivity and specific 
heat, the deformation and stress around the opening are shown to be very close to the normal 
thermal case. Similar results are also observed for cases with higher thermal expansion 
coefficient (BSC 2003b, Section 6.6.2). 
Source: BSC 2003b, Figure 6-88. 
NOTE: Waste package loading = 1.45 kW/m (linear heat source), ventilation flow rate = 15 m3/s. 
Figure B-9. Temperature at the Drift Crown as a Function of Time after Emplacement, NUFT 
Preclosure Thermal Sensitivity Calculations 
June 2004 
B.4.2.2 Uncertainties Associated with Off-Normal Thermal Scenarios 
Off-normal thermal scenarios with various ventilation shutdown durations (i.e., 1 day, 1 week, 
and 1 month) occurring at various preclosure times (e.g., 2 years, 5 years, 10 years) are 
considered in the analysis (BSC 2003b, Sections 6.4.3 and 6.6.1). The NUFT preclosure 
calculations are conducted for the off-normal thermal scenarios. The results demonstrate a rapid 
temperature increase of 14C in the 1-week shutdown case and less than 2C increase in the 
one-day shutdown case (Figure B-10). The initially-rapid increase of temperature declined 
rapidly after the normal ventilation was resumed. A case with a 1-month shutdown was also 
performed. The result shows a rapid temperature increase of 28C and a relatively slow decrease 
of temperature after the ventilation is resumed. The ventilation was assumed to be completely 
lost for 30 days in this case, which is judged to be conservative since backup fan capability and 
maintenance measures should ensure limited shutdown times. Additionally, natural air 
circulation due to the chimney effect induced by the heated air buoyancy effect and associated 
pressure gradient would occur naturally should the ventilation system be interrupted. In 
addition, loss of forced ventilation for as long as 30 days would also be very unlikely (BSC 
2004c). Therefore, this case is not considered a typical scenario in the design. 
B-16 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2003b, Figure 6-89. 
Figure B-10. Temperature at the Drift Crown for Base Case and Three Possible Off-Normal Scenarios 
The off-normal thermal scenarios are used for thermal-mechanical analyses. Predicted drift 
closures and major principal stresses in rock near the springline and the crown are shown in 
Figure B-11, under the off-normal thermal scenario 1 and 2. Compared with those predicted 
under the normal thermal condition, differences in rock displacements and stresses are not 
significant, even though the drift wall temperatures are 16C to 28C higher under the off-normal 
situations than the normal condition. Predicted yield zones and contours of strength-to-stress 
ratios also show a minor difference from the normal thermal case. 
June 2004 B-17 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2003b, Figure 6-143. 
Figure B-11. Time Histories of Drift Closures and Major Principal Stresses in Rock under In Situ and 
Thermal Loads for Various Off-Normal Thermal Conditions 
B.4.2.3 Uncertainties Associated with Waste Emplacement Sequence and Repository 
Edge Effect 
Effects of the waste emplacement sequence, in terms of time intervals of emplacement, are 
investigated using a two-drift NUFT preclosure calculation (BSC 2003b, Sections 6.4.4 
and 6.6.3). Several different emplacement time intervals were selected (e.g., 1 year, 5 years, and 
10 years) to simulate the temperature distribution and gradient compared to the simultaneous 
emplacement calculation. In addition to the waste emplacement sequence calculation, repository 
edge effect is also investigated by placing no heat source in the second drift during the entire 
preclosure period. Temperatures at the drift crown for the effects for waste emplacement 
sequence and repository edge are presented in Figures B-12 and B-13. The results of the 
emplacement sequence calculations exhibited minor temperature changes in the first and the 
June 2004 B-18 No. 4: Mechanical Degradation Revision 1 
second drifts from the normal thermal case with the temperature curves of the second drift 
shifted with delayed emplacement. 
Source: BSC 2003b, Figure 6-93. 
NOTE: Drift 1 curves are close to identical to base case. 
Figure B-12. Temperatures at the Drift Crown Resulting from Effects of the Waste Emplacement 
Sequence 
Source: BSC 2003b, Figure 6-94. 
Figure B-13. Temperatures at the Drift Crown Resulting from the Effects of the Waste Repository Edge 
A thermal-mechanical analysis, consisting of two drifts with 10 years delayed emplacement for 
the second drift, is used to investigate the impact of emplacement sequence and edge effect on 
drift stability. Time histories of predicted drift closures and major principal stresses in the rock 
near the springline and the crown of the second drift during heating for lithophysal rock strength 
June 2004 B-19 No. 4: Mechanical Degradation Revision 1 
Category 5 rock are presented in Figure B-14. A sharp increase in stress and displacement at 
around 10 years is shown both at the crown and at the springline. The magnitude of the 
maximum stress or closure is slightly less than the magnitude of the maximum stress predicted 
for the drifts located in the center of a panel. The results for the springline left and right are 
similar, indicating that the edge effect is insignificant. There are no noticeable differences in the 
predicted yield zone and safety factor contours compared to the normal thermal case. These 
results indicate that the emplacement sequence and edge effect has an insignificant impact on 
drift stability. 
June 2004 B-20 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2003b, Figure 6-152. 
NOTE: Drift closure is the net inward displacement of the drift across a given diameter. 
Figure B-14. Time Histories of Drift Closure and Major Principal Stresses in Rock under In Situ and 
Thermal Loads for the Second Drift with Consideration of Emplacement Sequence and 
Edge Effect, Rock Quality 5 
B.4.2.4 Conclusions 
In response to RDTME 3.06, a large parametric analysis of the impact of input rock mass 
properties and parameter assumptions on the prediction of unsupported emplacement drift 
stability in lithophysal and nonlithophysal rock units was performed. The sensitivity and 
June 2004 B-21 No. 4: Mechanical Degradation Revision 1 
uncertainty of unsupported drift stability was examined using a continuum-based numerical 
modeling method for the bounding ranges of mechanical and thermal rock mass properties for 
critical preclosure loading conditions, including in situ, thermal, and seismic stresses. The 
results are summarized here and described in detail in Scoping Analysis on Sensitivity and 
Uncertainty of Emplacement Drift Stability (BSC 2003b). A summary of the conclusions from 
this study are given in Section B.4.3. The analyses indicate that emplacement drifts within both 
lithophysal and nonlithophysal units are stable in an unsupported mode for all parameter 
variations and loading conditions. It is concluded that the ground support is not required to 
maintain drift stability, but that its primary purpose is retention of potentially loosened material 
along the excavation periphery. These analyses are considered to be sufficient for resolution of 
RDTME 3.06. 
B.5 REFERENCES 
B.5.1 Documents Cited 
BSC (Bechtel SAIC Company) 2003a. Underground Layout Configuration. 800-P0C-MGR0- 
00100-000-00E. Las Vegas, Nevada: Bechtel SAIC Company. ACC: ENG.20031002.0007. 
BSC 2003b. Scoping Analysis on Sensitivity and Uncertainty of Emplacement Drift Stability. 
800-K0C-TEG0-00600-000-000. Las Vegas, Nevada: Bechtel SAIC Company. ACC: 
ENG.20031125.0002. 
BSC 2003c. Ground Control for Emplacement Drifts for LA. 800-K0C-TEG0-00100-000-00A. 
Las Vegas, Nevada: Bechtel SAIC Company. ACC: ENG.20031016.0001 
BSC 2004a. Evaluation of Emplacement Drift Stability for KTI Resolutions. 800-KMC-SSE0- 
00200-000-00A. Las Vegas, Nevada: Bechtel SAIC Company. ACC: MOL.20040510.0199. 
BSC 2004b. Lithophysal Rock Mass Mechanical Properties of the Repository Host Horizon. 
800-K0C-SS00-00200-000-00Aa. Las Vegas, Nevada: Bechtel SAIC Company. ACC: 
MOL.20040510.0200. 
BSC 2004c. Subsurface Construction and Emplacement Ventilation. 800-P0C-MGR0-00200- 
000-00B. Las Vegas, Nevada: Bechtel SAIC Company. ACC: ENG.20040304.0002. 
CRWMS M&O (Civilian Radioactive Waste Management System Management & Operating 
Contractor) 2000. Ground Control for Emplacement Drifts for SR. ANL-EBS-GE-000002 
REV 00. Las Vegas, Nevada: CRWMS M&O. ACC: MOL.20000414.0875. 
Hoek, E.; Carranza-Torres, C.; and Corkum, B. 2002. Hoek-Brown Failure Criterion  2002 
Edition. 5th North American Rock Mechanics Symposium and 17th Tunnelling Association of 
Canada Conference: NARMS-TAC 2002, July 7-10, University of Toronto. Toronto, Ontario, 
Canada: Rocscience. Accessed March 17, 2003. TIC: 253954. 
Itasca Consulting Group 2002. Itasca SoftwareCutting Edge Tools for Computational 
Mechanics. Minneapolis, Minnesota: Itasca Consulting Group. TIC: 252592. 
June 2004 B-22 No. 4: Mechanical Degradation Revision 1 
NRC (U.S. Nuclear Regulatory Commission) 2002. Integrated Issue Resolution Status Report. 
NUREG-1762. Washington, D.C.: U.S. Nuclear Regulatory Commission, Office of Nuclear 
Material Safety and Safeguards. TIC: 253064. 
Reamer, C.W. and Williams, D.R. 2001. Summary Highlights of NRC/DOE Technical 
Exchange and Management Meeting on Repository Design and Thermal-Mechanical Effects. 
Meeting held February 6-8, 2001, Las Vegas, Nevada. Washington, D.C.: U.S. Nuclear 
Regulatory Commission. ACC: MOL.20010307.0511 through MOL.20010307.0521. 
Williams, N.H. 2002. Thermal Inputs for Evaluations Supporting TSPA-LA. Interoffice 
memorandum from N.H. Williams (BSC) to Distribution, September 16, 2002, 0911024159, 
with enclosures. ACC: MOL.20021008.0141. 
B.5.2 Codes, Standards, and Regulations 
10 CFR Part 63. Energy: Disposal of High-Level Radioactive Wastes in a Geologic Repository 
at Yucca Mountain, Nevada. Readily available. 
June 2004 B-23 No. 4: Mechanical Degradation INTENTIONALLY LEFT BLANK 
B-24 No. 4: Mechanical Degradation 
Revision 1 
June 2004 APPENDIX C 
ROCKFALL IN EMPLACEMENT DRIFTS 
(RESPONSE TO RDTME 3.15, RDTME 3.16, RDTME 3.17, RDTME 3.19, 
AND GEN 1.01 (COMMENT 3)) 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 Revision 1 
Note Regarding the Status of Supporting Technical Information 
This document was prepared using the most current information available at the time of its development. This 
Technical Basis Document and its appendices providing Key Technical Issue Agreement responses that were 
prepared using preliminary or draft information reflect the status of the Yucca Mountain Projects scientific 
and design bases at the time of submittal. In some cases this involved the use of draft Analysis and Model 
Reports (AMRs) and other draft references whose contents may change with time. Information that evolves 
through subsequent revisions of the AMRs and other references will be reflected in the License Application 
(LA) as the approved analyses of record at the time of LA submittal. Consequently, the Project will not 
routinely update either this Technical Basis Document or its Key Technical Issue Agreement appendices to 
reflect changes in the supporting references prior to submittal of the LA. 
June 2004 No. 4: Mechanical Degradation Revision 1 
APPENDIX C 
ROCKFALL IN EMPLACEMENT DRIFTS 
(RESPONSE TO RDTME 3.15, RDTME 3.16, RDTME 3.17, RDTME 3.19, 
AND GEN 1.01 (COMMENT 3)) 
This appendix provides a response for Key Technical Issue (KTI) agreements Repository Design 
and Thermal-Mechanical Effects (RDTME) 3.15, RDTME 3.16, RDTME 3.17, RDTME 3.19, 
and General Agreement (GEN) 1.01 (comment 3). The RDTME KTI agreements relate to 
providing the technical basis for the design, construction, and operation of the geologic 
repository operations area with respect to preclosure and postclosure performance objectives, 
taking into consideration long-term thermal-mechanical processes. 
C.1 KEY TECHNICAL ISSUE AGREEMENTS 
RDTME 3.15 
Provide field data and analysis of rock bridges between rock joints that are treated 
as cohesion in DRKBA modeling together with a technical basis for how a 
reduction in cohesion adequately accounts for thermal effects. The DOE will 
provide clarification of the approach and technical basis for how reduction in 
cohesion adequately accounts for thermal effects, including any additional 
applicable supporting data and analyses. Additionally, the adequacy of the 
cohesion reduction approach will be verified according to the approach described 
in Subissue 3, Agreement 19, of the Repository Design and Thermal-Mechanical 
Effects Technical Exchange. This will be documented in a revision to the Drift 
Degradation Analysis, ANL-EBS-MD-000027, expected to be available to NRC 
in FY 2003. 
C.1.1 RDTME 3.15, RDTME 3.16, RDTME 3.17, RDTME 3.19, and GEN 1.01 
(Comment 3) 
Agreements RDTME 3.15, RDTME 3.16, RDTME 3.17, and RDTME 3.19 were reached during 
the U.S. Nuclear Regulatory Commission (NRC)/U.S. Department of Energy (DOE) Technical 
Exchange and Management Meeting on Repository and Design Thermal-Mechanical Effects 
held February 6 to 8, 2001, in Las Vegas, Nevada (Reamer and Williams 2001). There have 
been no submittals related to these KTI agreements to the NRC. 
During the NRC/DOE Technical Exchange and Management Meeting on Thermal Operating 
Temperatures, held September 18 to 19, 2001, the NRC provided an additional comment related 
to RDTME 3.15, RDTME 3.16, RDTME 3.17, and RDTME 3.19 (Reamer and Gil 2001). This 
comment (GEN 1.01, comment 3) relates to the analysis of potential consequences of drift 
collapse. DOE provided an initial response to this comment at that meeting (Reamer and 
Gil 2001). 
The wording of these agreements and of DOEs initial response to the general agreement 
comment is as follows: 
June 2004 C-1 No. 4: Mechanical Degradation Revision 1 
RDTME 3.16 
Provide a technical basis for the DOE position that the method used to model joint 
planes as circular discs does not under-represent the smaller trace-length 
fractures. The DOE will analyze the available small trace-length fracture data 
from the Exploratory Studies Facility and Enhanced Characterization of the 
Repository Block, including their effect on block development. This will be 
documented in a revision to the Drift Degradation Analysis, ANL-EBS-MD- 
000027, expected to be available to NRC in FY 2003. 
RDTME 3.17 
Provide the technical basis for effective maximum rock size including 
consideration of the effect of variation of the joint dip angle. The DOE will 
provide the technical basis for effective maximum rock size including 
consideration of the effect of variation of the joint dip angle. This will be 
documented in revisions to the Drift Degradation Analysis, ANL-EBS-MD- 
000027, and the Rockfall on Drip Shield, CAL-EBS-ME-000001, expected to be 
available to NRC in FY 2003. 
RDTME 3.19 
The acceptability of the process models that determine whether rockfall can be 
screened out from performance assessment abstractions needs to be substantiated 
by the DOE by doing the following: (1) provide revised DRKBA analyses using 
appropriate range of strength properties for rock joints from the Design Analysis 
Parameters Report, accounting for their long-term degradation; (2) provide an 
analysis of block sizes based on the full distribution of joint trace length data from 
the Fracture Geometry Analysis Report for the Stratigraphic Units of the 
Repository Host Horizon, including small joints trace lengths; (3) verify the 
results of the revised DRKBA analyses using: (a) appropriate boundary 
conditions for thermal and seismic loading; (b) critical fracture patterns from the 
DRKBA Monte Carlo simulations (at least two patterns for each rock unit); 
(c) thermal and mechanical properties for rock blocks and joints from the Design 
Analysis Parameters Report; (d) long-term degradation of rock block and joint 
strength parameters; and (e) site-specific ground motion time histories appropriate 
for postclosure period; provide a detailed documentation of the analyses results; 
and (4) in view of the uncertainties related to the rockfall analyses and the 
importance of the outcome of the analyses to the performance of the repository, 
evaluate the impacts of rockfall in performance assessment calculations. DOE 
believes that the Drift Degradation Analysis is consistent with current 
understanding of the Yucca Mountain site and the level of detail of the design to 
date. As understanding of the site and the design evolve, DOE will: (1) provide 
revised DRKBA analyses using appropriate range of strength properties for rock 
joints from a design parameters analysis report (or other document), accounting 
for their long-term degradation; (2) provide an analysis of block sizes based on 
the full distribution of joint trace length data from the Fracture Geometry Analysis 
June 2004 C-2 No. 4: Mechanical Degradation Revision 1 
for the Stratigraphic Units of the Repository Host Horizon, ANL-EBS-GE- 
000006, supplemented by available small joint trace length data; (3) verify the 
results of the revised DRKBA analyses using: (a) appropriate boundary conditions 
for thermal and seismic loading; (b) critical fracture patterns from the DRKBA 
Monte Carlo simulations (at least two patterns for each rock unit); (c) thermal and 
mechanical properties for rock blocks and joints from a design parameters 
analysis report (or other document); (d) long-term degradation of joint strength 
parameters; and (e) site-specific ground motion time histories appropriate for 
postclosure period. This will be documented in a revision to the Drift 
Degradation Analysis, ANL-EBS-MD-000027, expected to be available to NRC 
in FY 2003. Based on the results of the analyses above and subsequent drip 
shield calculation revisions, DOE will reconsider the screening decision for 
inclusion or exclusion of rockfall in performance assessment analysis. Any 
changes to screening decisions will be documented in analyses prior to any 
potential license application. 
GEN 1.01 (Comment 3) 
None of the uncertainty and/or sensitivity analyses performed in the SSPA 
include the effects of drift collapse. Analyzing the potential consequences of drift 
collapse should be done to satisfy the basic TSPAI alternative conceptual model 
requirement. 
DOE Initial Response to GEN 1.01 (Comment 3) 
The SSPA rockfall sensitivity analyses (Sect. 6.3.4) were limited to examining the 
potential importance of three key uncertainties on rockfall. These uncertainties 
included 1) multiplier of fracture trace lengths, 2) Terzaghi correction factor, and 
3) number of Monte Carlo simulations. These subsystem analyses did not 
substantially change the results of the rockfall model from that presented in the 
SR wherein we concluded rockfall did not significantly impact performance. As 
efforts were focused on other aspects of EBS performance, DOE did not perform 
system level SSPA Volume 2 calculations that included rockfall. DOE is 
continuing to do uncertainty analyses and examining an alternative model to 
improve the basis for screening rockfall from performance assessment 
abstractions per KTI agreements RDTME 3.15, RDTME 3.16, RDTME 3.17, and 
RDTME 3.19. If it is determined from these additional analyses that rockfall may 
significantly impact repository performance then, rockfall will be evaluated for 
abstraction into the TSPA calculations for any potential LA. 
The foregoing statements focus on concerns regarding seismic, thermal, and time-dependent 
effects on rockfall and drift degradation. The effect on rockfall is further elaborated in 
Integrated Issue Resolution Status Report (NRC 2002, Section 3.3.2.4.4.1), where the NRC 
comments: 
 The rationale for the DOE approach is that the additional shear stress induced on 
fracture surfaces from a temperature distribution (thermal loading) or earthquake 
C-3 June 2004 No. 4: Mechanical Degradation Revision 1 
(seismic loading) and the weakening of fracture surfaces by time-dependent 
degradation can all be represented by the specified reduction of the cohesion and 
friction-angle parameters. DOE did not present a satisfactory mathematical basis 
to relate the cohesion reduction to the temperature distribution or the 
friction-angle reduction to the seismic loading to support an argument that the 
applied fracture-strength reductions appropriately represent the thermal and 
seismic loadings for the repository (NRC 2002, p. 3.3.2-29). 
 The seismic data used for the drift degradation analysis were the design basis 
seismic ground motions for both Category 1 and 2 events. These seismic ground 
motion parameters are appropriate for preclosure-related design and analysis but 
are not proper for any postclosure considerations (NRC 2002, pp. 3.3.2-30 to 
3.3.2-31). 
C.1.2 Related Key Technical Issues 
These RDTME agreements are related to: 
CLST 2.02This agreement requires rockfall loading estimates for the drip shield (postclosure) 
and waste package (preclosure) performance. RDTME KTI agreements 3.15, 3.16, 3.17, and 
3.19 address the estimation of rockfall (mass, shape, and momentum) induced by in situ, thermal, 
and seismic loading as well as time-dependent strength changes. Section 5 summarizes the 
seismic rockfall analyses for lithophysal and nonlithophysal rocks. 
CLST 2.08This agreement requires estimation of rockfall for examination of drip shield and 
waste package performance. RDTME KTI agreements 3.15, 3.16, 3.17, and 3.19 address the 
estimation of rockfall (mass, shape, and momentum) induced by in situ, thermal, and seismic 
loading as well as time-dependent strength changes. Section 5 summarizes the seismic rockfall 
analyses for lithophysal and nonlithophysal rocks. 
CLST 2.09This agreement requires examination of the drip shield and waste package 
mechanical response under seismic excitation. RDTME KTI agreements 3.15, 3.16, 3.17, and 
3.19 deal with the estimation of rockfall (mass, shape, and momentum) induced by in situ, 
thermal, and seismic loading as well as time-dependent strength changes. Section 5 summarizes 
the seismic rockfall analyses for lithophysal and nonlithophysal rocks. 
RDTME 3.08This agreement requires sensitivity analysis to establish design uncertainty for 
ground support design and drift degradation estimates with respect to the variability of the 
fracture system in the nonlithophysal and lithophysal rocks. RDTME KTI agreements 3.15, 
3.16, 3.17, and 3.19 are linked to 3.08 primarily through determination of rockfall from in situ, 
thermal, and seismic loading as well as through long-term strength changes in the rock mass. 
This document addresses the methodology for analyzing the fracture patterns in Sections 4.1 and 
4.2, and the postclosure drift degradation analyses in Section 5. Analysis of the variability of 
fracturing and its impact on ground support design is discussed in Ground Control for 
Emplacement Drifts for LA (BSC 2003). 
June 2004 C-4 No. 4: Mechanical Degradation Revision 1 
SDS 2.04This agreement requires a seismic risk assessment be conducted to provide the impact 
of site-specific ground motion on engineered barrier performance. This agreement is linked to 
the RDTME agreements in this appendix via the conduct of rockfall analyses. Section 5 
summarizes the seismic rockfall analyses for lithophysal and nonlithophysal rocks. 
TSPAI 2.02RDTME 3.17 is linked to TSPAI 2.02, Items 62, 78, and 79. RDTME 3.19 is 
related to TSPAI 2.02, items 62, 78, and 79. TSPAI 2.02 Items 58 and 62 related to the inclusion 
of rockfall and its potential mechanical impacts on engineered barriers (Item 58) and on the 
thermal-mechanical impacts of long term rock mass degradation on engineered barriers and 
potential hydrological changes in the rock mass. RDTME 3.05 deals specifically with the 
estimation of rock mass properties of the lithophysal rocks of the Topopah Spring formation and 
the development of rock mass material and numerical models for representing fracture, rockfall, 
and long-term degradation under in situ, thermal, and seismic loading. The estimates made for 
rockfall and long-term degradation and change of opening shape feed performance assessment 
studies of the engineered barriers. (Refer to Section 1 for further details on this integration.) 
C.3 RESPONSE 
Because RDTME 3.15, RDTME 3.16, RDTME 3.17, and RDTME 3.19 are related to rockfall in 
emplacement drifts, the response to each agreement is combined in this appendix. The approach 
to the analysis of seismic, thermal, and time-dependent effects on rockfall and drift degradation 
are summarized in this appendix. This information is further described in Drift Degradation 
Analysis (BSC 2004a) and summarized in Section 5. The drift degradation analysis includes the 
development and validation of rockfall models that approximate phenomena associated with 
various components of rock mass behavior anticipated within the repository horizon. 
The subject KTI agreements request greater detail in the use of the DRKBA rockfall model 
keyblock approach to estimating rockfall. Since the time of writing of these agreements, the 
approach to estimating drift degradation has changed, as presented in March 2002 and 
documented by Board (2003) and reviewed in DOE/NRC Appendix 7 and Technical Exchange 
C.2 RELEVANCE TO REPOSITORY PERFORMANCE 
Over time, changes will occur to both the stress condition and the strength of the rock mass as a 
result of thermal stress, seismic ground motion, and time-dependent strength degradation of the 
rock mass. These effects may cause rock blocks to become detached or loosen from the rock 
mass surrounding the emplacement drifts and eventually fall into the drift as ground support 
functionality diminishes during postclosure. Potential rockfall is a concern that could affect 
waste package and drip shield performance. In the preclosure period, although the amount of 
rockfall could impact ventilation and waste package surface conditions, there are no credible 
rock blocks, with or without ground support, that can cause breach of an emplaced waste 
package. During preclosure, as rockfall occurs, the size and shape of the emplacement drift will 
change. Changes in local drift geometry could affect percolation of groundwater around the 
drifts and create the potential for water seepage into drifts. Substantial postclosure rockfall could 
also impact waste package temperature due to the insulating effect of coarse granular material 
surrounding the drip shield. The analyses of drift degradation and rockfall are relevant for 
determining emplacement drift stability and repository performance. 
C-5 June 2004 No. 4: Mechanical Degradation Revision 1 
(Stablein and Gil 2003) meetings. The current approach to rockfall analyses no longer relies on 
the DRKBA numerical code. The limitations associated with DRKBA have been addressed 
through the use of the distinct element codes UDEC (BSC 2002a), a two-dimensional 
discontinuum code, and 3DEC (BSC 2002b), a three-dimensional discontinuum code. These 
codes can explicitly apply both seismic and thermal loads (as discussed in Section 5.3.2), which 
satisfies the requirements of RDTME 3.15, RDTME 3.16, RDTME 3.17, and RDTME 3.19 for 
the following reasons: 
First, the rock mass at the repository host horizon has been geologically characterized to support 
the rockfall modeling activities (see Section 2) and input requirements for UDEC and 3DEC. 
Drift degradation models have been developed for both nonlithophysal and lithophysal rock. A 
detailed description of the rock mass characteristics of the repository host horizon is provided in 
Drift Degradation Analysis (BSC 2004a, Section 6.1). The available rock mass geotechnical 
data, including fracture geometry (BSC 2004a, Sections 6.1.4.1, 6.1.6, and Appendix B), 
lithophysal abundance and geometric characteristics (BSC 2004a, Section 6.1.4.2 and Appendix 
O), and geotechnical rock properties (BSC 2004a, Section 6.1.3 and Appendix E) are sufficient 
to support drift degradation analyses using both continuum and discontinuum approaches. 
Second, the drift-scale temperature history is calculated throughout the preclosure and 
postclosure periods of the repository (see Section 5.3.2) (BSC 2004a, Section 6.2). The 
temperature history is used to calculate the thermal stress state that develops within the rock 
mass due to the heat energy released from the stored high-level radioactive waste and appropriate 
boundary conditions for thermal loading have been applied (BSC 2004a, Sections 6.2, 6.3.1.3, 
6.4.1.2, and Appendix W). Appropriate thermal properties have been used in the 
thermal-mechanical calculation (Section 3.2.5; BSC 2004a, Sections 4.1 and E5). 
Third, small trace-length fracture data have been analyzed, including their effect on block 
development (BSC 2004a, Section 6.3.3). The DRKBA probabilistic key-block code was used 
to analyze stability of blocks under static conditions generated from the fracture database 
including fractures of trace length greater than 1 m and including the small trace-length fracture 
database for fractures less than 1 m. About 10% more blocks with significantly smaller 
maximum volume blocks are formed when small trace length fractures are included in the 
analysis (BSC 2004a, Table 6-31). It was concluded that ignoring the small length fractures 
results in conservative analyses as larger mass blocks with higher waste package and drip shield 
impact energy are formed. 
Fourth, a nonlithophysal rockfall model was developed using 3DEC with the following features 
(see Sections 4.1 and 5.3.3): 
 Appropriate boundary conditions are provided for in situ, thermal and seismic loading 
(BSC 2004a, Sections 6.3.1.2 and 6.3.1.3). 
 Critical fracture patterns are included from multiple sampling from a synthetic rock 
mass volume that contains a realistic fracture population based on field mapping data 
(BSC 2004a, Section 6.1.6). 
June 2004 C-6 No. 4: Mechanical Degradation Revision 1 
 Appropriate thermal and mechanical properties of rock blocks and joints are used (BSC 
2004a, Appendix E). 
 Long-term degradation of joint strength parameters is considered (BSC 2004a, 
Section 6.3.1.5). 
 Site-specific ground motion time histories appropriate for both the preclosure (10-4 
annual hazard level) and the postclosure (10-5, 10-6, and 10-7 annual hazard levels) time 
period are included in the model (BSC 2004a, Section 4.1.5). Since the drifts collapse 
for all cases under the 10-7 motions, explicit modeling of the 10-8 annual hazard level 
was unnecessary. 
 The maximum rock size and shape is taken directly from the 3DEC output, which 
includes the variation in joint strike, dip, spacing, and persistence. The variation of joint 
geometry parameters is based on field mapping data from the ESF and ECRB, which 
have been input into the rockfall models. 
Finally, a lithophysal rockfall model was developed using UDEC with the following features 
(Sections 4.2 and 5.3.3): 
 Appropriate boundary conditions are provided for thermal and seismic loading (BSC 
2004a, Sections 6.4.2.2 and 6.4.2.3). 
 The rock mass is represented as an assembly of polygonal, elastic blocks in which the 
bond strength of the blocks is calibrated such that the overall mechanical behavior of the 
mass is consistent with the material model developed for the lithophysal rock (BSC 
2004a, Section 7.7.4). 
 The lithophysal rockfall model allows for the formation of stress-induced fractures 
between blocks (i.e., the formation of internal fracturing) and separation and instability 
(under the action of gravity or seismic shaking) of the rock mass around the drift (BSC 
2004a, Section 6.4). 
 Appropriate thermal and mechanical properties of rock blocks and joints are used (BSC 
2004a, Appendix E). 
 Long-term degradation of rock mass strength is considered (BSC 2004a, 
Section 6.4.2.4). 
 Site-specific ground motion time histories appropriate for both the preclosure (10-4 
annual hazard level) and the postclosure (10-6 annual hazard levels) time period are 
included in the model (BSC 2004a, Section 4.1.5). Since the 10-6 annual hazard level 
results in complete collapse of emplacement drifts in lithophysal rock, it was not 
necessary to analyze the 10-7 or 10-8 annual hazard levels. 
The information in this report is responsive to agreements RDTME 3.15, RDTME 3.16, 
RDTME 3.17, RDTME 3.19, and GEN 1.01 (Comment 3) made between the DOE and NRC. 
June 2004 C-7 No. 4: Mechanical Degradation Revision 1 
The report contains the information that DOE considers necessary for NRC review for closure of 
these agreements. 
C.4 BASIS FOR THE RESPONSE 
This section describes the strategy, technical basis, and approach for resolving the RDTME 3.15, 
3.16, 3.17, and 3.19. 
C.4.1 Overview of Resolution Strategy 
As a general approach for resolving the geomechanical issues related to the RDTME KTIs and 
addressing the associated NRC and DOE agreements, a resolution strategy was outlined in 
NOTE: USBR = U.S. Bureau of Reclamation. 
Resolution Strategy for Geomechanically-Related Repository Design and Thermal-Mechanical 
Effects (RDTME) (Board 2003), as discussed in Section 6.1. This strategy, as it relates to 
rockfall analyses associated with RDTME 3.15, 3.16, 3.17, and 3.19, is illustrated in Figures C-1 
and C-2 for nonlithophysal and lithophysal rocks, respectively. 
Source: Board 2003, Figure 32. 
June 2004 
Figure C-1. General Methodology for Rockfall Analyses in the Nonlithophysal Rocks 
C-8 No. 4: Mechanical Degradation Revision 1 
Figure C-2. General Methodology for Rockfall Analyses in the Lithophysal Rocks 
Source: Board 2003, Figure 33. 
In Drift Degradation Analysis (BSC 2004a) the analysis of seismic, thermal, and time-dependent 
effects on rockfall and drift degradation is included. It is the analysis that supports the resolution 
of these KTI agreements. 
A summary of the basis for the resolution of RDTME 3.15, RDTME 3.16, RDTME 3.17, and 
RDTME 3.19 is provided in the following section. 
C.4.2 Basis for Resolution of RDTME 3.15 
The emphasis of this agreement is to determine the adequacy of the DOE approach for 
accounting for thermal effects in the analyses of drift degradation. In a previous approach, 
thermal effects were indirectly accounted for through a reduction in joint cohesion using the 
DRKBA rockfall model. The joint cohesion reduction approach was used, since the DRKBA 
rockfall model did not have a mechanism to apply thermal stresses in the rock mass. This joint 
cohesion reduction approach is no longer used to account for thermal effects. Instead, the drift 
degradation analyses use UDEC and 3DEC to model thermal loads explicitly (BSC 2004a, 
Sections 6.2, 6.3, and 6.4), thus providing an improved thermal-mechanical model. A detailed, 
mountain-scale three-dimensional thermal analysis was performed to verify proper thermal 
boundary conditions and the impact of emplacement drift location on emplacement drift thermal 
history (BSC 2004a, Appendix C). The DRKBA approach now provides only a confirmatory 
role in the assessment of drift degradation. 
C-9 June 2004 No. 4: Mechanical Degradation Revision 1 
Thermal-mechanical modeling was performed to define drift stability under combined in situ and 
thermally induced stresses. The stability of the drift due to thermal loading was analyzed using 
both 3DEC and UDEC discontinuum models, as described in Drift Degradation Analysis (BSC 
2004a, Sections 6.3 and 6.4). For consistency with other thermal calculations performed for 
performance assessment at Yucca Mountain, the 3DEC and UDEC programs were not used to 
determine the rock mass temperatures. Instead, the evolution of the temperature field after waste 
emplacement was obtained using the hydro-thermal code NUFT, which is one of the component 
submodels of the line-averaged heat source, drift-scale, thermal-hydrologic (LDTH) model, 
described in Multiscale Thermohydrologic Model (BSC 2004b). The temperature or thermally 
induced stress fields were imported sequentially into 3DEC and UDEC for 45 time-steps after 
waste emplacement, making sure that the temperature change between two stages was relatively 
small (i.e., it did not result in a large stress change). The thermal stresses due to temperature 
changes were calculated, and the models were solved for equilibrium for the selected thermal 
times. 
The NUFT approach is two-dimensional and, thus, assumes a cross section through a series of 
infinitely long emplacement drifts. Therefore, this approach adequately represents the 
developing temperature distribution around emplacement drifts located centrally within the 
repository. 
Additional clarification and technical basis for this thermal approach are provided in 
Section 5.3.1 and in Drift Degradation Analysis (BSC 2004a, Section 6.2, Appendix C, and 
Appendix U). 
The adequacy of the current methods to account for thermal effects on drift degradation 
(i.e., using UDEC and 3DEC) has been validated in Drift Degradation Analysis (BSC 2004a, 
Sections 7.7 and 7.8). 
C.4.3 Basis for Resolution of RDTME 3.16 
The available small trace-length fracture data have been analyzed and included in Drift 
Degradation Analysis (BSC 2004a, Section 6.3.3), which documents their effect on block 
development. The probabilistic key-block code DRKBA, with its efficient key-block simulation 
algorithm, is used to assess the impact of small trace-length fractures on possible rockfalls. The 
DRKBA approach models small trace-length fractures as circular discs, which is consistent with 
the 3DEC approach for fractures greater than 1 m (see Section 4.1.1). The DKRBA rockfall 
assessment is based on static conditions by comparing the results of two cases. Case 1 provides 
the standard fracture representation, which includes multiple joint sets, with joint trace lengths 
greater than 1 m. Case 2 is identical to Case 1, except it includes an additional random joint set 
for small trace-length fracture data (i.e., less than 1 m). 
Figure C-3 presents the key-block analysis results in the format of cumulative frequency of 
occurrence for both cases. The block sizes predicted from Case 2 are, in general, similar to 
Case 1. As shown in Figure C-3, the inclusion of small trace-length fractures results in a higher 
frequency of smaller blocks (i.e., block volumes less than 0.05 m3). The maximum block 
predicted is 7.36 m3 for Case 1, without small trace-length fractures, compared to 3.25 m3 for 
Case 2, including the small trace-length fractures. The results also show that by considering the 
C-10 June 2004 No. 4: Mechanical Degradation Revision 1 
small trace-length fractures, more blocks would form. A total of 347 blocks were generated in 
Case 2 with inclusion of the small trace-length fractures, compared to 325 blocks predicted in 
Case 1 without the small trace-length fractures (BSC 2004a, Table 6-31). Approximately 10% 
more blocks are predicted when considering the small trace-length fractures. Since the inclusion 
of small trace-length fractures decreases the maximum block size with a relatively small increase 
in the number of blocks, it is concluded that small trace-length fractures have a minor impact on 
key-block development in the nonlithophysal units. 
Figure C-3. Impact of Small Trace-Length Fractures on Block Size Distribution 
Source: BSC 2004a, Figure 6-111. 
C.4.4 Basis for Resolution of RDTME 3.17 
The emphasis of this agreement is to verify the adequacy of the DOE approach in determining 
the geometry of various block sizes. In a previous approach, the geometry of blocks predicted 
was not available as direct output from the DRKBA rockfall model due to the limitation of the 
postprocessing capability of the DRKBA code. Therefore, the previous approach used 
UNWEDGE to calculate block geometry based on a systematic variation of the orientation of the 
three dominant joint sets. The approach for determining the effective maximum rock size has 
been revised in Drift Degradation Analysis (BSC 2004a, Appendix I). Varying the joint 
geometry input to UNWEDGE is no longer applied. The maximum rock size and shape is taken 
directly from the 3DEC output, which is based on stochastic variation in joint strike, dip, 
spacing, and persistence (see Section 4.1.2). The variation of joint geometry parameters is based 
on field mapping data from the ESF and the ECRB, which have been input into the rockfall 
model (BSC 2004a, Sections 6.1.6 and 6.3). 
The predicted rock blocks impacting the waste package and drip shield have many different sizes 
and shapes. Since the block geometry information is mainly used for drip shield impact 
calculations, the geometry of the largest blocks is provided in Drift Degradation Analysis (BSC 
2004a, Appendix I), showing six different views of each block with a listing of corner point 
C-11 June 2004 No. 4: Mechanical Degradation Revision 1 
coordinates. Rock block masses, including relative velocity components, impact location, and 
rockfall shapes, are supplied to three-dimensional engineering structural analysis of the drip 
shield for each set of analyses at 10-5, 10-6, and 10-7 annual exceedance frequencies. 
C.4.5 Basis for Resolution of RDTME 3.19 
Part 1This agreement addresses the inclusion of an appropriate range of geotechnical data from 
a design parameters analysis report (or other report). Drift Degradation Analysis is the approved 
report documenting the appropriate range of geotechnical data (BSC 2004a, Appendix E). In the 
current revision of Drift Degradation Analysis (BSC 2004a), the DRKBA analyses provide a 
confirmatory role only in the assessment of drift degradation. The primary analyses for 
degradation of nonlithophysal rock are provided using 3DEC (BSC 2004a, Section 6.3), while 
lithophysal rock is analyzed using UDEC (BSC 2004a, Section 6.4). An appropriate range of 
joint strength properties for nonlithophysal rock and estimated rock mass strength and moduli for 
lithophysal rocks has been applied in the drift degradation analyses, with the variability 
accounted for through parametric studies that cover a bounding range of parameters (see Section 
4.2.2.1, Table 4-2; Section 5.3.2.1.8, Table 5-6). 
Underground and surface excavations, which are designed to be stable after excavation, degrade 
with time, and some may eventually show partial or complete collapse. The main reason for 
these observations is that the strength of a rock mass exposed to humidity and temperature of the 
open atmosphere decays with time when it is loaded to a stress level higher than 50% to 60% of 
its short-term strength. The rate of strength decay depends on, among other parameters, rock 
type, stress state, relative humidity, and temperature. Stress corrosion is considered the main 
mechanism causing strength degradation of the rocks (Potyondy and Cundall 2001, Section 3). 
A small number of existing as well as ongoing uniaxial static fatigue tests on nonlithophysal rock 
samples from the Tptpmn form the basis for estimating the time-dependency of the intact rock. 
PFC1 (BSC 2002c), with an appropriate time-dependent strength model based on a stress 
corrosion crack growth model, was calibrated against the existing static fatigue data for 
nonlithophysal rocks and then used to extrapolate the impact of lithophysal porosity on the time 
dependency. The resulting estimate of the time-dependent change in the strength of the rock 
mass (in terms of loss of cohesive strength) is encapsulated in the drift scale UDEC model used 
for drift degradation analysis in lithophysal rocks. This work is reviewed in Section 5.3.3.2.4 
and in Drift Degradation Analysis (BSC 2004a, Appendix S). 
The drift stability impact due to the effect of rock joint degradation in nonlithophysal units is 
assessed based on a conservative estimate of the reduction of joint cohesion and friction angle 
(BSC 2004a, Section 6.3.1.5). The reduced joint strength parameters are estimated to be in the 
range of the residual state with joint cohesion reduced to zero and the joint friction angle reduced 
to 30. The reduced friction angle is a typical value for a smooth joint reported by Goodman 
(1980, p. 158) and is consistent with the direct shear test results (DTN: GS030283114222.001). 
Dilation angle is also conservatively presumed to be zero considering the asperities on fracture 
1 The PFC code used for time-dependent analysis was developed initially for analysis of time-dependent yield in 
tunnels at the Underground Research Laboratory, Atomic Energy Commission of Canada (Potyondy and Cundall 
2001). This same model, calibrated for the Topopah Spring Tuff data, is used in Drift Degradation Analysis (BSC 
2004a) for time-dependency analysis. 
C-12 June 2004 No. 4: Mechanical Degradation Revision 1 
surfaces had been sheared off, resulting in greater rockfall. The degraded joint strength and 
dilatational properties were applied in 10-6 seismic motion cases (BSC 2004a, Table 6-21). 
While a slight increase in rockfall is predicted for the degraded state, joint strength degradation 
has a minor impact on drift stability in nonlithophysal rock. 
Predictions of the time-dependent drift degradation in lithophysal rock were examined using the 
UDEC model described in Section C.3. The time-related change in drift shape and rockfall 
associated with strength reduction in the rock mass work is reviewed in Section 5.3.3.2.4 and in 
Drift Degradation Analysis (BSC 2004a, Appendix S). The subsequent load of the broken rock 
on the drip shield is estimated using three different approaches: analytical, continuum numerical 
modeling, and discontinuum numerical modeling (BSC 2004a, Section 6.4.2.5). Cohesion and 
tensile strength of the rock mass are considered to degrade to zero in the degradation model. 
Each of the methods uses certain conditions regarding caving of the rock above the drifts and 
transfer of the stresses within the broken rock mass. Those conditions make the model results 
(i.e., cave size and pressures on the drip shield) conservative in each of three approaches (i.e., the 
conditions result in higher pressures). The level of conservatism is the largest in the analytical 
results and the smallest in the approach that represents rock mass as a discontinuum. 
The predictions of pressure of the caved rock on the drip shield by the three modeling 
approaches are summarized in Figure C-4 (BSC 2004a, Section 6.4.2.5). The analytical model 
yields the largest loads due to overly conservative conditions. The continuum numerical model 
accounts more accurately for transfer of load by friction from the caved rock to the surrounding 
stable rock mass. Consequently, predicted loads for small bulking factors and large cavity size 
are much smaller than analytical predictions. When the bulking factor is large, the height of the 
cave becomes small. Stress arching cannot be realized within the small column of the caved 
rock and, consequently, predictions between analytical and continuum models are identical. The 
most accurate approach, using the discontinuum model, does not use an imposed condition about 
the shape of the caved region. It also correctly accounts for load transfer through the caved rock. 
The predictions of the pressures on the drip shield using this approach are smaller than the 
predictions of the analytical and continuum models for all values of the bulking factor. 
Part 2An analysis of block sizes based on the full distribution of joint trace length data has 
been included in the Drift Degradation Analysis (BSC 2004a, Sections 6.1.4 and 6.1.6). The 
joint trace length data used are consistent with the data from Fracture Geometry Analysis for the 
Stratigraphic Units of the Repository Host Horizon (CRWMS M&O 2000), and are discussed in 
Section 4.1. 
To improve the method for estimating rockfall in the repository host horizon, a fracture network 
texture representation has been developed for the four subunits comprising the repository host 
horizon (BSC 2004a, Section 6.1.6 and Appendix B). The fracture network texture 
representation provides the basis for geologically representative drift degradation scenarios. The 
synthetic fracture network in three dimensions has been constructed using the FracMan 
methodology. The actual data from the ESF main loop and ECRB Cross-Drift are used to 
condition the fracture network texture representation. These data are detailed line survey and 
full periphery geologic map data collected from the tunnel walls during and after construction. 
The fracture data include strike and dip, trace length, truncation style, and intensity/density of 
fracturing. 
June 2004 C-13 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 6-170. 
NOTE: Analytical methods are shown by solid lines. Continuum numerical modeling methods include Piping failure 
mechanism  numerical and Terzaghi failure mechanism  numerical. All other points are determined 
using a discontinuum numerical modeling approach. 
Figure C-4. Predictions of Pressure of the Degraded Lithophysal Rock on the Drip Shield as a Function 
of Bulking Factor 
The confidence building for the development of synthetic fracture geometries is both qualitative 
and quantitative. A visual comparison to mapped field data is used qualitatively to evaluate 
synthetic fracture networks. Although qualitative, the visual comparison of a synthetic full 
periphery geologic map with the observed full periphery geologic map allows a synthetic 
fracture network to be abstracted from the observed data. Qualitative comparisons are provided 
in Part 3b of this section. 
Quantitative comparison was done through comparison of orientation distributions (synthetic 
stereonets and observed stereonets), as well as comparison of trace length distribution and 
inter-fracture distance distribution, both synthetic and observed (BSC 2004a, Section 6.1.6). 
The analysis of block sizes also includes the available small joint trace length data (see 
Section C.4.3). 
Part 3As indicated in Section C.4.2, the DRKBA results now provide a confirmatory role in the 
assessment of drift degradation. 3DEC has replaced DRKBA as the primary code for analyzing 
structural block development in the nonlithophysal rock units. The 3DEC and DRKBA results 
June 2004 C-14 No. 4: Mechanical Degradation Revision 1 
are in good agreement, since the DRKBA block sizes are generally bounded by the 3DEC results 
(BSC 2004a, Section 7.8.4). 
Part 3aAppropriate boundary conditions for thermal and seismic loading have been included in 
the rockfall models (BSC 2004a, Sections 6.3.1.1, 6.4.2.2, and 6.4.2.3). At the initial 
consolidation stage and the later thermal loading period for the analysis of nonlithophysal rock 
using 3DEC, fixed velocity boundaries were used to ensure the boundary effect does not impact 
the stress distribution around the opening. For the 3DEC seismic analysis, a nonreflecting 
boundary is used for both the top and bottom of the model, whereas a free field boundary is 
imposed at the perimeter of the model (BSC 2004a, Figure 6-42). The free field boundaries 
ensure that plane waves propagating upward suffer no distortion at the boundary. A description 
of the free-field boundary is provided in Drift Degradation Analysis (BSC 2004a, Appendix H). 
Dynamic loading was applied at the bottom of the model as a prescribed stress boundary, and 
propagated vertically upward. 
For the seismic analysis of lithophysal rock using UDEC, quiet (nonreflecting) boundaries are 
used on the outside boundaries of each model (BSC 2004a, Section 6.4.2.2). These boundaries 
prevent reflection of outgoing seismic waves back into the model, which is an essential modeling 
consideration since the reflection has been accounted for in the ground motion time histories. 
Quiet boundaries are combined with free field boundaries on the vertical outside boundaries. 
The free field boundaries perform one-dimensional simulation of vertically propagating plane 
waves representing motion of truncated, semi-infinite medium. They prevent distortion of 
vertically propagating plane waves along the quiet boundaries. Dynamic loading was applied at 
the bottom of the model, as propagating vertically upward. Although the dynamic loading was 
specified as velocity histories, it was applied at the bottom model boundary as a stress boundary 
condition. The conversion of the ground motion velocity to input seismic stress is discussed in 
Drift Degradation Analysis (BSC 2004a, Section 6.4.2.2). 
For the thermal analysis of lithophysal rock using UDEC, boundary conditions are discussed in 
Drift Degradation Analysis (BSC 2004a, Section 6.4.2.3). The UDEC model does not perform 
complete thermal-mechanical simulation. Instead, temperature fields calculated with the code 
NUFT for 1.45 kW/m and 50 years of forced ventilation are imported into UDEC. Two cases of 
ventilation efficiency are considered, as discussed in Section 5.3.1: 90% and 70% heat removal. 
Stresses are calculated for each new temperature state based on the temperature increment (from 
the previous temperature state) and the coefficient of thermal expansion. 
Part 3bThe previous DRKBA approach used a built-in joint generator for simulating joints in 
three-dimensional space. The DRKBA joint generator was limited in its ability to fully populate 
the model space, and typically created a joint density that was highest in the center of the model 
region. For the 3DEC analyses, the software FracMan (USGS 1999) provides fracture geometry 
input, which is used to provide an improved method of generating joints. 
A representative FracMan simulation of the actual fracture network has been constructed based 
on standard detailed line survey and full periphery geologic map data (BSC 2004a, 
Section 6.1.6). These data consist of fractures with trace lengths of 1 m or greater. The premise 
to this simulation is that a cube 100 m on a side results in a representative fracture network 
within which emplacement drifts can be excavated. The fractures are simulated, and their 
June 2004 C-15 No. 4: Mechanical Degradation Revision 1 
location, orientation, and size are inputs for the rockfall analyses. Individual 100 m cubes are 
constructed for each lithostratigraphic unit. For structural rockfall analyses, the Tptpll and 
Tptpmn units are representative of lithophysal and nonlithophysal rock, respectively, within the 
repository. 
For nonlithophysal rock, 50 FracMan fracture patterns have been sampled from the 100 m 
Tptpmn FracMan cube and analyzed using 3DEC. The number of fracture patterns sampled was 
determined based on a sufficiency assessment to ensure that rockfall data (including block size, 
relative impact velocity, and impact energy of the rock to the drip shield) are converging to a 
stable value (BSC 2004a, Appendix K). These FracMan fractures are drawn from the same 
fracture population used in the DRKBA analyses (BSC 2004a, Sections 6.3.1.1 and 6.4.3), which 
includes joint geometry data collected from geologic mapping in both the ESF and the ECRB 
Cross-Drift. Comparisons of full periphery geologic maps from the ECRB Cross-Drift to 
simulated full periphery geologic maps from the FracMan cube are shown in Figures C-5 and 
C-6 for the lithophysal and nonlithophysal rocks, respectively. Fracture intensity, trace length, 
and orientation are similar for both the field data and the synthetic FracMan data (see 
Section 4.1.2). 
Source: BSC 2004a, Figure 6-20. 
Figure C-5. Comparison of Full Periphery Geologic Maps from the Tptpll in the ECRB Cross-Drift (a) 
with Simulated Full Periphery Geologic Maps from the FracMan Cube (b) 
June 2004 C-16 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 6-28. 
Figure C-6. Comparison of Full Periphery Geologic Maps from the Tptpmn in the ECRB ESF to 
Simulated Full Periphery Geologic Maps from the FracMan Cube 
Part 3cThermal and mechanical properties for rock blocks and joints are available in the 
Technical Data Management System, as documented in Drift Degradation Analysis (BSC 2004a, 
Section 4.1 and Appendix E). 
Sensitivity calculations for thermal properties were conducted for a range including the mean 
value down to one standard deviation less than the mean for thermal conductivity and specific 
heat (BSC 2004a, Sections 6.2, 6.3.1.3, and 6.4.1.2). The sensitivity case results in an 
approximately 23C higher peak temperature compared with the base case with minor impact to 
the rockfall prediction. 
Based on the discussion provided in Section 3.2, a sufficient amount of intact rock physical and 
mechanical properties data have been collected for the nonlithophysal rock units. Conversely, 
the amount of intact rock physical and mechanical properties data for the lithophysal units is 
June 2004 C-17 No. 4: Mechanical Degradation Revision 1 
limited. To account for the uncertainty of intact data for the lithophysal units in the UDEC 
lithophysal rockfall model, five categories of rock properties were included in the model to 
assess the impact of the bounding ranges in intact properties data. The difference of rockfall 
prediction for the range of properties considered is provided in Drift Degradation Analysis (BSC 
2004a, Section 6.4), and in Appendix A, which address RDTME 3.05. 
Part 3dLong-term joint strength degradation in nonlithophysal rock and long-term rock mass 
strength degradation in lithophysal rock has been documented in Drift Degradation Analysis 
(BSC 2004a, Sections 6.3.1.5 and 6.4.2.4). The drift stability impact due to the effect of rock 
joint degradation in nonlithophysal units is assessed based on a conservative estimate of the 
reduction of joint cohesion and friction angle (BSC 2004a, Section 6.3.1.5). While a slight 
increase in rockfall is predicted for the degraded state, joint strength degradation has a minor 
impact on drift stability in nonlithophysal rock. 
The time-related change in drift shape and rockfall associated with strength reduction in the rock 
mass work in lithophysal rock is reviewed in Section 5.3.3.2.4 and in Drift Degradation Analysis 
(BSC 2004a, Appendix S). Cohesion and tensile strength of the lithophysal rock mass are 
considered to degrade to 0 in the degradation model, resulting in collapse of the drift. The 
subsequent load of the broken rock on the drip shield was then estimated. Additional discussion 
of long-term strength degradation is provided in Part 1 of this section. 
Part 3eSite-specific ground motion time histories were developed based on a site response 
model. The modeling approach implements a random-vibration theory, equivalent-linear 
formulation to calculate site response effects on ground motions. A detailed description for the 
development of the site specific ground motion time histories is provided in Development of 
Earthquake Ground Motion Input for Preclosure Seismic Design and Postclosure Performance 
Assessment of a Geologic Repository at Yucca Mountain, NV (BSC 2004c). 
Site-specific ground motions for three postclosure levels of annual probability of exceedance, 
10-5, 10-6, and 10-7, are included in Drift Degradation Analysis (BSC 2004a, Section 6.3.1.2.1). 
A total of 15 sets of Point B ground motion (i.e., ground motion developed at the repository 
horizon) were selected for each annual postclosure hazard level. The multiple sets ensure a 
reasonable distribution of spectral shapes and time history durations. For each set of ground 
motion, two horizontal components (H1 and H2) and one vertical component (V) of acceleration, 
velocity, and displacement are supplied. Figure C-7 shows example H1 velocity time histories 
for four annual hazard levels. A preclosure ground motion level (10-4 annual probability of 
exceedance) is included in this figure for comparison to the postclosure levels. 
These site-specific ground motions are included in both the 3DEC nonlithophysal rockfall model 
and the UDEC lithophysal rockfall model, as documented in Drift Degradation Analysis (BSC 
2004a, Sections 6.3.1.2 and 6.4.2.2). 
June 2004 C-18 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 6-43. 
C.4.6 Conclusions 
The current approach to rockfall analyses no longer relies on the DRKBA numerical code. The 
limitations associated with DRKBA have been addressed through the use of the distinct element 
codes UDEC, a two-dimensional discontinuum code, and 3DEC, a three-dimensional 
Figure C-7. Examples of Ground Velocity Time Histories with Truncated Duration for Analysis 
Part 4The impacts of rockfall on performance are discussed in Section 1.2.2.2. The rockfall 
results described in this appendix provide input to performance assessment calculations that 
evaluate rockfall impacts. The impact of seismic-induced rockfall on engineered barrier system 
components, including the drip shield, the waste package, and the fuel cladding, has been 
evaluated in Seismic Consequence Abstraction (BSC 2004d). The impact of drift shape changes 
caused by rockfall has been evaluated in seepage models for performance assessment (BSC 
2004e). These performance calculations provide probability distributions of relevant parameters 
affected by rockfall for use in the TSPA-LA. 
June 2004 C-19 No. 4: Mechanical Degradation Revision 1 
discontinuum code. These codes can explicitly apply both seismic and thermal loads, which 
satisfies the requirements of RDTME 3.15, RDTME 3.16, RDTME 3.17, and RDTME 3.19. 
Drift degradation has the potential to affect drip shield integrity directly, waste package integrity 
indirectly, and thermal-hydrologic environments within drifts. The results of this modeling and 
analysis activity provide rockfall data to support structural analyses of the ground support 
system, the drip shield, and waste package. The drift degradation analysis also provides the 
changes in drift profile due to rockfall, which supports analyses of seepage into the emplacement 
drift during the period of compliance for postclosure performance. 
C.5 REFERENCES 
Contractor) 2000. Fracture Geometry Analysis for the Stratigraphic Units of the Repository 
C.5.1 Documents Cited 
Board, M. 2003. Resolution Strategy for Geomechanically-Related Repository Design and 
Thermal-Mechanical Effects (RDTME). REV 00. Las Vegas, Nevada: Bechtel SAIC Company. 
ACC: MOL.20030708.0153. 
BSC (Bechtel SAIC Company) 2002a. Software Code: UDEC. V3.1. PC/WINDOWS 
2000/NT 4.0. 10173-3.1-00. 
BSC 2002b. Software Code: 3DEC. V2.01. PC/WINDOWS 2000/NT 4.0. 10025-2.01-00. 
BSC 2002c. Software Code: PFC2D. V2.0. PC/WINDOWS 2000/NT 4.0. 10828-2.0-00. 
BSC 2003. Ground Control for Emplacement Drifts for LA. 800-K0C-TEG0-00100-000-00A. 
Las Vegas, Nevada: Bechtel SAIC Company. ACC: ENG.20031016.0001. 
BSC 2004a. Drift Degradation Analysis. ANL-EBS-MD-000027 REV 03A. Las Vegas, NV: 
Bechtel SAIC Company. ACC: MOL.20040513.0081. 
BSC 2004b. Multiscale Thermohydrologic Model. ANL-EBS-MD-000049 REV 01. Las 
Vegas, Nevada: Bechtel SAIC Company. ACC: DOC.20040301.0004. 
BSC 2004c. Development of Earthquake Ground Motion Input for Preclosure Seismic Design 
and Postclosure Performance Assessment of a Geologic Repository at Yucca Mountain, NV with 
Errata. MDL-MGR-GS-000003 REV 00. Las Vegas, Nevada: Bechtel SAIC Company. ACC: 
DOC.20031201.0001; DOC.20040401.0004. 
BSC 2004d. Seismic Consequence Abstraction. MDL-WIS-PA-000003 REV 0, with errata. 
Las Vegas, Nevada: Bechtel SAIC Company. ACC: DOC.20030818.0006; 
DOC.20040218.0002. 
BSC 2004e. Abstraction of Drift Seepage. MDL-NBS-HS-000019 REV 00 ICN 01. Las Vegas, 
Nevada: Bechtel SAIC Company. ACC: DOC.20031112.0002; DOC.20040223.0001. 
CRWMS M&O (Civilian Radioactive Waste Management System Management and Operating 
Host Horizon. ANL-EBS-GE-000006 REV 00. Las Vegas, Nevada: CRWMS M&O. ACC: 
MOL.20000918.0286. 
June 2004 C-20 No. 4: Mechanical Degradation Revision 1 
Goodman, R.E. 1980. Introduction to Rock Mechanics. New York, New York: John Wiley & 
Sons. TIC: 218828. 
NRC (U.S. Nuclear Regulatory Commission) 2002. Integrated Issue Resolution Status Report. 
NUREG-1762. Washington, D.C.: U.S. Nuclear Regulatory Commission, Office of Nuclear 
Material Safety and Safeguards. TIC: 253064. 
Potyondy, D. and Cundall, P. 2001. The PFC Model for Rock: Predicting Rock-Mass Damage at 
the Underground Research Laboratory. Report No. 06819-REP-01200-10061-R00. Toronto, 
Ontario, Canada: Ontario Power Generation, Nuclear Waste Management Division. 
TIC: 253569. 
Reamer, C.W. and Gil, A.V. 2001. Summary Highlights of NRC/DOE Technical Exchange and 
Management Meeting of Range on Thermal Operating Temperatures, September 18-19, 2001. 
Washington, D.C.: U.S. Nuclear Regulatory Commission. ACC: MOL.20020107.0162. 
Reamer, C.W. and Williams, D.R. 2001. Summary Highlights of NRC/DOE Technical 
Exchange and Management Meeting on Repository Design and Thermal-Mechanical Effects. 
Meeting held February 6-8, 2001, Las Vegas, Nevada. Washington, D.C.: U.S. Nuclear 
Regulatory Commission. ACC: MOL.20010307.0511 through MOL.20010307.0521. 
Stablein, N.K. and Gil, A.V. 2003. Summary Highlights of NRC/DOE Technical Exchange 
Meeting on Repository Design and Thermal-Mechanical Effects Key Technical Issue. Meeting 
held May 6 to 8, 2003, Las Vegas, Nevada. Washington, D.C.: U.S. Nuclear Regulatory 
Commission. ACC: MOL.20030708.0032. 
USGS (U.S. Geological Survey) 1999. Software Code: FracMAN. V.2.511. PC/Windows NT. 
10114-2.511-00. 
C.5.2 Data, Listed by Data Tracking Number 
GS030283114222.001. Direct Shear Data from Selected Samples of the Topopah Spring Tuff. 
Submittal date: 02/20/2003. 
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C-22 No. 4: Mechanical Degradation 
Revision 1 
June 2004 APPENDIX D 
ANALYSIS OF LOAD COMBINATION AND MODELING APPROACH 
RELATED TO EVALUATION OF EMPLACEMENT DRIFT STABILITY 
(RESPONSE TO RDTME 3.02, RDTME 3.10, AND RDTME 3.13) 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 Revision 1 
Note Regarding the Status of Supporting Technical Information 
This document was prepared using the most current information available at the time of its development. This 
Technical Basis Document and its appendices providing Key Technical Issue Agreement responses that were 
prepared using preliminary or draft information reflect the status of the Yucca Mountain Projects scientific 
and design bases at the time of submittal. In some cases this involved the use of draft Analysis and Model 
Reports (AMRs) and other draft references whose contents may change with time. Information that evolves 
through subsequent revisions of the AMRs and other references will be reflected in the License Application 
(LA) as the approved analyses of record at the time of LA submittal. Consequently, the Project will not 
routinely update either this Technical Basis Document or its Key Technical Issue Agreement appendices to 
reflect changes in the supporting references prior to submittal of the LA. 
June 2004 No. 4: Mechanical Degradation Revision 1 
APPENDIX D 
D.1 KEY TECHNICAL ISSUE AGREEMENTS 
RDTME 3.02 
Provide the critical combinations of in-situ, thermal, and seismic stresses, together 
with their technical bases, and their impacts on ground support performance. The 
DOE will examine the critical combinations of in-situ, thermal, and seismic 
stresses, together with their technical bases and their impacts on preclosure 
ground support performance. These results will be documented in a revision to 
the Ground Control for Emplacement Drifts for SR, ANL-EBS-GE-000002 (or 
other document) supporting any potential license application. This is expected to 
be available to NRC in FY 2003. 
RDTME 3.10 
Provide technical basis for the assessment that two-dimensional modeling for 
emplacement drifts is considered to be adequate, considering the fact that neither 
the in-situ stress field nor the principle fracture orientation are parallel or 
perpendicular to emplacement drift orientation. The DOE will provide the 
technical bases for the modeling methods used in ground control analysis in a 
revision to the Ground Control for Emplacement Drifts for SR, ANL-EBS-GE- 
000002 (or other document) supporting any potential license application. This is 
expected to be available to NRC in FY 2003. 
ANALYSIS OF LOAD COMBINATION AND MODELING APPROACH 
RELATED TO EVALUATION OF EMPLACEMENT DRIFT STABILITY 
(RESPONSE TO RDTME 3.02, RDTME 3.10, AND RDTME 3.13) 
This appendix provides a response for Key Technical Issue (KTI) agreements Repository Design 
and Thermal-Mechanical Effects (RDTME) 3.02, 3.10, and 3.13. These agreements relate to 
issues regarding the critical load combinations and modeling approaches related to the evaluation 
of emplacement drift stability and ground support performance during the preclosure period. 
These issues are addressed together in this appendix due to their related content. 
June 2004 
D.1.1 RDTME 3.02, RDTME 3.10, and RDTME 3.13 
Agreements RDTME 3.02, RDTME 3.10, and RDTME 3.13 were reached during the 
U.S. Nuclear Regulatory Commission (NRC)/U.S. Department of Energy (DOE) Technical 
Exchange and Management Meeting on Repository Design and Thermal-Mechanical Effects 
held February 6 to 8, 2001, in Las Vegas, Nevada (Reamer and Williams 2001). There have 
been no submittals related to these KTI agreements to the NRC. 
The wording of these agreements is as follows: 
D-1 No. 4: Mechanical Degradation Revision 1 
RDTME 3.13 
Provide technical justification for boundary conditions used for continuum and 
discontinuum modeling used for underground facility design. The DOE will 
provide the technical justification for boundary conditions used in modeling for 
preclosure ground control analyses in a revision to the Ground Control for 
Emplacement Drifts for SR, ANL-EBS-GE-000002 (or other document) 
supporting any potential license application. This is expected to be available to 
NRC in FY 2003. 
These agreements focus on issues regarding the effects of critical load combinations and 
numerical modeling-related parameters on emplacement drift stability and ground support 
performance. The concerns are further elaborated in the Integrated Issue Resolution Status 
Report (NRC 2002, Section 2.1.7). The NRC concerns are paraphrased as follows: 
 DOE has set performance criteria for several structures, systems, and components that 
call for a design against the worst-case load combinations. In the stability analyses of 
emplacement drifts for site recommendation (BSC 2001), the worst-case load 
combination was assumed to be achieved by superimposing seismic loading on thermal 
loading at the time when the drift-wall temperature was close to its peak value. Since 
the effect of critical load combinations depend to a large extent on the potential failure 
modes of structures, systems, and components, several different loading combinations 
need to be considered to determine the appropriate load combinations for design (NRC 
2002, pp. 2.1.7-8 to 2.1.7-9). 
 Concern about the appropriateness of two-dimensional thermal-mechanical modeling of 
emplacement drifts arises because the in situ horizontal principal stresses and several of 
the fracture sets are oblique to the proposed drift alignment (252 azimuth, that is 
S72W) (BSC 2003a, Section 5.1.4). The ambient minimum principal stress is 
horizontal and oriented N60W-N65 W, which is 40 to 45 degrees from the 
drift-normal plane, which is the assumed orientation of the minimum principal stress for 
the two-dimensional modeling. Also, the dip direction of the subhorizontal fractures lies 
in the 40 to 60 degree range (i.e., 10 to 30 degrees from the drift orientation). Therefore, 
the two-dimensional models are not favorably oriented to detect slip on the 
subhorizontal fractures. Three-dimensional modeling may be necessary to determine the 
effects of these structural features that are oblique to the drift alignment. Other areas for 
which three-dimensional modeling may also be necessary include (1) stability of the 
turnout area, which may be subjected to a combination of vertical tension and highhorizontal 
compression; (2) effects of greater heat conduction rates through the drift 
floor because steel members in the floor (invert and pallet) that are in direct or indirect 
contact with the waste package provide a faster heat-flow path into the rock; (3) stability 
of the structural components of the invert and the interaction of the transverse beams 
with the drift wall under heated conditions; and (4) effects of ground-surface topography 
drift-parallel thermal gradients on thermal stress and, consequently, drift stability (NRC 
2002, pp. 2.1.7-12 to 2.1.7-13). 
June 2004 D-2 No. 4: Mechanical Degradation Revision 1 
 Thermal-mechanical analyses of the emplacement drifts were conducted using a 
drift-scale model truncated at a distance of 50 m above and below the emplacement drift 
axis. The base of the model was held at zero vertical displacement, whereas the model 
top was held at constant normal traction equivalent to the preemplacement in situ stress. 
Such a model is inappropriate because it allows excessive free upward thermal 
expansion, thereby interfering with the development of thermally induced stress that is 
consistent with the geometry of the emplacement area (NRC 2002, p. 2.1.7-10). 
D.1.2 Related Key Technical Issue Agreements 
Agreements RDTME 3.05 (addressed in Appendix A) and RDTME 3.06 (addressed in 
Appendix B) are related to RDTME 3.02, RDTME 3.10, and RDTME 3.13. RDTME 3.05 
addresses rock-mass mechanical properties of lithophysal rock. The rock-mass properties 
estimate, based on the approach described in resolution of RDTME 3.05, is used for resolution of 
RDTME 3.02, RDTME 3.10, and RDTME 3.13. RDTME 3.06 addresses sensitivity and 
uncertainty related to evaluation of emplacement drift stability. The results from the parameter 
study conducted for resolution of RDTME 3.06 assist the resolution of RDTME 3.02 and 
RDTME 3.13. 
D.3 RESPONSE 
The analyses that address resolution of RDTME 3.02, RDTME 3.10, and RDTME 3.13 are 
summarized in this appendix. Detailed information is provided in Evaluation of Emplacement 
Drift Stability for KTI Resolutions (BSC 2004a). Additional detail on rock mass properties 
D.2 RELEVANCE TO REPOSITORY PERFORMANCE 
The preclosure safety analysis is used to demonstrate the safety of the proposed design and 
operations in the geologic repository operations area with regard to the overall preclosure 
performance objectives through a systematic examination of the site, design, potential hazards, 
initiating events and their resulting event sequences, and the potential radiological exposures to 
workers and the public (10 CFR 63.112). This safety analysis demonstrates that the ground 
control system is not required to prevent or mitigate credible rockfall. This demonstration relies 
upon analyses that show that the waste package does not breach when impacted by credible rock 
blocks. The emplacement drifts are an array of horizontal tunnels trending at 72 azimuth. Each 
drift will have a nominal diameter of 5.5 m and will be separated from the adjacent drifts by a 
nominal center-to-center distance of 81 m (BSC 2003a, Section 5.3.1; Williams 2002). The 
emplacement drifts provide the subsurface access and openings for the structures, systems, and 
components used for emplacement and retrieval operations. 
The rock-mass surrounding the emplacement drifts will be subjected to loadings from in situ, 
thermal, and seismic stresses. The performance of emplacement drifts and ground support is 
analyzed based on the technically justified numerical modeling approach with appropriate 
boundary conditions and load combinations. The emplacement drift stability analysis 
demonstrates (1) the stability of the emplacement drifts with or without the installation of ground 
support, and (2) satisfaction of ground support performance, under the critical combinations of in 
situ, thermal, and seismic stresses during the preclosure period. 
D-3 June 2004 No. 4: Mechanical Degradation Revision 1 
derivation and numerical model development and validation are given in the technical basis 
document. The items covered in this appendix include: 
 Analysis for resolution of RDTME 3.02 
 Describe each type of load, including the bounds of each and how these loads are 
estimated 
 Identify potential critical load combinations for emplacement drift stability and 
ground support performance 
 Conduct analyses based on critical load combinations 
 Analysis for resolution of RDTME 3.10 
 Discuss the ground support functions and their implementation in a numerical model 
 Conduct two- and three-dimensional analyses using assumed and true in situ stress 
conditions and compare their results 
 Provide the basis for the acceptability of the analyses 
 Analysis for resolution of RDTME 3.13 
 Describe and justify model dimension and boundary conditions used in numerical 
modeling 
 Conduct two-dimensional analysis for sensitivity of variations in model dimension. 
The summary and conclusions from various analyses responding to these RDTME issues are 
provided below: 
D.3.1 Response to RDTME 3.02 
 The critical load combinations include in situ plus preclosure seismic stress and in situ 
plus preclosure thermal plus seismic stress. In the latter load combination, two 
potentially critical thermal conditions are: (1) when the drift wall temperature reaches 
its peak value, and (2) when the preclosure period (or the considered repository service 
life) ends, which gives the longest duration of heating or thermal expansion (BSC 2004a, 
Section 6.1). 
 Under the load combinations considered, the emplacement drifts are predicted to be 
stable during the preclosure period, with a minimum factor of safety of 2 for the weakest 
estimate lithophysal rock strength category (BSC 2004a, Section 6.1). 
 For the estimated weak rock strength (Category 1; see Section 4.2.2, Table 4-2 of the 
technical basis document for definition) with low deformation modulus, the load 
combination of in situ plus preclosure seismic stress governs the stability of 
June 2004 D-4 No. 4: Mechanical Degradation Revision 1 
emplacement drifts. Changes in stresses during the seismic ground motions overshadow 
those induced by heating over the entire preclosure period. This is because the increase 
in temperature is relatively moderate and the rock mass modulus is relatively low, 
resulting in a low thermally induced stress (BSC 2004a, Section 6.1). 
 Stresses in the rock adjacent to emplacement drifts are sensitive to the elevated 
temperatures, especially for the higher strength category rock (e.g., Category 5, which 
has a high modulus value). The ratio of modulus to strength controls stability sensitivity 
to the thermal load. If the ratio is large, the effect of heating on the drift stability is 
greater. For lithophysal rock, the ratio increases with rock strength category number. 
The worst conditions of thermal loading are for the best quality rock mass (Category 5; 
see Section 4.2.2 of the technical basis document for definition). The maximum stresses 
are generally anticipated to be associated with the peak temperature following waste 
emplacement. For emplacement drifts excavated in good quality rock, a load 
combination that includes thermal effect may be the most important from a stability 
standpoint (BSC 2004a, Section 6.1). 
 The duration of heating has an effect on the performance of emplacement drifts and 
ground support. This is because an increasing amount of heat will be transferred into the 
rock mass over time, even though the temperature on the drift wall is decreasing. The 
more heat accumulated within the rock, the more thermal expansion the rock will have, 
resulting in additional rock deformations and stresses (BSC 2004a, Section 6.1). 
 For rock bolt performance, the in situ and preclosure seismic load combination is the 
governing loading state. The peak axial forces in rock bolts installed near the springline 
and the crown are induced under combined in situ and preclosure seismic loading 
conditions. During the early stage of heating, the axial forces in bolts actually decrease 
due to the difference in coefficient of thermal expansion between rock bolt steel and 
rock mass and may become compressive if the axial loads are relatively low prior to 
heating. Depending on the length of heating, however, the axial forces in rock bolts are 
predicted to increase with time. This suggests that a thermal condition at the end of the 
preclosure period should also be combined with the preclosure seismic loading condition 
as a potentially governing load combination (BSC 2004a, Section 6.1). 
D.3.2 Response to RDTME 3.10 
Use of two-dimensional analyses for evaluation of emplacement drift stability and ground 
support performance is a generally-accepted practice. In these two-dimensional analyses, the 
bounding scenarios are considered in terms of rock properties and load conditions. Results based 
on the bounding scenarios are usually more conservative than those from three-dimensional 
analyses using more realistic input data of in situ stress field and fractures (BSC 2004a, 
Section 6.5). 
To assess the effect of fractures in rock mass on instability due to wedge-type failures, a 
three-dimensional discontinuum model is required (BSC 2004b). The predicted rockfall 
information can then be used to evaluate the performance of ground support based on uncoupled 
calculations (see Appendix F of this document). The size, mass, and location of blocks can be 
June 2004 D-5 No. 4: Mechanical Degradation Revision 1 
used to perform simple static structural calculations to ensure that the ground support is capable 
of maintaining stability. 
Results from two-dimensional stability analyses from FLAC models are compared with those 
from the three-dimensional FLAC3D models. It appears that the two-dimensional models yield 
more conservative results and, therefore, are justified for use in the ground support design (BSC 
2004a, Section 6.5). 
D.3.3 Response to RDTME 3.13 
The thermal-mechanical impacts of three models with vertical dimensions of 50, 100, and 200 m 
and corresponding thermal-mechanical boundary conditions are considered. The analyses are 
based on the rock mass properties for the lithophysal unit and base thermal loading and forced 
ventilation conditions (BSC 2004a, Section 6.7). 
Use of a model with a vertical dimension of 100 m is adequate from the standpoint of ground 
support design. Additional increase in the model dimension has little effect on the predicted drift 
stability and ground support performance and is not necessary (BSC 2004a, Section 6.7). 
Reduction of the model dimension to 25 m measured from the drift center to the top boundary 
may result in an overestimate of drift closures and stresses near the drift opening (BSC 2004a, 
Section 6.7). 
The information in this report is responsive to agreements RDTME 3.02, RDTME 3.10, and 
RDTME 3.13 made between the DOE and NRC. This report contains the information that the 
DOE considers necessary for NRC review for closure of these agreements. 
D.4 BASIS FOR THE RESPONSE 
This section describes the technical basis and approach for resolving RDTME 3.02 
(Section D.4.2), RDTME 3.10 (Section D.4.3), and RDTME 3.13 (Section D.4.4). 
D.4.1 Overview of Resolution Strategy 
A general strategy for resolving the geomechanical issues related to the RDTME KTIs is 
outlined in Resolution Strategy for Geomechanically-Related Repository Design and 
Thermal-Mechanical Effects (RDTME) (Board 2003). This technical basis document provides a 
summary of the testing and analyses that have been conducted based on this strategy. Although 
this technical basis document is primarily concerned with postclosure performance issues, the 
rock mass properties derived from these studies and the modeling tools developed are also 
applied to preclosure issues such as emplacement drift stability and ground support design. A 
number of documents describing preclosure applications have been developed, 
including: Scoping Analysis on Sensitivity and Uncertainty of Emplacement Drift Stability (BSC 
2003b) and Evaluation of Emplacement Drift Stability for KTI Resolutions (BSC 2004a). These 
documents were developed specifically to resolve RDTME KTI agreements related to preclosure 
issues. These documents present detailed information regarding the analyses of emplacement 
drift stability and ground support performance, and support the resolution of these KTI 
agreements. 
June 2004 D-6 No. 4: Mechanical Degradation Revision 1 
D.4.2 Basis for Resolution of RDTME 3.02 
The focus of this agreement is the technical basis for selecting the critical1 load combinations of 
in situ, thermal, and seismic stresses to support an evaluation of emplacement drift stability and 
ground support performance during the repository preclosure period. 
D.4.2.2 Selection of Load Combinations 
The following load combinations have been considered to support a design analysis of 
emplacement drift stability and ground support performance during the repository preclosure 
period (BSC 2003c): 
 In situ 
 In situ plus seismic 
 In situ plus thermal (as a function of time) 
 In situ plus thermal plus seismic. 
Of these combinations, the last three are considered more important in terms of governing the 
design. 
1 The term, critical means that these load combinations have the greatest potential impact on drift stability. Critical 
is not meant to imply that the load combinations result in impending instability. 
D.4.2.1 Consideration of Sources of Loads 
In a design analysis of emplacement drift stability and ground support performance, stresses 
resulting from three sources are considered: in situ (including excavation effects), thermal 
(radioactive waste heat decay), and seismic stresses. In situ stresses are present before drift 
excavation and will be altered in the vicinity of openings due to drift excavation. Thermal 
stresses occur following the initiation of waste emplacement and are transient in nature. The 
magnitude of thermally induced stresses at a particular location in the vicinity of the 
emplacement drifts is dependent upon the position relative to the emplaced waste packages and 
the airflow rate used in ventilation during the preclosure period. Seismic stresses are dynamic 
and transitory in nature. The magnitude of seismically induced stresses 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 drift openings. 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. 
Detailed discussion on how each of these three loads is estimated and applied in the design 
analysis of emplacement drift stability and ground support performance during the preclosure 
period is presented in Evaluation of Emplacement Drift Stability for KTI Resolutions (BSC 
2004a, Section 6.1). The in situ, thermal, and seismic loads considered in the analysis of drift 
degradation are discussed in Section 5. 
D-7 June 2004 No. 4: Mechanical Degradation Revision 1 
For the rock mass adjacent to the emplacement drift and ground support components that are in 
compression due to excavation-induced rock displacements, the critical load combinations will 
be (1) in situ plus thermal and (2) in situ plus thermal plus seismic. 
In Situ and Thermal Loading CombinationThere are two significant times during the 
preclosure period that are critical from the ground-support design perspective. The first time is 
the point at which the drift wall reaches its peak temperature (Figure D-1). Including this point 
in the analysis of emplacement drift stability is appropriate because the increase in compressive 
stresses in the rock mass immediately adjacent to the drift is proportional to the increase in the 
rock temperature. 
The second time is the end of the preclosure period. The emplacement drifts will be loaded with 
waste packages sequentially, together with ventilation provided. Once the loading is completed, 
the drift wall temperature will go down, as long as a ventilation airflow rate is maintained. As 
indicated in Figure D-1, the drift wall temperature peaks at about 2 years after emplacement for a 
forced ventilation airflow rate of 15 m3/s. The drift wall temperature then gradually decreases 
with time due to the effect of continuous ventilation. However, the rock mass temperature away 
from the drift will continue to rise after the drift wall temperature peaks because the ventilation 
cannot remove 100% of the heat generated by the waste package (see rock temperatures at about 
19 m from the skin of drift opening, shown in Figure D-1). As the temperature of the rock mass 
continues to rise, it causes additional expansion of the rock mass, which results in thermal stress 
increase that must be accounted for in estimation of drift stability and ground support 
performance. As more rock mass is heated, the cumulative effect of continuous rock thermal 
expansion plays a key role in governing the applied preclosure applied stresses. 
June 2004 D-8 No. 4: Mechanical Degradation Revision 1 
Source: DTN: MO0306MWDALAFV.000, ANSYS-LA-Fine.xls and la600c24.rth. 
NOTE: These are temperatures predicted at 600 m from the drift air inlet. 
Figure D-1. Rock Temperatures as a Function of Time at the Drift Wall and Various Positions within the 
Rock Mass 
Thus, there are two time frames (i.e., from the beginning of emplacement to the time when the 
drift wall temperature peaks, and then the end of the preclosure period) to be considered in 
evaluating the effects of thermal load and its combination with other loads. 
The in situ plus thermal load combination may be considered less critical, or less conservative, 
when a potential failure mode of rock mass or ground support components is due to tension. For 
example, rock bolts are generally in tension under excavation-induced rock deformation. The 
coefficient of thermal expansion for steel of rock bolts is usually similar to or higher than that of 
the rock mass. The elevated rock temperatures will result in induced compression in the bolts, 
which, as a result, will offset their preexisting tensile forces. Thus, thermal load is less important 
to rock bolt performance. 
In Situ, Thermal and Seismic Loading CombinationBecause the seismically-induced ground 
motions will induce both tensile and compressive stresses in the rock mass and in rock bolts, 
their effects should be combined with those of all other loads. The seismic load is dynamic in 
nature, while the in situ and thermal loads are essentially static in comparison. The required 
safety margin for the former (dynamic load) is usually lower than that for the latter (static load). 
For this reason, a load combination without the seismic load, such as in situ plus thermal, is also 
considered as critical due to a higher factor of safety required. 
June 2004 D-9 No. 4: Mechanical Degradation Revision 1 
D.4.2.3 Effect and Sensitivity of Various Load Combinations 
Two-Dimensional Continuum Modeling ApproachFrom a ground support design 
perspective, stability of emplacement drifts is judged by overall rock mass displacements and 
stresses. A two-dimensional plane-strain thermal-mechanical parameter analysis is used to 
assess the stability of emplacement drifts. The two-dimensional finite-difference code FLAC 
(Itasca Consulting Group 2002) is used for the analysis. A combination of in situ, thermal, and 
seismic loadings is included in the analysis. The continuum-based analyses assume the rock 
mass conforms to a Mohr-Coulomb failure criteria, with property variations for lithophysal rock 
described in Section 4.2 of this technical basis document. 
Range in Variation of Rock Mass Properties for Lithophysal RockIn the FLAC models, 
rock mass properties for the lithophysal rock properties reflect the effects of lithophysae and 
fractures on rock mass properties; their values are presented in Evaluation of Emplacement Drift 
Stability for KTI Resolutions (BSC 2004a, Table 4-5a, as well as Section 4.2, Table 4-2, and 
Appendix A of this technical basis document). For the bounding case scenarios, the rock mass 
mechanical properties corresponding to strength Category 1 (the weakest rock), and strength 
Category 5 (the strongest rock), are specifically considered. Under the in situ or seismic loading 
conditions, the stability of emplacement drifts is governed by the mechanical properties 
associated with the weakest rock, so use of the Category 1 rock mass properties represents the 
worst case scenario. However, under the thermal loading condition, the drift stability may 
depend on the modulus of the rock mass, since a relatively stiff rock (i.e., a good quality rock) 
may result in higher thermal stresses under elevated temperatures (because the thermal stress is 
proportional to modulus). Use of the Category 5 rock mass properties that give a relatively high 
modulus value are considered as a bounding case scenario in the analysis. 
In Situ Stress State Variation and Rock Mass Yield CriterionThe measured vertical stress 
component at the Yucca Mountain site is maximum and equal to the gravitational component. 
The ratio of the minimum and maximum horizontal stress to vertical is estimated at 0.36 and 
0.62 (see Table 2-2 of the technical basis document). Since a two-dimensional analysis is 
conducted here, bounding horizontal-to-vertical stress ratios (K0), 0.3 and 1.0, are used in the 
analyses to encompass the potential range of in situ horizontal components in the plane normal to 
the emplacement drift axis (BSC 2004a, Table 6.1-1). Only results for a K0 value of 0.3 are 
reviewed here as it provides the most critical combination of in situ stresses. The mechanical 
failure response of rock mass is assumed to conform to Mohr-Coulomb yield criterion. The use 
of the Mohr-Coulomb yield criteria is a typical assumption for rock (Hoek 2000) and was 
verified for lithophysal tuff in Section 4.2.5 of the technical basis document. 
Model Configuration and Boundary ConditionsFigure D-2 illustrates the configuration of a 
typical FLAC model. The vertical dimension of the model is 100 m, and the horizontal is 81 m, 
equal to the drift spacing. The boundary conditions associated with the analyses for in situ, 
thermal, and seismic loading conditions are shown in Figure D-3. Evaluation of Emplacement 
Drift Stability for KTI Resolutions (BSC 2004a, Section 6.1.3) contains details on how each type 
of load is applied in the analyses. 
June 2004 D-10 No. 4: Mechanical Degradation Source: BSC 2004a, Figure 6.1-5. 
Figure D-2. Configuration of a Typical FLAC Model: (a) Mesh; (b) Rock Bolts (in Meters) 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 D-11 Revision 1 
Figure D-3. Geometry and Boundary Conditions for a Typical Two-Dimensional Model 
Additionally, to evaluate the sensitivity of rock bolt performance under various load 
combinations, rock bolts (such as the Swellex bolts proposed for the emplacement drifts (BSC 
2003c, Section 6.3.2) are also included in the FLAC models. These bolts are simulated by rock 
bolt elements available as a standard feature within the FLAC program. As shown in 
Figure D-2, there are a total of ten 3-m-long bolts in each modeled cross section. The drift axial 
spacing of each row of rock bolts is 1.25 m. Two-dimensional modeling of the bolts with regular 
spacing in the drift axial direction involves averaging the three-dimensional effect over the 
distance between the adjacent rows of bolts. Linear scaling of rock bolt material properties is 
required in the FLAC models, as described by Itasca Consulting Group (2002), to account for the 
impact of spacing in the axial direction. This scaling is achieved by dividing the actual property 
by the bolt spacing along the drift. The material properties that need to be scaled include the 
modulus of elasticity, the tensile strength, and the normal and shear bonding stiffness, Kbond and 
Sbond of the bolt-borehole contact. Rock bolt properties are derived from laboratory or field pull 
June 2004 D-12 No. 4: Mechanical Degradation Revision 1 
testing. Data used in these analyses are summarized in Evaluation of Emplacement Drift 
Stability for KTI Resolutions (BSC 2004a, Section 6.1.4). Axial force outputs from the models 
are then multiplied by the bolt spacing to obtain the actual loads. 
In considering the combination of in situ, thermal, and seismic loading conditions, two subcases 
are analyzed. One subcase imposes the seismic motions at 2 years after waste emplacement, 
which corresponds to a state that the temperature reaches its peak value on the drift wall; and the 
other subcase applies the seismic load at 50 years, which is the end of the preclosure ventilation 
period. 
Results of AnalysesResults of the analyses for the three different load combinations and two 
bounding rock mass categories (i.e., Categories 1 and 5 only) are shown in Figures D-4 to D-23. 
The observations based on a comparison of these results can be summarized as follows: 
 The analyses define the most important loading combinations in regard to tunnel 
stability and rock bolt loading. In all cases analyzed, tunnels are stable with 
approximate strength-to-stress ratios of 2 to 4 for the lowest and highest lithophysal 
strength categories, respectively. Even though modeling of rock bolts was inherent in 
these analyses as shown in Appendix B, the tunnels are stable in an unsupported fashion. 
Therefore, the role of ground support is not to ensure opening stability under preclosure 
in situ, thermal and seismic loadings, but to retain any small, loosened rock particles. 
 Drift closure, defined as the relative displacement between the invert and the crown for 
the vertical closure and between the walls at the springline for the horizontal closure (see 
Figure D-2b), are not significantly affected by either preclosure heating or preclosure 
seismic ground motions, as indicated in Figures D-4 and D-5. Heating after waste 
emplacement induces additional drift closure, especially in the horizontal direction, but 
the increases are on the order of millimeters. Changes in the drift closures during the 
seismic ground motion are generally independent of whether the drift is heated or not. 
Compared to heating from waste packages, seismic ground motions have greater impact 
on the drift closures, but the impact is inconsequential in comparison to initial closure of 
the drift from excavation alone and do not affect drift stability (see Figures D-4a and 
D-4b). 
 Under the load combinations and bounding lithophysal rock mass properties considered, 
the emplacement drifts are predicted to be stable. The stress paths2 near the springline 
and the crown, as shown in Figure D-6, remain within the Mohr-Coulomb yield criteria, 
meaning that the rock mass adjacent to the emplacement drifts behaves elastically under 
all load combinations. A stress path located below the yield criteria indicates that the 
transient stress variation is such that the rock mass remains within the elastic range. 
 For the weakest lithophysal rock case (Category 1), the load combination of in situ plus 
seismic is considered to be critical for the stability of emplacement drifts. Changes in 
2 The term, stress path refers to a history of the principal stress at a point in the rock mass as it undergoes the 
transient loading and unloading associated with heating and cooling or seismic stressing. Plotting of the transient 
stress path on a standard plot of principal stresses with superimposed Mohr-Coulomb yield criteria (e.g., Figure B-2) 
allows easy identification of the location, timing, and extent of yield of the rock mass. 
June 2004 D-13 No. 4: Mechanical Degradation Revision 1 
stresses during the seismic ground motions are greater than those induced by heating 
over the entire preclosure period (see Figure D-7). This is because the increase in rock 
mass temperature is relatively moderate and the rock mass modulus value is also low, 
resulting in a low thermally induced stress. 
 Stresses in the rock adjacent to emplacement drifts increase with elevated temperatures, 
especially for the stiffer rock (see Figure D-8). The maximum stresses are generally 
anticipated to be associated with the peak temperature occurring at 2 years following 
waste emplacement. For emplacement drifts excavated in good quality rock, a load 
combination that includes thermal effect may potentially be critical. 
 Overall factors of safety or strength-to-stress ratio for emplacement drifts are not 
significantly affected by variations in loading conditions, as shown in Figures D-9 and 
D-10. They are controlled by the quality or strength of the rock mass considered. For 
the weakest lithophysal rock (Category 1), the average factor of safety in a region of 
3-m-thick annulus around the drift is about 2 for the governing load combination, while 
for the strong rock (Category 5), it is greater than 4. 
 For rock bolt performance, the in situ and seismic load combination is governing. As 
shown in Figures D-11 and D-12, the peak axial forces in rock bolts installed near the 
springline and the crown are induced under combined in situ and seismic loading 
conditions. Under thermal loading conditions, however, the axial forces in bolts initially 
decrease, and then gradually increase with time. The axial forces may become 
compressive if they are relatively low prior to heating (see Figure D-12). By comparing 
Figure D-11 with Figure D-12, it can also be seen that the axial forces in bolts are very 
sensitive to the stiffness of rock: the lower the stiffness is, the higher the tensile axial 
forces. It should be noted that the in situ drift scale heating test, which was located in 
nonlithophysal rock (rock mass modulus at least as high as Category 5 lithophysal rock) 
was heated to temperatures of nearly 200oC, followed by a cool-down phase. The rock 
bolts utilized for ground support in this drift have showed no signs of loss of function. 
As discussed in Section 4.2.7 of this technical basis document, the thermally-induced 
roof tangential compressive stresses were actually high enough to equal the rock mass 
compressive strength at the maximum temperature to induce minor spalling. Even under 
this extreme condition, the rock bolts continued functioning and maintained suspension 
of wire mesh in the roof. 
 The duration of heating has some impact on the performance of emplacement drifts and 
ground support. This is because an increasing amount of heat will be transferred into the 
rock mass over time, even though the temperature on the drift wall is decreasing. The 
more heat accumulated within the rock, the more thermal expansion of the rock, 
resulting in additional rock deformation and stressing. Axial forces in rock bolts are also 
predicted to increase with time due to continuous heating. This indicates that a thermal 
condition at the end of the preclosure period should be combined with seismic loading 
condition as a potentially critical load combination. 
June 2004 D-14 No. 4: Mechanical Degradation Revision 1 
 Inclusion of the seismic condition is considered critical, and the load combination of in 
situ and thermal alone does not control the emplacement drift stability and ground 
support performance. 
The observations made above are based on the analyses using rock mass properties for the 
lithophysal units. Similar observations can be made if rock mass properties for the 
nonlithophysal units are used (BSC 2003c, Sections 6.1.2 and 6.1.3). 
June 2004 D-15 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 6.1-6. 
NOTE: The displacement values shown in (b) are initialized prior to seismic shaking. 
Figure D-4. Drift Closure for Various Load Combinations (Category 1 and K0 = 0.3): (a) In Situ plus 
Thermal; (b) In Situ plus Seismic and In Situ plus Thermal plus Seismic 
June 2004 D-16 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 6.1-16. 
NOTE: The displacement values shown in (b) are initialized prior to seismic shaking. 
Figure D-5. Drift Closures for Various Load Combinations (Category 5 and K0 = 0.3): (a) In Situ plus 
Thermal; (b) In Situ plus Seismic and In Situ plus Thermal plus Seismic 
June 2004 D-17 No. 4: Mechanical Degradation Source: BSC 2004a, Figures 6.1-7 and 6.1-17. 
Figure D-6. Stress Paths near Springline and Crown for Various Load Combinations (a) Lithophysal 
Rock Strength Category 1 (Lowest Strength and Modulus) and (b) Lithophysal Rock 
Strength Category 5 (Highest Strength and Modulus) 
D-18 No. 4: Mechanical Degradation 
Revision 1 
June 2004 Revision 1 
Source: BSC 2004a, Figure 6.1-8. 
0 = 0.3): (a) In Situ plus Thermal; (b) In Situ plus Seismic and In Situ plus Thermal 
Figure D-7. Major Principal Stresses near Drift Opening for Various Load Combinations (Category 1 
and K 
plus Seismic 
June 2004 D-19 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 6.1-18. 
0 = 0.3): (a) In Situ plus Thermal; (b) In Situ plus Seismic and In Situ plus Thermal 
Figure D-8. Major Principal Stresses near Drift Opening for Various Load Combinations (Category 5 
and K 
plus Seismic 
June 2004 D-20 No. 4: Mechanical Degradation Source: BSC 2004a, Figure 6.1-9. 
Figure D-9. Contours of Strength-to-Stress Ratios around Emplacement Drifts under Various Load 
Combinations (Category 1 and K0 = 0.3) 
No. 4: Mechanical Degradation D-21 
Revision 1 
June 2004 Source: BSC 2004a, Figure 6.1-19. 
Figure D-10. Contours of Strength-to-Stress Ratios around Emplacement Drifts under Various Load 
Combinations (Category 5 and K0 = 0.3) 
No. 4: Mechanical Degradation D-22 
Revision 1 
June 2004 Source: BSC 2004a, Figure 6.1-10. 
Figure D-11. Axial Forces in Rock Bolts Installed near Springline and Crown for Various Load Combinations 
(Category 1 and K0 = 0.3): (a) In Situ plus Thermal; (b) In Situ plus Seismic and In Situ plus Thermal 
plus Seismic 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 D-23 Revision 1 
Source: BSC 2004a, Figure 6.1-20. 
0 = 0.3): (a) In Situ plus Thermal; (b) In Situ plus Seismic 
Figure D-12. Axial Forces in Rock Bolts Installed near Springline and Crown for Various Load 
Combinations (Category 5 and K 
and In Situ plus Thermal plus Seismic 
June 2004 D-24 No. 4: Mechanical Degradation Revision 1 
D.4.2.4 Uncertainty Associated with Analyses of Load Combination Effect 
The following factors may contribute to uncertainties associated with the analyses of load 
combination effects on the emplacement drift stability and ground support performance: 
 Variation of modeling approaches. The analyses presented in this appendix for the 
RDTME 3.02 agreement are based on a two-dimensional continuum approach. This 
approach is considered appropriate from the standpoint of ground support design. In the 
study of drift degradation, a three-dimensional analysis based on a discontinuum 
approach is used to assess the effect of fractures on rockfall (Sections 5.3.1 and 5.3.2 of 
this technical basis document) (BSC 2004b). The predictions of potential rockfalls 
under various load combinations during the preclosure period are provided in 
Appendix F of this document. The conclusion from these studies is that unsupported 
drifts in fractured rock subjected to preclosure thermal and seismic loading are stable. 
 Variation in thermal properties and ventilation shutdown. Appendix B (Section B.4.2) 
presents results from BSC 2003b, Sections 6.4.2 and 6.6.2, in which the impact of 
variability of thermal conductivity, coefficient of thermal expansion, and specific heat 
on drift stability were examined. The impact of variation of these parameters as well as 
credible periods of ventilation shutdown on stability of unsupported drifts in the 
preclosure was found to be negligible. 
D.4.2.5 Summary of Analyses for Resolution of RDTME 3.02 
The critical preclosure load combinations include in situ plus seismic and in situ plus thermal 
plus seismic. In the latter load combination, two potentially critical thermal conditions 
are: (1) when the drift wall temperature reaches its peak value, and (2) when the preclosure 
period ends. These load combinations, as well as bounding ranges in rock mass properties and in 
situ stress conditions, were considered in the evaluation of emplacement drift stability and 
ground support performance. The conclusion from these analyses, coupled with similar analyses 
for unsupported drifts presented in Appendix B, show that the emplacement drifts are stable with 
a minimum strength to stress ratio of approximately 2 for the lowest strength lithophysal rock. 
D.4.3 Basis for Resolution of RDTME 3.10 
The emphasis of this agreement concerns the technical basis for using two-dimensional 
continuum models for evaluation of emplacement drift stability and ground support performance 
during the repository preclosure period. 
D.4.3.1 Methods of Analyses for Ground Support Design 
Both empirical and numerical methods are widely used in mining and tunneling industries for 
design of ground support. The empirical methods are primarily tools for assessing the needs for 
initial ground support. They may also be used to develop preliminary estimates of the final 
ground support system(s) to be used. Design issues such as personnel safety, constructibility, 
and geologic mapping requirements may be factored into the design of the ground support 
system at this stage. Then, with the aid of computer modeling, the stability of underground 
openings may be further assessed and the recommended ground support system analyzed. 
D-25 June 2004 No. 4: Mechanical Degradation Revision 1 
In the case of a waste repository, evaluation of the stability of emplacement drifts and the 
performance of ground support relies on numerical methods, due primarily to the complex nature 
of the combined in situ, thermal, and seismic loads. 
Emplacement drifts will be excavated primarily in two different types of densely welded tuff: 
lithophysal and nonlithophysal (BSC 2003a, Table II-2). The drift response to anticipated loads 
will depend on the characteristics of rock where the drifts are located. The approaches used in 
design analyses reflect this difference. 
Consideration of Geologic Features in Choice of Modeling Approach D.4.3.1.1 
D.4.3.1.1.1 Analysis for Drifts in Lithophysal Rocks 
The application of numerical methods to assessing the stability of repository drifts in the 
lithophysal rock depends on the rock mass structure, material properties, and failure 
mechanisms. As discussed in the Section 2.3 of this technical basis document, the characteristics 
of this type of rock is primarily controlled by the presence of lithophysae and the short fractures 
interconnecting lithophysae, particularly in the lower lithophysal unit (Board 2003) (Section 3.4 
of this technical basis document). Joint sets with a spacing of greater than 2 m also exist, but are 
not judged to be a key factor that governs the deformation and failure mechanism (BSC 2004b). 
Since the lithophysae roughly uniformly distributed through the lithophysal units, it is reasonable 
to assume that a mechanical constitutive model, dependent on porosity and matrix strength 
properties, can capture the general failure mechanisms. In addition, the lithophysal cavity radius 
is much smaller than the radius of emplacement drifts. Individual lithophysal cavities will, in 
general, have negligible impact on the predicted overall stability of the drifts, as long as the 
effect of lithophysae is accounted for in the properties used in the analysis (see Section 4.2 of 
this technical basis document for a discussion about the modeling approach for lithophysal 
rocks). Therefore, it is appropriate to use a two-dimensional continuum approach based on 
equivalent rock mass properties derived from a constitutive model (Mohr-Coulomb) representing 
the lithophysal rocks for analysis of stress distribution, deformation and yield estimation. 
Quantitative analysis of rockfall in emplacement drifts needed for postclosure performance 
assessment purposes, requires a discontinuum approach to modeling since the continuum 
approach is not capable of explicit modeling of fracture or detachment of blocks from the rock 
mass. 
In the equivalent continuum approach, the rock structural features such as lithophysae and 
short-length fractures are not modeled explicitly. Rather, the effects of these features are 
reflected in the material properties and failure mechanisms defined. The general stress 
distribution, deformation, and yield predicted by this approach is approximately equivalent to 
models that explicitly represent the fractures as the overall response of the rock mass is assumed 
to conform to a Mohr-Coulomb yield condition. Section 4.2 of this technical basis document 
describes the process of development of rock mass properties and verification of the mechanical 
constitutive model for lithophysal rock in detail. 
D-26 June 2004 No. 4: Mechanical Degradation Revision 1 
D.4.3.1.1.2 Analysis for Drifts in Nonlithophysal Rocks 
The strength and modulus of intact blocks of the nonlithophysal rock mass is large in comparison 
to the applied preclosure stress levels. Primary structural features of the nonlithophysal rocks are 
four fracture sets with a mean spacing on the order of 1 m or more (see Section 2.3 of this 
technical basis document). The mechanical response of emplacement drifts in the nonlithophysal 
rocks is generally assumed to be anisotropic and three-dimensional in nature as the 
fracture set orientations are not perpendicular and parallel the drift axis. With the discontinuum 
approach, interaction between joints and rock blocks that are bounded by joints can be simulated 
explicitly and in detail. The characteristics of joints, such as joint geometry and frequency, are 
the direct inputs to numerical models. Such detail is necessary when the model is used to 
explicitly calculate rockfall, as is the case in postclosure performance analysis. 
However, when the goal of the analysis is to estimate overall stability of the excavations, and, in 
particular, to examine stress distributions, deformations and yield potential for ground support 
estimation, a discontinuum modeling approach is not required. A sophisticated threedimensional 
discontinuum model is rarely used in the mining or tunneling industry for ground 
support design purposes. Typical industrial applications of three-dimensional discontinuum 
modeling include problems such as examination of the impacts of faults or block detachment in 
large excavations in blocky ground. Common practice (Hoek 2002) in excavation stability 
assessment and ground support design is the use of empirical methods based on geotechnical 
classification for initial support specification. The empirical methods are often supplemented by 
the use of continuum-based numerical modeling using estimated equivalent rock mass properties 
for detailed analysis of excavation stability. A two-dimensional model based on the equivalent 
continuum approach is conventionally accepted for simulating the response of jointed rock mass. 
The success of the two-dimensional continuum approach is due primarily to the conservatism 
built in the two-dimensional analysis and the high factor of safety used in most of the 
conventional tunnel and mining design. 
D.4.3.1.2 Consideration of Loading Conditions 
In situ stress conditions at the repository host horizon are not isotropic in the horizontal plane. 
The vertical stress, resulting from overburden gravitational weight, is the major principal stress 
component. The major horizontal principal stress is estimated to be 62% of the vertical stress 
acting in the N15E, whereas the minor horizontal principal stress is estimated to be 36% of the 
vertical stress acting in the N75W. This was observed from field in situ stress measurements 
(DTN: MO0007RIB00077.000). In a two-dimensional analysis, this anisotropic stress condition 
cannot be simulated. In addition, the emplacement drifts are oriented at an azimuth of 72 
(N72E) (BSC 2003a, Section 5.1.4). This indicates that the drifts are neither perpendicular to 
nor parallel with the horizontal principal stress components. In a two dimensional, crosssectional 
tunnel design analysis, vertical and horizontal stresses are applied to the model 
horizontal and vertical boundaries. Thus, the drift longitudinal axis is assumed to orient 
perpendicular to, or parallel with, the principal stresses. The effect of this model simplification 
on the ground support design depends on how the in situ stress condition is considered in the 
two-dimensional analysis. 
D-27 June 2004 No. 4: Mechanical Degradation Revision 1 
In Ground Control for Emplacement Drifts for LA (BSC 2003c), the bounding scenarios of in 
situ stress conditions were considered for the design calculation of ground support. The 
horizontal stress was assumed to be isotropic, but horizontal-to-vertical stress ratio (K0) of 0.3 
and 1.0 were used. These were considered the lower and upper bounds of anticipated stress 
conditions (Sun 2002, Table 3-2). Results from these bounding conditions capture the worst case 
scenario and, therefore, are considered more conservative than those based on the anisotropic 
condition. 
D.4.3.1.3 Consideration of Drift Behavior 
The purpose of the analyses discussed in Evaluation of Emplacement Drift Stability for KTI 
Resolutions (BSC 2004a, Section 6.3), is to examine drift stability to preclosure loading 
conditions and rock mass property variability. As discussed above, a two-dimensional 
continuum modeling approach is considered adequate for this purpose. Rockfall induced by 
preclosure seismic ground motions is examined using three-dimensional discontinuum models 
(Appendix F). However, for the purposes of emplacement drift ground support design, 
two-dimensional models are used to predict the drift deformation, stability, and approximate drift 
strength to stress ratios for bounding ranges of conditions for unsupported drifts. The ground 
support function and requirements are determined from the deformations and depth of yield from 
these models in a conventional fashion without explicitly modeling the ground support. As 
described below, additional modeling is performed in which the rock bolts are represented 
explicitly within the model to verify their functionality. 
In some repository nonemplacement areas, such as intersections, three-dimensional modeling is 
used to examine stability since the geometry of these areas dictates a three-dimensional stress 
condition. For this reason, the design analyses for the intersections and other openings where 
three-dimensional behavior dominates are based on three-dimensional models (BSC 2004c, 
Section 6.5.1.2). 
D.4.3.2 Effect of Ground Support Components in Numerical Model 
In general, the rock mass elastic deformation as a result of tunnel excavation is largely complete 
before rock support is installed because the support is typically installed some distance behind 
the advancing tunnel face. Thus, the rock support will actually experience only small strains 
from deformations related to tunnel excavation if the rock mass is not actively yielding. Since 
elastic conditions are predicted for emplacement drift excavation, the primary loading of tunnel 
support is the result of preclosure thermal and seismic loading. 
Ground support components installed in emplacement drifts, such as proposed rock bolts and 
perforated steel sheets, also have a very limited effect on the strains and stresses in rock mass 
predicted in numerical models. The effect is limited because, as a reinforcement member, rock 
bolts have much lower stiffness compared to that of the rock mass. Therefore, from a practical 
standpoint, since the emplacement drifts do not undergo yield and significant closure strain, 
ground support components have negligible impact on the modeled results. Results from the 
analyses of unsupported emplacement drifts can be used for estimation of drift stability and 
deformation, and thus provide an indication of the stability state and, therefore, the ground 
support requirements. 
D-28 June 2004 No. 4: Mechanical Degradation Revision 1 
D.4.3.3 Adequacy of Two-Dimensional Modeling 
To further demonstrate the adequacy of two-dimensional modeling for evaluation of 
emplacement drift stability and ground support performance, several comparisons between 
two-dimensional and three-dimensional analyses are presented below. 
D.4.3.3.1 Comparison of Drift and Regional Scale Thermal-Mechanical Analyses 
In Drift Degradation Analysis (BSC 2004b, Section 6.2 and Appendix C), coupled 
thermal-mechanical processes in the rock mass surrounding the repository drifts are examined 
using a drift-scale calculation as well as a coupled regional- and drift-scale calculation. 
The drift-scale calculation (both thermal and mechanical) considers an infinite extent 
(perpendicular and in the direction of the drifts) of the repository. Consequently, the problem is 
two-dimensional in nature and only a single drift need be included in the calculation with a 
symmetry boundary condition on a plane midway between the emplacement drifts (BSC 2004b, 
Section 6.2). The thermal part of the drift-scale calculation was performed by the NUFT 
thermal-hydrology software (LLNL 2002) simulating two-dimensional drift-scale 
thermal-hydrologic behavior. The temperature history results from the NUFT software 
calculation were imported to the FLAC code model that calculates the thermal strain and stress 
around an emplacement drift. Algorithms for importing temperatures from NUFT to other 
models is described in Drift Degradation Analysis (BSC 2004b, Appendix U) 
The coupled regional- and drift-scale thermal-mechanical calculation was conducted to support 
this drift-scale calculation by assessing repository-scale effects, including edge effects and the 
effects of finite repository size and depth on predicted temperatures and stresses (BSC 2004b, 
Appendix C) (Section 5.3.1 of this technical basis document). These calculations are threedimensional, 
and the analysis was carried out in two steps. First, the regional-scale thermalmechanical 
calculation was used to determine temperature and stress changes on the scale of the 
entire mountain. Second, the drift-scale thermal-mechanical analysis was performed such that 
boundary conditions for temperature and stress fields (functions of time) were determined from 
the regional-scale calculation. Thus, this calculation did not use any simplifying assumptions 
(e.g., infinite extent of the repository) for the boundary conditions. Both components of the 
regional- and drift-scale thermal-mechanical calculations were performed using FLAC3D (BSC 
2002). 
Stresses in the drift wall and crown for conditions in the middle of the repository for 10 years 
after waste emplacement, as predicted by the drift-scale calculations (FLAC) and coupled 
regional- and drift-scale calculations (FLAC3D), are shown in Figure D-13. 
Agreement of the tangential stresses in the crown is good, as seen in Figure D-13. The vertical 
stress in the wall predicted from the FLAC drift-scale model is slightly higher than that from the 
FLAC3D coupled regional- and drift-scale calculation, but the difference is not significant. 
These results suggest that use of a two-dimensional drift-scale model, as adopted in the ground 
support design analysis, is justified and is more conservative if loading conditions with bounding 
loads are considered. 
June 2004 D-29 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004b, Figure 6-34. 
NOTE: Stresses shown are positive in compression. 
Figure D-13. Comparison between the FLAC Code and FLAC3D Code Predictions of Stresses around 
the Drift after 10 Years of Heating 
D.4.3.3.2 Comparison of Two-Dimensional (FLAC) and Three-Dimensional (FLAC3D) 
Drift-Scale Mechanical Analyses 
FLAC3D Model with Anisotropic Stress ConditionA three-dimensional FLAC3D model 
based on an anisotropic in situ stress condition was developed (BSC 2004a, Section 6.5.3.2). 
The major horizontal principal stress is 62% of the vertical stress acting in the N15E, whereas 
the minor horizontal principal stress is 36% of the vertical stress acting in the N75W. In 
addition, the emplacement drift is oriented at an azimuth of 72 (N72E). The model is 
constructed so that the y-axis is parallel to the drift axis (pointing to N72E), x-axis pointing 
N162E, and z-axis upward. The horizontal components of principal stresses are not along x- or 
y-axes. In addition to the normal stress components, an initial shear component is also defined. 
To achieve the anisotropic stress condition, the model boundaries are fixed in three orthogonal 
(x, y, z) directions. Figure D-14 shows a perspective view of model mesh and a horizontal 
section (plan) view of the initial principal stress field. The orientation of the short bars depicted 
in Figure D-16b indicates the direction of the horizontal principal stresses. It is evident that an 
anisotropic stress field is developed, and this stress field is neither perpendicular nor parallel to 
the drift axis. 
The FLAC3D model is run for the in situ stress condition using the lithophysal rock mass 
properties for rock strength Categories 1, 3, and 5. The predicted drift vertical and horizontal 
June 2004 D-30 No. 4: Mechanical Degradation Revision 1 
0 
closures are compared in Figure D-15 with those from FLAC analyses with K0 values of 0.3, 0.5, 
and 1.0. Drift closures predicted from the FLAC3D analysis are generally consistent with those 
from the FLAC analyses using a K0 value of 0.5. This suggests that the effect of an isotropic 
stress field with a K0 value of 0.5 is similar to the actual stress condition (see Section 2.3.3 of 
this technical basis document). The results also indicate that the scenarios represented by the K 
values of 0.3 and 1.0 can bound the anticipated stress condition. Using the K0 values of 0.3 
and 1.0 overestimates the vertical and horizontal closures by about 5 mm and 25 mm, 
respectively, for the weakest rock case (Category 1). Therefore, use of the two-dimensional 
modeling with the bounding stress conditions is appropriate and justified. 
D.4.3.4 Summary of Analyses for Resolution of RDTME 3.10 
From the overall preclosure drift stability and ground support design perspective, use of 
two-dimensional modeling is conventional and justified as long as the bounding scenarios in 
terms of loading and rock conditions are considered. However, if the problem of interest is the 
effect of fractures on rockfalls, three-dimensional analysis based on discontinuum approach is 
required. 
D-31 June 2004 No. 4: Mechanical Degradation Source: BSC 2004a, Figure 6.5-4. 
Figure D-14. FLAC3D Model Mesh and Initial Principal Stress Field 
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Revision 1 
June 2004 D-32 Source: BSC 2004a, Figure 6.5-5. 
Figure D-15. Comparisons of Drift Closures from FLAC and FLAC3D Predictions for Various Rock Mass 
Categories 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 D-33 Revision 1 
D.4.4 Basis for Resolution of RDTME 3.13 
As discussed in Section D.4.2, thermal-mechanical analyses of the emplacement drifts were 
conducted using a drift-scale model truncated at a distance of 50 m above and below the 
centerline of emplacement drifts (100 m total model vertical dimension). The base of the model 
was fixed in the vertical direction, whereas the top of the model was prescribed a constant 
normal traction equivalent to the in situ stress and given no constraint in displacements (see 
Figure D-3). NRC questioned the appropriateness of such a model in simulating the 
development of thermally induced stress and believed that the model might allow excessive free 
upward thermal expansion during heating (NRC 2002, p. 2.1.7-10). To investigate the sensitivity 
of predicted drift and ground support performance to variations in the model (vertical) dimension 
and boundary conditions, additional thermal-mechanical analyses have been conducted. 
Sections D.4.4.1 to D.4.4.3 describe the results obtained from this endeavor. 
D.4.4.1 Selection of Model Dimensions and Boundary Conditions 
For mechanical analyses of a deep excavation in hard rock (i.e., the diameter of the tunnel is 
much less than the depth to ground surface), a model dimension equal to about 5 to 10 times the 
diameter is generally sufficient to preclude boundary effects on mechanical predictions (Itasca 
Consulting Group 2002, Sections 3.2.1 and 3.3.4.2). This is particularly true when the rock mass 
behaves in an elastic fashion. It is not necessary to fully extend the top boundary of the model to 
the ground surface, as long as a normal traction is applied to the top boundary that is equal to the 
load applied by the remaining overburden. In this case the top boundary is allowed to move in 
the vertical direction as the stress conditions dictate. 
For thermal analyses of the effect of heating that continues more than thousands of years, a large 
model dimension is required or the boundary is set at the location where the boundary condition 
is known or can be easily defined. For example, the lateral boundaries for thermal analyses of 
emplacement drifts are located at the middle of the pillar between two adjacent drifts because 
these boundaries can be conservatively assumed to be adiabatic due to thermal symmetry. The 
top and bottom boundaries cannot be located too close to the emplacement drifts because the 
region of influence by heating in emplacement drifts may eventually extend to the ground 
surface within thousands of years. A ground surface that is about 400 m from the emplacement 
drifts as the top boundary is used in the thermal analyses of the ventilation model for predicting 
the temperatures during the preclosure period (BSC 2004d). To ensure consistency of the 
thermal-mechanical analyses described here with the thermal analyses of the ventilation model, 
transient temperature distributions from the ventilation model are imported to the mechanical 
models used for thermal-mechanical analyses of emplacement drift stability and ground support 
performance for the license application design (BSC 2003c). The procedure for importing 
temperatures from the thermal model to the mechanical model are described in (BSC 2004b, 
Appendix U). 
The model output parameters of interest in thermal-mechanical analyses are the principal stresses 
(and stress path) and displacement in the rock mass adjacent to an emplacement drift during 
heating. Heating in rock mass is expected to result in thermal strain and stress in the regions at a 
distance of several times the drift diameter from the drift opening, such as near the middle of 
pillar or the ground surface. The thermal strain and stress in these regions may not have adverse 
D-34 June 2004 No. 4: Mechanical Degradation Revision 1 
effect on emplacement drift stability and ground support performance. So selection of the model 
dimension should be based on the criterion that the selected dimension is judged adequate if any 
additional increase in the model dimension with associated boundary conditions has no, or 
negligible, impact on the calculated principal stresses and excavation wall displacement. 
To further examine whether the model dimension used in the thermal-mechanical analyses of 
ground support design is appropriate, ,sensitivity study was conducted (BSC 2003b, 
Section 6.2.1). In this sensitivity study, three different total model vertical dimensions, equal to 
50, 100, and 200 m, were used. 
A total vertical dimension of 50 m (upper and lower boundaries of 25 m from the drift 
centerline) is considered as a lower bound in terms of the opening size because the top and the 
bottom boundaries are located at a distance of about five times the drift diameter (5.5 m). 
Further reduction in the vertical dimension may result in some degree of boundary effect on the 
results of interest (Itasca Consulting Group 2002). On the other hand, a vertical dimension of 
200 m is considered as a reasonable upper bound for the thermal-mechanical analyses. 
As shown in Figure D-16, noticeable increases of temperature occur at a distance of 25 m above 
and below the drift center, while those at a distance of 50 m or 100 m from the drift are less than 
1C over the entire preclosure period. Therefore, results from the model using the upper bound 
vertical dimension can be used to examine boundary impacts for the base case of a total model 
dimension of 100 m (50 m up and down from drift centerline) as was assumed for analyses 
presented in Evaluation of Emplacement Drift Stability for KTI Resolutions (BSC 2004a, 
Section 6.7). 
Figure D-17 shows the configurations and meshes of three different models for these additional 
analyses. To allow comparison, the figure also presents a close-up view of the mesh refinement. 
It is noted that the mesh sizes within a region of 10  10 m around the drift opening are actually 
identical in these three models. 
Boundary conditions for the models with different vertical dimensions are identical, as shown in 
Figure D-3, except the magnitude of normal traction applied on the top of a model and the 
prescribed (imported from the ventilation thermal model) transient temperatures on both the top 
and the bottom boundaries. The normal traction is equivalent to the vertical component of in situ 
stress at the elevation of the top boundary of a model, and therefore dependent on the overburden 
thickness above that boundary. For an emplacement drift excavated at a depth of 400 m, the 
overburden thickness is 375, 350, and 300 m for the model vertical dimensions of 50, 100, and 
200 m, respectively. As a result, in situ stresses are the same at any comparable location within 
the models with different vertical dimensions. 
D.4.4.2 Description of Analyses for Examination of the Impact of Model Dimension 
Continuum thermal-mechanical analyses based on the models with different vertical dimensions 
are performed for combined in situ, preclosure thermal, and preclosure seismic loading 
conditions. The seismic ground motions are applied at 50 years after waste emplacement. The 
analyses are conducted using lithophysal rock mass properties for bounding strength 
Categories 1 and 5 and bounding K0 values of 0.3, 0.5, and 1.0. For comparison purposes, only 
June 2004 D-35 No. 4: Mechanical Degradation Revision 1 
the results associated with a K0 value of 0.5 are presented because the conclusions drawn are 
applicable to those for K0 values of 0.3 and 1.0. 
D.4.4.2.1 Effects of Model Dimension on Predicted Temperature Distributions 
In Figure D-18, temperature contours at 10 and 50 years following waste emplacement are 
compared for total model vertical dimensions of 50, 100, and 200 m. The transient temperatures 
from the ventilation model, with 400 m dimension, are imported into all internal elements of 
these truncated models. However, at the upper and lower boundaries, a constant temperature (at 
any given time step) is assumed. Therefore, it is possible that slight temperature effects could 
occur near the boundaries. It can be seen that temperature distributions within comparable 
regions of the models with vertical dimensions of 100 and 200 m are nearly identical, and only a 
small degree of discrepancy occurs near the corners of the smallest model (50-m vertical 
dimension) compared to those in the other two models. This small discrepancy is a result of the 
boundary effect described previously. Overall, temperatures within an annulus of about 3 to 
4 times the drift diameter are not distinguishable among the three models. 
Effects of Model Dimension on Predicted Rock Displacements 
Comparisons of the drift closures induced under combined in situ and thermal loads are shown in 
Figure D-19. There are almost no noticeable differences in both vertical and horizontal drift 
closures for the Category 1 rock mass. Only a small discrepancy exists for Category 5 rock mass 
between the predicted closures from the smallest model (50-m vertical dimension) and those 
from the other two modeled vertical dimensions. It appears that thermally induced rock 
displacements in the model with a small vertical dimension are slightly more sensitive to 
elevated temperatures. Results clearly indicate that use of a vertical dimension of 100 m is 
adequate, and there is no evidence that use of a vertical dimension of 100 m will underestimate 
the effect of thermal loading conditions. 
Drift closures due to preclosure seismic ground motions (1x10-4 annual exceedance) that occur at 
50 years after heating are shown in Figure D-20. As shown, the effect of model size variations is 
D.4.4.2.2 
negligible. 
D.4.4.2.3 Effects of Model Dimension on Predicted Stress Changes 
Time histories of the major principal stresses at locations for models of various dimension are 
shown in Figure D-21. Similar to what is observed for the drift closures, changes in the model 
dimension only have an impact on the smallest model when used in conjunction with Category 5 
rock mass properties. Increase in the model size beyond 100 m has no, or minimal effect on 
stress. These can also be observed from the stress paths, shown in Figure D-22. 
Also included in Figure D-22, are the stress paths at locations about 6 m from the wall measured 
horizontally and 4 m from the crown measured vertically. The differences among these three 
models are negligible. 
The effect of preclosure seismic ground motions (1x10-4 annual exceedance) on stresses is shown 
in Figure D-23. Results indicate that the predicted changes of stresses due to ground motions are 
nearly independent of the model dimension. 
D-36 June 2004 No. 4: Mechanical Degradation Source: BSC 2004a, Figure 6.7-1. 
Figure D-16. Time Histories of Rock Temperatures on Model Boundaries 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 D-37 Source: BSC 2004a, Figure 6.7-2. 
Figure D-17. Comparisons of Configurations and Meshes for Models with Different Vertical Dimensions 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 D-38 Source: BSC 2004a, Figure 6.7-4. 
Figure D-18. Comparisons of Temperature Distributions in Models with Different Vertical Dimensions at 
10 and 50 Years after Waste Emplacement 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 D-39 Revision 1 
Figure D-19. Comparisons of Drift Closures for Models with Different Vertical Dimensions 
Source: BSC 2004a, Figure 6.7-5. 
June 2004 D-40 No. 4: Mechanical Degradation Source: BSC 2004a, Figure 6.7-6. 
Figure D-20. Comparisons of Drift Closures under Seismic Ground Motions for Models with Different 
Vertical Dimensions 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 D-41 Source: BSC 2004a, Figure 6.7-13. 
Figure D-21. Comparisons of Major Principal Stresses for Models with Different Vertical Dimensions 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 D-42 Revision 1 
Figure D-22. Comparisons of Stress Paths for Models with Different Vertical Dimensions 
Source: BSC 2004a, Figure 6.7-14. 
June 2004 D-43 No. 4: Mechanical Degradation Source: BSC 2004a, Figure 6.7-15. 
Figure D-23. Comparisons of Major Principal Stresses under Seismic Ground Motions for Models with 
Different Vertical Dimensions 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 D-44 Revision 1 
D.4.4.2.4 Effects of Model Dimension on Predicted Strength to Stress Ratios for the 
Rock Mass Surrounding Emplacement Drifts 
Contours of strength-to-stress ratios for emplacement drifts from different model dimensions are 
compared in Figures D-24 and D-25 at 50 years after waste emplacement. The minimum 
strength to stress ratio ranges from about 2 for drifts in the weakest lithophysal rock (Category 1) 
to greater than 4 in the strongest lithophysal rock (Category 5). In the pillar, the factor of 
strength to stress ratio is well above 5 for the cases analyzed. The results are consistent among 
all models and are not sensitive to the model sizes. 
June 2004 D-45 No. 4: Mechanical Degradation Source: BSC 2004a, Figure 6.7-18. 
Figure D-24. Comparisons of Strength-to-Stress Ratios at 50 Years after Waste Emplacement in Models 
with Different Vertical Dimensions (Category 1, K0 = 0.5) 
No. 4: Mechanical Degradation D-46 
Revision 1 
June 2004 Revision 1 
Source: BSC 2004a, Figure 6.7-27. 
with Different Vertical Dimensions (Category 5, K0 = 0.5) 
D.4.4.2.5 
Figure D-25. Comparisons of Strength-to-Stress Ratios at 50 Years after Waste Emplacement in Models 
June 2004 
Effects of Model Dimension on Predicted Axial Force in Bolts 
Axial forces in various bolts at different stages of the repository preclosure period are illustrated 
in Figures D-26 and D-27. By comparing the plots from different models, it can be seen that the 
predicted axial forces from these models are generally consistent, and the difference is 
negligible. 
D-47 No. 4: Mechanical Degradation Source: BSC 2004a, Figure 6.7-30. 
Figure D-26. Comparisons of Axial Forces (N) in Bolts at 50 Years after Waste Emplacement in Models 
with Different Vertical Dimensions (Category 1, K0 = 0.5) 
No. 4: Mechanical Degradation D-48 
Revision 1 
June 2004 Source: BSC 2004a, Figure 6.7-33. 
Figure D-27. Comparisons of Axial Forces in Bolts at 50 Years after Waste Emplacement in Models with 
Different Vertical Dimensions (Category 5, K0 = 0.5) 
No. 4: Mechanical Degradation D-49 
Revision 1 
June 2004 Revision 1 
D.5 REFERENCES 
D.5.1 Documents Cited 
Board, M. 2003. Resolution Strategy for Geomechanically-Related Repository Design and 
Thermal-Mechanical Effects (RDTME). REV 00. Las Vegas, Nevada: Bechtel SAIC Company. 
ACC: MOL.20030708.0153. 
BSC (Bechtel SAIC Company) 2001. Ground Control for Emplacement Drifts for SR. 
ANL-EBS-GE-000002 REV 00 ICN 01. Las Vegas, Nevada: Bechtel SAIC Company. 
ACC: MOL.20010627.0028. 
BSC 2002. Software Code: FLAC3D. V2.1. PC/WINDOWS 2000/NT 4.0. 10502-2.1-00. 
D.4.4.3 Effect of Model Dimension on Rockfall Predictions Discontinuum Modeling 
The dimension of the three-dimensional discontinuum models becomes an important issue due to 
the large run times required for conducting the dynamic rockfall analyses. The postclosure 
rockfall study utilized a base-case 3DEC discontinuum model with a size of 25  27.5  25 m 
(BSC 2004b) (Section 5.3.2.1.2 of this technical basis document). The region with detailed 
fractures imported from the FracMan model is one diameter at the side of the drift and two 
diameters on top of the drift. Analyses with three model dimensions were conducted to 
investigate the sensitivity of rockfall prediction to model. 
The study indicates that the distinct blocks increase as the model dimension increases. The 
model, with a dimension smaller than that of the base-case model, underestimates the amount of 
rockfall, whereas the base-case and the larger models predict roughly the same amount of 
rockfall. Generally speaking, the base case appears to be adequate to provide a reasonable 
answer for rockfall (BSC 2004b, Section 6.3.1.6.4). 
June 2004 
D.4.4.4 Summary of Analyses for Resolution of RDTME 3.13 
The comparisons presented above indicate that use of a model with a vertical dimension of 
100 m, in conjunction with the thermal-mechanical boundary conditions described here is 
adequate to minimize boundary influence. Additional increase in the model dimension beyond 
this level does not significantly improve the accuracy of prediction of emplacement drift 
performance. Reduction in the total model vertical dimension from 100 m to 50 m results in an 
overestimate of the drift closures and stresses near the drift opening. A total model dimension of 
100 m was used as the base case condition in ground support analysis studies presented in 
Evaluation of Emplacement Drift Stability for KTI Resolutions (BSC 2004a, Section 6.7. 
In conclusion, assuming a free-moving boundary located at 50 m above the drift centerline, 
coupled with importing temperatures from a large scale ventilation model, are sufficient to 
minimize both mechanical and thermal boundary effects. Use of the 100 m total model 
dimension will not result in an underestimate of potential failure mechanisms of emplacement 
drifts, and therefore, is justified. 
D-50 No. 4: Mechanical Degradation Revision 1 
BSC 2003a. Underground Layout Configuration. 800-P0C-MGR0-00100-000-00E. Las Vegas, 
Nevada: Bechtel SAIC Company. ACC: ENG.20031002.0007. 
BSC 2003b. Scoping Analysis on Sensitivity and Uncertainty of Emplacement Drift Stability. 
800-K0C-TEG0-00600-000-000. Las Vegas, Nevada: Bechtel SAIC Company. 
ACC: ENG.20031125.0002. 
BSC 2003c. Ground Control for Emplacement Drifts for LA. 800-K0C-TEG0-00100-000-00A. 
Las Vegas, Nevada: Bechtel SAIC Company. ACC: ENG.20031016.0001. 
BSC 2004a. Evaluation of Emplacement Drift Stability for KTI Resolutions. 800-KMC-SSE0- 
00200-000-00A. Las Vegas, Nevada: Bechtel SAIC Company. ACC: MOL.20040510.0199. 
BSC 2004b. Drift Degradation Analysis. ANL-EBS-MD-000027 REV 03A. Las Vegas, 
Nevada: Bechtel SAIC Company. ACC: MOL.20040513.0081. 
BSC 2004c. Ground Control for Non-Emplacement Drifts for LA. 800-KMC-SSD0-00700-000- 
00A. Las Vegas, Nevada: Bechtel SAIC Company. ACC: ENG.20040302.0022. 
BSC 2004d. Ventilation Model and Analysis Report. ANL-EBS-MD-000030 REV 03 ICN 03, 
Errata 2. Las Vegas, Nevada: Bechtel SAIC Company. ACC: DOC.20031216.0002; 
DOC.20040202.0004; DOC.20040325.0003. 
Hoek, E. 2000. Practical Rock Engineering, 2000 Edition. Toronto, Ontario, Canada: 
RocScience. TIC: 253544. 
Itasca Consulting Group 2002. Itasca SoftwareCutting Edge Tools for Computational 
Mechanics. Minneapolis, Minnesota: Itasca Consulting Group. TIC: 252592. 
LLNL (Lawrence Livermore National Laboratory) 2002. Software Code: NUFT. V3.0s. 
Sun/SunO.S. 5.6 & 5.7. 10088-3.0s-01. 
NRC (U.S. Nuclear Regulatory Commission) 2002. Integrated Issue Resolution Status Report. 
NUREG-1762. Washington, D.C.: U.S. Nuclear Regulatory Commission, Office of Nuclear 
Material Safety and Safeguards. TIC: 253064. 
Reamer, C.W. and Williams, D.R. 2001. Summary Highlights of NRC/DOE Technical 
Exchange and Management Meeting on Repository Design and Thermal-Mechanical Effects. 
Meeting held February 6-8, 2001, Las Vegas, Nevada. Washington, D.C.: U.S. Nuclear 
Regulatory Commission. ACC: MOL.20010307.0511 through MOL.20010307.0521. 
Sun, Y. 2002. Ground Control Methodology for Emplacement Drifts. TDR-GCS-GE-000002 
REV 00. Las Vegas, Nevada: Bechtel SAIC Company. ACC: MOL.20021118.0097. 
Williams, N.H. 2002. Thermal Inputs for Evaluations Supporting TSPA-LA. Interoffice 
memorandum from N.H. Williams (BSC) to Distribution, September 16, 2002, 0911024159, 
with enclosures. ACC: MOL.20021008.0141. 
June 2004 D-51 No. 4: Mechanical Degradation Revision 1 
D.5.2 Codes, Standards, and Regulations 
10 CFR Part 63. Energy: Disposal of High-Level Radioactive Wastes in a Geologic Repository 
at Yucca Mountain, Nevada. Readily available. 
D.5.3 Data, Listed by Data Tracking Number 
MO0007RIB00077.000. In Situ Rock Conditions. Submittal date: 07/18/2000. 
MO0306MWDALAFV.000. ANSYS-La-Fine Ventilation. Submittal date: 06/23/2003. 
D-52 June 2004 No. 4: Mechanical Degradation SITE-SPECIFIC PROPERTIES OF THE HOST ROCK 
(RESPONSE TO RDTME 3.04) 
No. 4: Mechanical Degradation 
APPENDIX E 
Revision 1 
June 2004 Revision 1 
Note Regarding the Status of Supporting Technical Information 
This document was prepared using the most current information available at the time of its development. This 
Technical Basis Document and its appendices provide Key Technical Issue Agreement responses that were 
prepared using preliminary or draft information reflect the status of the Yucca Mountain Projects scientific 
and design bases at the time of submittal. In some cases this involved the use of draft Analysis and Model 
Reports (AMRs) and other draft references whose contents may change with time. Information that evolves 
through subsequent revisions of the AMRs and other references will be reflected in the License Application 
(LA) as the approved analyses of record at the time of LA submittal. Consequently, the Project will not 
routinely update either this Technical Basis Document or its Key Technical Issue Agreement appendices to 
reflect changes in the supporting references prior to submittal of the LA. 
June 2004 No. 4: Mechanical Degradation Revision 1 
APPENDIX E 
SITE-SPECIFIC PROPERTIES OF THE HOST ROCK 
(RESPONSE TO RDTME 3.04) 
This appendix provides a response for Key Technical Issue (KTI) agreement Repository Design 
and Thermal-Mechanical Effects (RDTME) 3.04. This agreement relates to the physical, 
thermal, and mechanical properties of the host rock used as part of license application activities. 
E.1.1 RDTME 3.04 
Agreement RDTME 3.04 was reached during the U.S. Nuclear Regulatory Commission 
(NRC)/U.S. Department of Energy (DOE) Technical Exchange and Management Meeting on 
Repository and Design Thermal-Mechanical Effects held February 6 to 8, 2001, in Las Vegas, 
Nevada (Reamer and Williams 2001). There has been no submittal related to this KTI agreement 
to the NRC. 
The wording of the agreement is as follows: 
RDTME 3.04 
Provide in the Design Parameter Analysis Report (or some other document) sitespecific 
properties of the host rock, as a minimum those included in the NRC 
handout, together with the spatial and temporal variations and uncertainties in 
such properties, as an update to the information contained in the March 1997 
Yucca Mountain Site Geotechnical Report. The DOE will: (1) evaluate the 
adequacy of the currently available measured and derived data to support the 
potential repository licensing case and identify areas where available data may 
warrant additional field measurements or testing to reduce uncertainty. DOE will 
provide a design parameters analysis report (or other document) that will include 
the results of these evaluations, expected to be available to NRC in FY 2002; and 
(2) acquire data and/or perform additional analyses as necessary to respond to the 
needs identified in (1) above. The DOE will provide these results prior to any 
potential license application. 
RDTME Agreement 3.04 appears in the NRCs Integrated Issue Resolution Status Report (NRC 
2002, Appendix A). This agreement focuses on providing an update to the project rock 
properties geotechnical database in the form of a geotechnical parameters report. Specific NRC 
concerns include addressing the uncertainty, spatial variability, and temporal variability of these 
rock properties at the repository site and identifying where additional data gathering activities or 
analyses or both may be warranted to reduce the existing uncertainties. The NRC also requested 
an evaluation of the adequacy of the project geotechnical database for supporting the repository 
licensing case. 
E.1 KEY TECHNICAL ISSUE AGREEMENT 
E-1 June 2004 No. 4: Mechanical Degradation Revision 1 
E.1.2 Related Key Technical Issue Agreements 
RDTME 3.05 (Appendix A)This KTI agreement requires the characterization of the rock mass 
mechanical properties of lithophysal rock, which includes a description of the uncertainty and 
variability of these properties. Section 3.2.1 of this technical basis document summarizes the 
site-specific properties of lithophysal rock and Section 4.2 discusses the development of rock 
mass properties for lithophysal rock. The spatial variability of lithophysal rock properties based 
on their correlation with rock porosity is addressed in Sections 2.3.2 and 5.3.2.2. 
RDTME 3.07This KTI agreement requires a description of how mechanical intact rock 
strength is expected to vary over time as a result of sustained loading. The response to the 
agreement provides a summary of laboratory testing, field investigations, scientific and 
engineering analyses, and modeling efforts used in estimating and confirming the timedependent 
performance of rock mass and rock fracture strength. Section 5.3.2.2.4 discusses 
many of these issues, and the agreement will be fully addressed separate from this technical basis 
document. 
RDTME 3.08 (Appendix F)This KTI agreement requires a description of the uncertainty and 
variability of the fracture geometry in nonlithophysal and lithophysal rocks and a corresponding 
sensitivity analysis with respect to fracture geometry to establish design uncertainty for ground 
support design and drift degradation estimates. Sections 3.2, 4.1, and 5.3 discuss these issues. 
RDTME 3.15, 3.16, 3.17, and 3.19 (Appendix C)These KTI agreements require a summary of 
in situ fracture geometry measurements, including small trace-length fracture data, and a 
description of the representative modeling of fractures for the repository horizon. They also 
require a discussion of the uncertainties in the thermal and mechanical properties of rock blocks 
and fractures, including their long-term degradation. Sections 3.2, 4.1, and 5.3 discuss these 
issues. 
June 2004 
E.2 RELEVANCE TO REPOSITORY PERFORMANCE 
The license application will include an adequate summary of thermal-mechanical properties 
needed for subsurface design, preclosure repository safety analysis, and repository performance 
after permanent closure. It is necessary to demonstrate that data inputs for license application are 
a traceable to an appropriate source, are transparent to users of the data, and are sufficient for 
their intended use. In particular, the NRC regulations employ a risk analysis approach for 
license application requires that rock property inputs have defined uncertainties, as well as 
established spatial and temporal variations that can be propagated through the products that 
support the repository performance licensing case. 
To facilitate meeting these objectives, a preliminary collection of site-specific subsurface 
geotechnical properties and parameters was made and evaluated for its adequacy to support 
license application. As a result of this evaluation, additional laboratory, field, computational, 
and geostatistical work was initiated to improve the geotechnical database for lithophysal rock, 
to better characterize the spatial representations of fractures and rock properties for the 
repository, and to describe the temporal variation of rock parameters. While this work was being 
accomplished, a comprehensive effort was made to improve the defensibility of the historical 
E-2 No. 4: Mechanical Degradation Revision 1 
data and to summarize the important contextual information associated with particular rock 
parameters in order to ensure appropriate use of the parameters by users of the data. Finally, a 
current summary of site-specific subsurface geotechnical parameters was prepared by the DOE, 
including a summary of parameter uncertainties and variability, which updates data in Yucca 
Mountain Site Geotechnical Report (CRWMS M&O 1997). The updated summary of acquired 
and developed geotechnical properties was issued as Subsurface Geotechnical Parameters 
Report (BSC 2003a) and is supplemented by information found in Drift Degradation Analysis 
(BSC 2004a). Sections 3 and 4 of this technical basis document summarize the characterization 
of these rock properties. 
E.3 RESPONSE 
DOE has evaluated the adequacy of the previously available data, acquired additional data to 
reduce data uncertainty, and published an update of geotechnical data and parameters in 
Subsurface Geotechnical Parameters Report (BSC 2003a). This section describes actions 
leading to its development. 
In 1997, the DOE issued Yucca Mountain Site Geotechnical Report (CRWMS M&O 1997). 
Included in the report were the site-specific data resulting from field and laboratory testing 
available through June 1996 (CRWMS M&O 1997, p.1-1). This report evaluated the sufficiency 
of the existing geotechnical data available in the geotechnical database for a viability assessment 
of constructing a repository at Yucca Mountain, and for license application. 
In the ensuing years, substantial effort was made to acquire the additional data considered 
necessary to support the design and analyses for the license application. These investigative 
activities encompassed a range of data obtained from a number of sources including field 
measurements, laboratory testing, in situ testing, numerical modeling, and analytical assessment. 
Results of historical as well as recent activities were summarized in Subsurface Geotechnical 
Parameters Report (BSC 2003a). The report incorporates the additional data available to the 
project as of the summer of 2003. The primary source data are based on qualified data obtained 
within a framework of the DOE quality assurance procedures. Where available, additional 
(corroborative) data are also presented in Subsurface Geotechnical Parameters Report (BSC 
2003a). 
The general properties focused on in Subsurface Geotechnical Parameters Report (BSC 2003a) 
included fracture geometry parameters; rock density and porosity data; intact rock thermal and 
mechanical parameters; rock mass quality estimates; and estimated rock mass physical, thermal, 
and mechanical parameters. Site geology, stratigraphy, stratigraphic nomenclature, and 
lithostratigraphic structural features are also included to provide a framework for presentation of 
the rock and rock mass property data. 
Subsurface Geotechnical Parameters Report (BSC 2003a) presents simple statistical summaries 
of parameters where sufficient data are available. Where an incomplete or inadequate 
understanding of rock behavior and parameters existed, other approaches were used to gain a 
deeper understanding of rock behavior and to estimate the uncertainty of rock unit parameters. 
For example, to better understand the mechanical behavior of lithophysal rock, a conceptual 
model of lithophysal rock was devised, implemented within suitable computational codes, and 
June 2004 E-3 No. 4: Mechanical Degradation Revision 1 
used in addition to the large-core test results. In particular, the large core mechanical tests 
(including room dry and saturated conditions) and computational results were used to relate 
strength and modulus to lithophysal porosity, which is the primary factor controlling the 
variability of the mechanical behavior in lithophysal rocks. The spatial variation of mechanical 
rock properties was then developed using a surrogate porosity process in which field 
measurements of lithophysal voids were used to build a simple geostatistical model that was then 
coupled with the propertyporosity correlations. Preliminary modeling of time degradation of 
mechanical rock parameters is presented in Section 5.3.2.2.4 and is fully addressed by response 
to RDTME 3.07, to be provided separately. 
A general discussion of the adequacy of geotechnical parameters for underground design and 
modeling purposes is given in Subsurface Geotechnical Parameters Report (BSC 2003a). The 
updated summary of laboratory tests, field tests, field mapping, construction records, numerical 
modeling, and geostatistical modeling provides sufficient supporting rock data for design and 
analysis needs. Several key geotechnical analysis and modeling reports support the license 
application. The adequacy of the geotechnical database for these analyses is discussed in 
Sections 5 and 6 and, individually, in the following reports: 
 Drift Degradation Analysis (BSC 2004a) 
 Ground Control for Emplacement Drifts for LA (BSC 2003b) 
 Ground Control for Non-Emplacement Drifts for LA (BSC 2004b) 
 Scoping Analysis on Sensitivity and Uncertainty of Emplacement Drift Stability (BSC 
2003c). 
The information in this report is responsive to agreement RDTME 3.04 made between the DOE 
and NRC. The report contains the information that DOE considers necessary for NRC review 
for closure of this agreement. 
E.4 BASIS FOR THE RESPONSE 
The scope and methodology for achieving an adequate set of site-specific data for characterizing 
the complex subsurface conditions was progressively developed from a series of geotechnical 
review panel meetings held in 2001 and 2002, and from work proposals prepared and reviewed 
by its participants. The review panel consisted of BSC personnel and geomechanics experts 
from Sandia National Laboratories, U.S. Bureau of Reclamation, U.S. Geological Survey, 
University of Nevada, Las Vegas, University of Nevada, Reno, New England Research, Itasca 
Consulting Group, and Nick Barton and Associates. The DOE approach for addressing the 
identified geomechanical needs is summarized in Section 1.2.3 of this technical basis document 
and in Resolution Strategy for Geomechanically-Related Repository Design and 
Thermal-Mechanical Effects (RDTME) (Board 2003). 
The overall approach adopted to resolve these geomechanical issues consisted of obtaining new 
laboratory and field data, supplemented by utilizing a combination of analyses, studies, and 
June 2004 E-4 No. 4: Mechanical Degradation Revision 1 
calculations to maximize the utility of the available site-specific data. The following steps were 
identified and taken: 
1. Collection and geotechnical analysis of the existing (2001) geological and 
geotechnical characterization data for suitability to support the analyses listed in E.3. 
The analysis identified additional laboratory and in situ thermal-mechanical testing 
needs. 
2. Development and execution of a laboratory and field-testing plan. This consisted of 
extensive mapping of features within Yucca Mountains excavated tunnels, the 
mechanical testing of lithophysal rock given the limited data set existing in 2001, 
augmenting the thermal-mechanical test database, and planning for time-dependent 
testing needed to better characterize the temporal variation of rock properties. 
3. Development, calibration, and validation of numerical models capable of representing 
the thermal-mechanical behavior of lithophysal and nonlithophysal rocks and 
geostatistical models to characterize the geologic spatial variability (lithophysal 
porosity and fracture geometry). 
4. Utilization of these computational models to supplement the material properties 
database and to explore the impact of spatial variability of properties on the 
geomechanical response in rock, primarily lithophysal rocks. 
5. Exploration by means of numerical sensitivity studies to further investigate the impact 
of parameter uncertainty on preclosure ground support and postclosure drift 
degradation and seismic stability. 
A preliminary 2001 summary of site-specific geotechnical properties of the host rock appears in 
Subsurface Geotechnical Parameters Report (BSC 2003a, Attachment IV). A qualitative 
analysis of this summary resulted in an assessment of data needs required for license application. 
These needs and a description of new measurement data and analyses are described below. 
Field Collection of Fracture Geometry and Rock Mass Quality DataFor nonlithophysal 
rock, field characterization was carried out using a large database of meter-scale fracture 
geometry data produced from mapping the ESF and ECRB tunnels. Evaluation of this database 
of the tunnel full-periphery structure maps produced a statistical database, by fracture set, of 
fracture geometry data, such as fracture orientation, spacing, and trace length (BSC 2003a, 
Section 8.8.2; BSC 2004a, Sections 4.11 and 6.14, Appendix B). A secondary aspect of the 
program involved collecting rock quality data for rock classification purposes for 
lithostratigraphic units (BSC 2003a, Section 8.8.3) and mapping smaller scale fractures (less than 
1 m) in the ESF and ECRB tunnels. Section 2.3.1 of this technical basis document discusses the 
characterization of fractures. 
Field Characterization of Lithophysal Rock FeaturesFor lithophysal rock within the 
repository host horizon, the primary aspect of field characterization involved mapping the shape, 
size, and abundance of lithophysae, spots, and rims (BSC 2003a, Section 8.8.4). This augmented 
the existing field characterization from geophysical borehole logging to indirectly determine 
June 2004 E-5 No. 4: Mechanical Degradation Revision 1 
vertical variation of bulk density and porosity data. Section 2.3.2 discusses the characteristics 
and distribution of lithophysae. 
Laboratory Characterization of Nonlithophysal Rock PropertiesLaboratory tests were 
performed on intact core samples of nonlithophysal rock. This laboratory database has been 
analyzed to identify the key factors impacting the thermal and mechanical properties of intact 
rock (BSC 2003a, Sections 8.2, 8.3, and 8.4). The general mechanical and thermal behavior of 
nonlithophysal rock is adequately understood and sufficient parameter summaries have been 
prepared. Direct shear testing of five fractures was performed to help characterize the 
mechanical behavior of fractures by fracture set in nonlithophysal rock (BSC 2003a, 
Section 8.6). Section 3.2.1 discusses the mechanical properties of intact rock, and Section 3.2.4 
discusses the mechanical properties of fractures. 
Laboratory Characterization of Lithophysal Rock PropertiesSeveral important thermal and 
mechanical rock properties are strongly dependent on rock porosity and lithophysal rock sample 
size. The mechanical elastic and rock strength parameters and thermal conductivity represent 
typical examples. Additional laboratory testing has been conducted to better characterize the 
thermal intact rock properties of thermal conductivity, heat capacity, and coefficient of thermal 
expansion (BSC 2003a, Section 8.3). Additional mechanical testing of large core specimens of 
lithophysal rock was undertaken to better characterize rock parameter dependencies on porosity 
and size. Results and analysis are presented in Subsurface Geotechnical Parameters Report 
(BSC 2003a, Section 8.4) and in Lithophysal Rock Mass Mechanical Properties of the 
Repository Host Horizon (BSC 2004c). However, sampling difficulties and laboratory testing 
limitations made it impractical to adequately characterize of lithophysal rock parameters, so 
computational methods were developed to augment the limited data. Section 3.2.1 discusses the 
mechanical properties of intact rock, including correlation of these properties with rock porosity. 
In Situ Field TestingAdditional thermal and mechanical field testing was planned and carried 
out along the ESF and ECRB Cross-Drift to improve characterization of rock properties and 
validation of property estimates (BSC 2002a). A summary of in situ thermal and mechanical 
rock tests is reported in Subsurface Geotechnical Parameters Report (BSC 2003a, Sections 8.3 
and 8.7). The in situ testing program extends from small-scale borehole testing of thermal and 
mechanical behaviors to meter-scale mechanical and thermal testing of rock blocks (e.g., the 
Single Heater Test, plate loading tests, slot tests) to drift-scale tests in both repository host 
horizon nonlithophysal (Drift Scale Test) and lithophysal rock (planned for the future for 
purposes of confirmation). In particular, in situ lithophysal rock testing (slot tests) was necessary 
to help confirm models and ranges of rock behavior. Section 3.2.3 discusses the in situ 
mechanical testing of nonlithophysal and lithophysal rocks. 
Analysis and Modeling of Thermal Rock Mass PropertiesNew analyses of thermal 
laboratory and field measurements have been carried out to better characterize and model 
thermal properties. The most important of these recent efforts include Thermal Testing 
Measurements Report (BSC 2002a), Heat Capacity and Thermal Expansion Coefficients 
Analysis Report (BSC 2003d), Thermal Conductivity of the Potential Repository Horizon Model 
Report (BSC 2002b), and Thermal Conductivity of Non-Repository Lithostratigraphic Layers 
(BSC 2004d). The conceptual, analytical, and numerical modeling of thermal behavior of both a 
nonlithophysal and lithophysal rock mass is relatively well developed and is summarized in the 
June 2004 E-6 No. 4: Mechanical Degradation Revision 1 
Subsurface Geotechnical Parameters Report (BSC 2003a, Section 8.3) for the rock mass thermal 
properties of thermal conductivity, heat capacity, and coefficient of thermal expansion. 
Section 3.2.5 discusses the thermal properties of the repository host rocks. 
Derived Estimate of Nonlithophysal Rock Mass PropertiesThe mechanical rock mass 
behavior of nonlithophysal rock is controlled largely by the geometry of fractures that separates 
the relatively strong and stiff pieces of intact rock. The rock mass parameters of the repository 
host horizon nonlithophysal rock from tunnel data have been developed from an established 
empirical approach and are summarized in Subsurface Geotechnical Parameters Report (BSC 
2003a, Section 8.5). A statistically representative fracture model of the repository site was 
required for addressing the spatial variability of properties and for meeting design and analysis 
needs. Fracture model results have been developed using the FracMan program and are reported 
in Subsurface Geotechnical Parameters Report (BSC 2003a, Section 9.3) and Drift Degradation 
Analysis (BSC 2004a, Section 6.1.6 and Appendix B). 
Derived Estimate of Lithophysal Rock Mass PropertiesThe mechanical rock mass behavior 
of lithophysal rock is controlled largely by stress conditions and voids within the rock that can 
lead to failure of the bulk lithophysal rock. Lithophysal Rock Mass Mechanical Properties of the 
Repository Host Horizon (BSC 2004c) provides a detailed analysis of the relationship between 
porosity and mechanical rock properties and develops estimates of the rock mass mechanical 
properties of the lithophysal units of the Topopah Spring Tuff. The strength and modulus 
base-case estimates and ranges are developed based on laboratory testing of lithophysal rock and 
supplemented by numerical modeling. Numerical modeling of lithophysal rock was used to 
study the dependence of rock behavior on sample size, and lithophysal geometry (e.g., shape, 
size, and spatial distribution of voids). The development, calibration, and validation of these 
numerical models of lithophysal rock are summarized in the above report and Subsurface 
Geotechnical Parameters Report (BSC 2003a, Sections 9.1 and 9.2). These numerical 
approaches are shown to be successful in reproducing observed complex rock behavior, such as 
allowing for the development of interlithophysae fracturing. Section 4.2 and the response to 
RDTME 3.05 (Appendix A) provides further explanation of how lithophysal rock mass 
properties were developed. 
Spatial Variation of the Mechanical Properties of Lithophysal RockThe spatial variability 
of lithophysal rock mechanical parameters could not be directly developed from laboratory and 
field data. Instead, an indirect approach was adopted that combines the modeled spatial variation 
of lithophysal porosity in the field with the previously developed correlations between rock 
porosity and its mechanical properties. A preliminary simulation of the spatial variation of 
lithophysal porosity has been completed and is described in Subsurface Geotechnical 
Parameters Report (BSC 2003a, Section 9.4). Further discussion of this surrogate property 
approach for estimating the spatial variability of mechanical and thermal properties is described 
in Subsurface Geotechnical Parameters Report (BSC 2003a, Section 10.2). The application of 
this approach for lithophysal rock is developed in Lithophysal Rock Mass Mechanical Properties 
of the Repository Host Horizon (BSC 2004c). The spatial variability of lithophysal rock 
properties based on their correlation with rock porosity is addressed in Sections 2.3.2 and 5.3.2.2. 
Temporal Variation of the Mechanical Properties of Lithophysal RockEstimates of the 
temporal variation in rock properties will be addressed separately in conjunction with 
June 2004 E-7 No. 4: Mechanical Degradation Revision 1 
RDTME 3.07. A discussion of static fatigue testing and estimates of long-term strength is 
provided in Drift Degradation Analysis (BSC 2004a, Appendix S). This is based on laboratory 
static-fatigue testing of lithophysal rock as well as reliance on numerical modeling. The 
numerical modeling program is required due both to the limited number of rock samples tested 
under sustained loading and the relatively short time available in laboratories relative to the 
10,000-year prediction of behavior. Conceptual models have been developed and implemented 
within computational codes to better understand and predict the time-related mechanical 
behavior of lithophysal rock. For nonlithophysal rock, the temporal variation in mechanical rock 
mass behavior is assumed to substantially relate to the degradation of the fracture strength 
parameters. A summary of laboratory static fatigue testing, as well as time-dependent numerical 
modeling and residual fracture strength data will be provided in the next revision of Subsurface 
Geotechnical Parameters Report (BSC 2003a) and will be specifically addressed in the response 
to RDTME 3.07. A summary of the long-term loading effect on rock strength is discussed in 
Section 5.3.2.2.4. 
ConclusionsThe uncertainty related to several key thermal and mechanical geotechnical 
properties has been substantially reduced as a result of additional data that have been collected 
and analyzed since 1997. This new work is summarized in Subsurface Geotechnical Parameters 
Report (BSC 2003a), Lithophysal Rock Mass Mechanical Properties of the Repository Host 
Horizon (BSC 2004c), and Drift Degradation Analysis (BSC 2004a) and specifically includes: 
 Updated analysis of geophysical bulk density and porosity from Yucca Mountain 
boreholes 
 Updated analysis of rock mass quality ratings and classifications of the underground 
openings 
 Field mapping and analysis of lithophysae and other features in lithophysal rock 
 Field mapping and analysis of small-scale fractures within the repository horizon 
 Updated analysis and modeling of field fracture geometry within the repository horizon 
 Creation of a new database of thermal and mechanical properties of intact rock of the 
Yucca Mountain stratigraphy, with an updated analyses of these intact rock properties 
 Direct shear testing of repository horizon rock fractures and an updated analysis of 
mechanical fracture properties of the Yucca Mountain stratigraphy 
 New estimates of rock mass thermal properties of the Yucca Mountain stratigraphy 
 New estimates of nonlithophysal rock mass mechanical properties of the Yucca 
Mountain stratigraphy 
 Laboratory testing of large cores of lithophysal rock 
 Numerical testing to better characterize the properties of lithophysal rock 
June 2004 E-8 No. 4: Mechanical Degradation Revision 1 
 New estimates of lithophysal rock mass mechanical properties in the repository horizon 
 In situ thermal-mechanical field testing of rock within the repository horizon 
 Static fatigue testing of repository horizon rock 
 Estimates of the spatial variation of rock mass mechanical properties based on the spatial 
variation of lithophysae 
 Estimates of the temporal variation of rock properties based on laboratory results and 
numerical analysis. 
Completion of the above work, together with the anticipated completion of in-progress activities, 
is considered sufficient to reduce thermal and mechanical rock property uncertainties to 
acceptable risk levels for license application. 
BSC 2004b. Ground Control for Non-Emplacement Drifts for LA. 800-KMC-SSD0-00700-000- 
E.5 REFERENCES 
Board, M. 2003. Resolution Strategy for Geomechanically-Related Repository Design and 
Thermal-Mechanical Effects (RDTME). REV 00. Las Vegas, Nevada: Bechtel SAIC Company. 
ACC: MOL.20030708.0153. 
BSC (Bechtel SAIC Company) 2002a. Thermal Testing Measurements Report. ANL-NBS-HS- 
000041 REV 00. Las Vegas, Nevada: Bechtel SAIC Company. ACC: MOL.20021004.0314. 
BSC 2002b. Thermal Conductivity of the Potential Repository Horizon Model Report. 
MDL-NBS-GS-000005 REV 00. Las Vegas, Nevada: Bechtel SAIC Company. 
ACC: MOL.20020923.0167. 
BSC 2003a. Subsurface Geotechnical Parameters Report. 800-K0C-WIS0-00400-000-00A. 
Las Vegas, Nevada: Bechtel SAIC Company. ACC: ENG.20040108.0001. 
BSC 2003b. Ground Control for Emplacement Drifts for LA. 800-K0C-TEG0-00100-000-00A. 
Las Vegas, Nevada: Bechtel SAIC Company. ACC: ENG.20031016.0001. 
BSC 2003c. Scoping Analysis on Sensitivity and Uncertainty of Emplacement Drift Stability. 
800-KOC-TEG0-00600-000-000. Las Vegas, Nevada: Bechtel SAIC Company. ACC: 
ENG.20031125.0002. 
BSC 2003d. Heat Capacity and Thermal Expansion Coefficients Analysis Report. ANL-NBS- 
GS-000013 REV 00. Las Vegas, Nevada: Bechtel SAIC Company. 
ACC: DOC.20030820.0002. 
BSC 2004a. Drift Degradation Analysis. ANL-EBS-MD-000027 REV 03A. Las Vegas, NV: 
Bechtel SAIC Company. ACC: MOL.20040513.0081. 
00A. Las Vegas, Nevada: Bechtel SAIC Company. ACC: ENG.20040302.0022. 
June 2004 E-9 No. 4: Mechanical Degradation Revision 1 
BSC 2004c. Lithophysal Rock Mass Mechanical Properties of the Repository Host Horizon. 
800-K0C-SS00-00200-000-00Aa. Las Vegas, Nevada: Bechtel SAIC Company. 
ACC: MOL.20040510.0200. 
BSC 2004d. Thermal Conductivity of Non-Repository Lithostratigraphic Layers. MDL-NBS- 
GS-000006 REV 00, with errata. Las Vegas, Nevada: Bechtel SAIC Company. 
ACC: DOC.20030815.0001; DOC.20040115.0001. 
CRWMS M&O (Civilian Radioactive Waste Management System Management and Operating 
Contractor) 1997. Yucca Mountain Site Geotechnical Report. B00000000-01717-5705-00043 
REV 01. Two volumes. Las Vegas, Nevada: CRWMS M&O. ACC: MOL.19971017.0736; 
MOL.19971017.0737. 
NRC (U.S. Nuclear Regulatory Commission) 2002. Integrated Issue Resolution Status Report. 
NUREG-1762. Washington, D.C.: U.S. Nuclear Regulatory Commission, Office of Nuclear 
Material Safety and Safeguards. TIC: 253064. 
Reamer, C.W. and Williams, D.R. 2001. Summary Highlights of NRC/DOE Technical 
Exchange and Management Meeting on Repository Design and Thermal-Mechanical Effects. 
Meeting held February 6-8, 2001, Las Vegas, Nevada. Washington, D.C.: U.S. Nuclear 
Regulatory Commission. ACC: MOL.20010307.0511 through MOL.20010307.0521. 
June 2004 E-10 No. 4: Mechanical Degradation APPENDIX F 
DESIGN SENSITIVITY AND UNCERTAINTY ANALYSES OF THE FRACTURE 
PATTERNS WITH DISCONTINUUM APPROACH 
(RESPONSE TO RDTME 3.08 AND RDTME 3.12) 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 Revision 1 
Note Regarding the Status of Supporting Technical Information 
This document was prepared using the most current information available at the time of its development. This 
Technical Basis Document and its appendices providing Key Technical Issue Agreement responses that were 
prepared using preliminary or draft information reflect the status of the Yucca Mountain Projects scientific 
and design bases at the time of submittal. In some cases this involved the use of draft Analysis and Model 
Reports (AMRs) and other draft references whose contents may change with time. Information that evolves 
through subsequent revisions of the AMRs and other references will be reflected in the License Application 
(LA) as the approved analyses of record at the time of LA submittal. Consequently, the Project will not 
routinely update either this Technical Basis Document or its Key Technical Issue Agreement appendices to 
reflect changes in the supporting references prior to submittal of the LA. 
June 2004 No. 4: Mechanical Degradation Revision 1 
APPENDIX F 
(RESPONSE TO RDTME 3.08 AND RDTME 3.12) 
This appendix provides a response for Key Technical Issue (KTI) agreements Repository Design 
and Thermal-Mechanical Effects (RDTME) 3.08 and 3.12. These agreements focus on 
U.S. Nuclear Regulatory Commission (NRC) concerns regarding the design sensitivity, 
uncertainty of fracture patterns, and applying site-specific ground motion to the stability of the 
drifts and the ground support system. These agreements are addressed together in this appendix 
due to the similar nature of the issues. 
F.1.1 RDTME 3.08 and RDTME 3.12 
Agreements RDTME 3.08 and 3.12 were reached during the NRC/U.S. Department of Energy 
(DOE) Technical Exchange and Management Meeting on Repository and Design 
Thermal-Mechanical Effects held February 6 to 8, 2001, in Las Vegas, Nevada (Reamer and 
Williams 2001). There has been no submittal related to these KTI agreements to the NRC. 
The wording of these agreements are as follows: 
RDTME 3.08 
Provide the design sensitivity and uncertainty analyses of the fracture pattern 
(with respect to Subissue 3, Component 1). The DOE will provide sensitivity and 
uncertainty analysis of fracture patterns (based on observed orientation, spacing, 
trace length, etc) on the preclosure ground control system design in a revision to 
the Ground Control for Emplacement Drifts for SR, ANL-EBS-GE-000002 (or 
other document) supporting any potential license application. This is expected to 
be available to NRC in FY 2003. 
RDTME 3.12 
Provide dynamic analyses (discontinuum approach) of ground support system 
performance using site specific ground motion time history as input. The DOE 
will provide appropriate analyses to include dynamic analyses (discontinuum 
approach) of preclosure ground support systems, using site specific ground 
motion time histories as input, in a revision to the Ground Control for 
Emplacement Drifts for SR, ANL-EBS-GE-000002 (or other document) 
supporting any potential license application. This is expected to be available to 
NRC in FY 2003. 
DESIGN SENSITIVITY AND UNCERTAINTY ANALYSES OF THE FRACTURE 
PATTERNS WITH DISCONTINUUM APPROACH 
June 2004 
F.1 KEY TECHNICAL ISSUE AGREEMENTS 
F-1 No. 4: Mechanical Degradation Revision 1 
NRC concerns are further elaborated in the Integrated Issue Resolution Status Report 
(NRC 2002, Section 2.1.7) and are paraphrased as follows: 
 Discontinuum models used in the thermal-mechanical analyses for site recommendation 
(CRWMS M&O 2000) were based on a regular fracture pattern composed from the 
mean fracture-set attitudes (dip and dip direction) and spacing, but the uncertainties in 
the fracture-set properties and their effects on the calculated results were not discussed 
(NRC 2002, p. 2.1.7-13). 
 Seismic loading was represented in the models as a sinusoidal velocity history with a 
frequency of 10 Hz, an amplitude equal to the estimated peak ground velocity for the 
site, and a duration of 3 seconds for site recommendation (CRWMS M&O 2000). The 
site-specific ground-motion time history would differ from the model velocity history in 
terms of frequency content, amplitude variation, and duration of loading, so a 
comparison of the two might examine the total energy delivered to the rock in either 
case and the amount of that energy available to cause rock failure (e.g., by fracture slip) 
(NRC 2002, pp. 2.1.7-13 to 2.1.7-14). 
F.1.2 Related Key Technical Issue Agreements 
Agreements RDTME 3.02 and RDTME 3.13 (Appendix D) are related to RDTME 3.08 and 
RDTME 3.12. RDTME 3.02 addresses the critical combination of in situ, thermal, and seismic 
stresses. RDTME 3.13 addresses the technical justification for boundary conditions for the 
numerical models. 
F.3 RESPONSE 
A summary of analyses for resolving agreements RDTME 3.08 and RDTME 3.12 follow. These 
analyses are discussed in Sections 4.1 and 5.3 and covered in detail in Evaluation of 
Emplacement Drift Stability for KTI Resolution (BSC 2004a). The approach for addressing the 
KTIs include: 
 Development of a three-dimensional stochastically-defined synthetic fracture network 
using the FracMan program, and based on the mapped fracture orientation, spacing, and 
F.2 RELEVANCE TO REPOSITORY PERFORMANCE 
The preclosure safety analysis is used to demonstrate the safety of the geologic repository 
operations area with regard to the overall preclosure performance design and operations 
objectives. The safety strategy for the preclosure operating period is to demonstrate that the 
ground control system is not required to prevent or mitigate credible rockfall. This 
demonstration relies upon analyses that show that the waste package does not breach when 
impacted by credible rock blocks. 
The jointed rock mass surrounding the emplacement drifts for nonlithophysal rock will be 
subjected to loadings from in situ stress, thermal loading, and seismic ground motion. The 
results of the sensitivity and uncertainty analyses for fracture patterns provide the design basis 
rock blocks for evaluation of the waste package. 
F-2 June 2004 No. 4: Mechanical Degradation Revision 1 
trace length from repository rock units exposed in the ESF and ECRB Cross-Drift 
excavations. 
 Development of site-specific ground-motion time histories with representative frequency 
content and variability, amplitude variation, and duration of loading for seismic shaking. 
 Development of a three-dimensional dynamic discontinuum (3DEC) model that 
incorporates the synthetic fracture model developed from the FracMan program to 
address the issues of design sensitivity and uncertainty of fracture patterns. A total of 32 
fracture patterns are selected to represent the variability of fracture patterns. 
 Conducting the dynamic discontinuum analyses using site-specific ground-motion time 
histories for unsupported drifts. 
The summary and conclusions from the discontinuum analyses are provided below: 
 A three-dimensional synthetic fracture network was generated for the uncertainty and 
sensitivity analysis of fracture patterns. The generation was based on the observed 
orientation, spacing, and trace length (Section 4.1 of this technical basis document). 
 The 3DEC model, incorporating the stochastically-defined synthetic fracture network 
that addresses the design sensitivity and uncertainty of fracture patterns for preclosure 
analysis, was used to conduct analysis for identification of potential unstable blocks 
(Section 5.3.3 of this technical basis document). Dynamic discontinuum analyses using 
site-specific ground-motion time histories were conducted to provide correct input of 
frequency content, amplitude variation, and duration of loading for seismic shaking 
(BSC 2004a, Section 6.3). 
 Limited areas of unstable blocks were observed in the discontinuum analyses for the 
unsupported openings when subjected to seismic shaking. A total of 32 simulations with 
stochastically-defined fracture patterns were conducted for preclosure ground motions. 
Calculation of the ground support response based on an uncoupled method (i.e., the 
ground support was not modeled directly in the 3DEC analysis) demonstrated that the 
factor of safety of the rock bolt tension capacity versus the median weight of unstable 
wedges is more than 10. The combination of rock bolts and steel sheets provides ample 
support for the extreme scenario with unfavorable block orientation (BSC 2004a, 
Section 6.3). 
The information in this appendix is responsive to agreements RDTME 3.08 and RDTME 3.12 
made between the DOE and NRC. This report contains the information that the DOE considers 
necessary for NRC review for closure of these agreements. 
F.4 BASIS FOR THE RESPONSE 
F.4.1 Development of Three-Dimensional Synthetic Fracture Network Using FracMan 
The description of the fracture characteristics at the repository host horizon is provided in 
Section 2.3.1 of this technical basis document. Relatively short trace length and partially 
June 2004 F-3 No. 4: Mechanical Degradation Revision 1 
nonpersistent joints are observed at the ESF and ECRB Cross-Drift (BSC 2004b, Section 6.1). 
Analysis of seismic response in emplacement drifts in fractured rock is a three-dimensional 
problem requiring the rock mass to be represented as an explicitly fractured assemblage. The 
FracMan program (Section 4.1 of this technical basis document) was used to provide a 
statistically-similar fracture network geometry as compared to that observed in the exploratory 
tunnels. This fracture network was used to develop a representative volume of jointed rock mass 
within which example emplacement drifts could be excavated. The existing fracture mapping 
database provides the basic input to the FracMan program, which develops sets of planar, 
circular fractures that conform to the statistical variability of the geometric characteristics of the 
input data. A detailed description of the synthetic fracture network generation process is 
provided in Drift Degradation Analysis (BSC 2004b, Section 6.1). 
F.4.2 Development of Site-Specific Ground-Motion Time History 
A description regarding the development of the site-specific ground motion time history is 
provided in Development of Earthquake Ground Motion Input for Preclosure Seismic Design 
and Postclosure Performance Assessment of a Geologic Repository at Yucca Mountain, NV 
(BSC 2004c). The site-specific ground-motion time history was developed based on a site 
response model. The modeling approach implements a random-vibration theory and 
equivalent-linear modeling formulation to calculate site response effects on ground motion. 
Resulting seismic velocity time histories for the mean annual exceedance probability of 10-4 are 
shown in Figure F-1. Two horizontal components (H1 and H2) and one vertical component (V) 
of acceleration, velocity, and displacement are supplied, as discussed in Section 5.3.2.1.3 of this 
technical basis document. 
F-4 June 2004 No. 4: Mechanical Degradation Revision 1 
Source: DTN: MO0306SDSAVDTH.000. 
Figure F-1. Time Histories of Velocity Components of Seismic Motion at Repository Horizon, Mean 
Annual Exceedance Probability of 10-4 
June 2004 
F.4.3 Three-Dimensional Discontinuum Analysis of Jointed Rock Mass 
The three-dimensional discontinuum analysis is used for simulating the mechanical behavior of 
the jointed rock mass in the nonlithophysal units for loading conditions in which the stability 
response will be controlled by the fractures. The possibility that wedge-type failure occurs in the 
lithophysal units has been investigated and found to be very small (approximately 1 block per 
kilometer) with consideration of the variation of fracture patterns (BSC 2004b, Section 6.4.3). A 
minor sidewall shear failure mechanism is considered more appropriate for the lithophysal units, 
as discussed in Sections 4.2 and 5.3.3.2 of this technical basis document. 
The 3DEC program was selected for its capability of simulating a jointed rock mass under both 
thermal and seismic loadings (see Section 4.1.1 of this technical basis document). The jointed 
rock mass is represented as a number of intact rock blocks separated by interface planes whose 
mechanical behavior is represented by a standard Coulomb slip criterion. The intact blocks are 
subdivided into tetrahedral finite difference zones and can be assigned a suitable mechanical 
constitutive law (Itasca Consulting Group 2002). Because of the high intact rock strength in the 
nonlithophysal units (Section 3.2.1 of this technical basis document), rock blocks are considered 
to behave elastically. Ground support is not included in the model so that a general stability of 
F-5 No. 4: Mechanical Degradation Revision 1 
the unsupported openings is assessed. Unstable blocks identified in the 3DEC analyses are then 
used in a separate, uncoupled ground support calculation presented in Section F.4.4. 
To account for the heterogeneous nature of the jointed medium, 32 fracture patterns were 
selected from a 100m cube of stochastically-defined FracMan rock mass for preclosure 
consideration (see Section 4.1 of this technical basis document). Justification of conducting 
32 fracture patterns to cover the range of variability of the fracture patterns is provided in Drift 
Degradation Analysis (BSC 2004b, Appendix K). A representative tunnel volume, 
approximately two tunnel diameters around the emplacement drift centroid and 25 m long, is 
created at each of these locations to contain fractures generated in FracMan. This volume is 
considered sufficient to generate an appropriate representation of damaged rock and of sufficient 
length (approximately five times the tunnel diameter and approximately 10 times the median 
joint trace length) to provide a representative volume of unstable blocks for estimation of the 
rockfall hazard. 
One of the 3DEC models including the region with detailed fractures imported from FracMan is 
shown in Figure F-2. Three vertical and five horizontal cross-sectional views are included in 
Figure F-2 to illustrate the fractures and blocks around the excavation. The dominant northwest 
trending subvertical fracture set is clearly shown in the cross-sectional views. That portion of the 
rock mass in which fractures do not form blocks is shown in green, while distinct blocks are 
identified by other colors. Some of the fractures shown in the cross-sectional views were 
artificially generated during the block-cutting process or to facilitate mesh generation and are 
bonded and not allowed to slip (BSC 2004b, Section 6.3.1). The drip shield is represented as a 
stiff block fixed to the invert of the drift. Although the drip shield does not exist during the 
preclosure period, the drip shield block is placed to collect information on the locations and 
relative velocities of the rockfall impact. Modeling parameters, such as the mechanical 
properties of intact rock and joints, in situ stress, and boundary conditions used in the 3DEC 
model, are provided in Drift Degradation Analysis (BSC 2004b, Section 6.3). 
June 2004 F-6 No. 4: Mechanical Degradation Source: BSC 2004a, Figure 6.3-11. 
Figure F-2. 3DEC Model Geometry and Cross Sections 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 F-7 Revision 1 
Figure F-3 compares the input ground motion (H1) with the recorded velocities at the center of 
the model. The results confirm the correct wave inputs and proper wave propagation in the 
3DEC model. Time histories of normal and shear stresses for joints close to the opening were 
recorded during seismic shaking in the 3DEC model. The stress paths of selected fracture 
subcontacts are plotted against the Coulomb slip criterion, as shown in Figure F-4. The fracture 
properties assigned to the model are reviewed in Section 3.2.4 of this technical basis document. 
The coordinates of the subcontacts and their relative location to the drift are presented in 
Figure F-5. The in situ stress state before seismic shaking is also included as the orange square 
for each subcontact location in Figure F-4. The in situ stress state at subcontact a is very close to 
the Coulomb slip criterion. The fracture containing this subcontact has undergone yielding 
during seismic motions. Seismic shaking-induced normal and shear stress are in the range of 
6 MPa. The stress paths for fractures containing subcontacts b and c are well under the yield 
criterion and representative of most of the rock mass. 
Source: BSC 2004a, Figure 6.3-13. 
Figure F-3. Comparison of Input Seismic Wave and Recorded Velocities (H1) in the 3DEC Model 
June 2004 F-8 No. 4: Mechanical Degradation Source: BSC 2004a, Figure 6.3-14. 
Figure F-4. Stress Path at Selected Fracture Subcontacts (3DEC Simulation 16) 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 F-9 Revision 1 
Source: BSC 2004a, Figure 6.3-15. 
Figure F-5. Information for the Subcontacts for Stress Path Presentation (3DEC Simulation 16) 
The unstable blocks resulting from preclosure seismic shaking were identified in the 3DEC 
simulations, with results summarized in Table F-1. The total number and volume of unstable 
blocks for each simulation are summed and ranked in order in Table F-1. Figure F-6 presents the 
histogram and the cumulative frequency of occurrence for the size of the unstable blocks. The 
blocks are generally small. The maximum rockfall block mass predicted is 2.72 MT, with a 
median block size of 0.12 MT. The histogram shows it is highly skewed toward small-size 
blocks, which is the same trend depicted for rockfall for postclosure ground motions shown in 
Section 5.3.2.1.6 of this technical basis document. 
June 2004 F-10 No. 4: Mechanical Degradation Table F-1. Total Number and Volume of Unstable Blocks for Each 3DEC Simulation 
Rank Order 
1 
2 
3 
4 
5 
6 
7 
8 
9 
10 
11 
12 
13 
14 
15 
16 
17 
18 
19 
20 
21 
22 
23 
24 
25 
26 
27 
28 
29 
30 
31 
32 
Source: BSC 2004a, Table 6.3-5. 
Synthetic 
Fracture Pattern 
Number 
29 
37 
98 
5 
30 
25 
17 
42 
23 
99 
10 
65 
75 
39 
102 
6 
16 
33 
59 
26 
4 
19 
21 
7 
3 
22 
79 
78 
9 
24 
14 
27 
3DEC 
Simulation 
Case Number 
38 
39 
56 
23 
16 
28 
25 
41 
22 
40 
19 
44 
34 
45 
33 
24 
55 
35 
43 
18 
53 
20 
15 
32 
29 
14 
31 
36 
21 
42 
27 
17 
No. 4: Mechanical Degradation 
Number of 
Unstable Blocks 
62 
29 
28 
22 
35 
16 
16 
13 
15 
17 
13 
22 
23 
5 
9 
10 
10 
5 
15 
17 
9 
4 
5 
1 
4 
3 
2 
6 
5 
3 
3 
1 
F-11 
Revision 1 
Percentile 
100 
97 
94 
91 
88 
84 
81 
78 
75 
72 
69 
66 
63 
59 
56 
53 
50 
47 
44 
41 
38 
34 
31 
28 
25 
22 
19 
16 
13 
9 
6 
3 
Total Volume of 
Unstable 
Blocks (m3) 
7.169 
3.487 
3.433 
3.129 
2.684 
2.481 
2.375 
1.979 
1.686 
1.286 
1.232 
1.186 
0.990 
0.858 
0.709 
0.576 
0.570 
0.565 
0.544 
0.544 
0.405 
0.238 
0.197 
0.181 
0.152 
0.146 
0.140 
0.129 
0.108 
0.098 
0.094 
0.045 
June 2004 Revision 1 
Source: BSC 2004a, Figure 6.3-16. 
Attachment I). 
Figure F-6. Histogram for Block Mass 
The unstable blocks are identified in each simulation. Limited areas of unstable blocks are 
observed for most of the simulations. A typical case is shown in Figure F-7 with 10 unstable 
blocks and a total volume of 0.57 m3. The side view, perspective view, and cross-sectional view 
in through which most of the rockfall occurred are provided. The blocks are small, and most of 
them are clustered in the one location. Figure F-8 presents the results for the worst case in terms 
of the unstable block volume. A total of 62 unstable blocks and a volume of 7.17 m3 are 
predicted for this case. Most of the blocks are clustered at four locations. A cluster of blocks 
with a pyramid shape is formed at the roof crown. The cross section shown in Figure F-8 
intersects this pyramid cluster. A catalog of unstable blocks identified in each simulation is 
provided in Evaluation of Emplacement Drift Stability for KTI Resolutions (BSC 2004a, 
F-12 June 2004 No. 4: Mechanical Degradation Source: BSC 2004a, Figure 6.3-17. 
Figure F-7. Unstable Blocks for Typical Case (Simulation 55, 50 Percentile) 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 F-13 Source: BSC 2004a, Figure 6.3-18. 
Figure F-8. Unstable Blocks for the Case with Greatest Amount of Rockfall (Simulation 38; 100 
Percentile) 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 F-14 Revision 1 
F.4.4 Uncoupled Ground Support Calculation 
Frictional rock bolts and thin perforated steel sheets (held to the rock surface with the rock bolts) 
are selected as the ground support methods for the emplacement drifts (BSC 2003, Section 6.3). 
The design for the ground support system utilizes 3-m-long bolts and perforated steel sheets with 
approximately 240 coverage around the tunnel periphery above the invert. Radially oriented 
rock bolts with faceplates provide the required holding capacity and are placed in an 
approximately 240 coverage pattern around the drift periphery (Figure F-9). Bolts are spaced 
1.25 m apart radially. The longitudinal spacing of the rows of rock bolts is also 1.25 m. Both 
the frictional rock bolts and the perforated steel sheets are made of Stainless Steel Type 316 with 
thickness of 3 mm (equivalent to Stainless Steel Type 316). 
NOTE: The diameter of the rock bolt is 54 mm, and the thickness of the perforated plate is 3 mm. 
Figure F-9. Ground Support Methods Recommended for Emplacement Drifts 
For the uncoupled ground support calculation, six cases were selected to cover the range of 
predicted rockfall caused by preclosure seismic shaking. These six cases have rank order 
numbers 1, 2, 3, 7, 9, and 17 (listed in Table F-1) corresponding to three cases of greatest volume 
of rockfall, two above-average volume cases, and the median volume case. This selection 
F-15 June 2004 No. 4: Mechanical Degradation provides a conservative catalog of unstable blocks for ground support calculation. Figure F-10 
shows the cross section of unstable blocks with superimposed ground supports for each selected 
simulation. 
Source: BSC 2004a, Figure 6.3-20. 
No. 4: Mechanical Degradation 
Revision 1 
Figure F-10. Identified Unstable Blocks with Ground Support 
F-16 June 2004 The uncoupled ground support calculation concerning the bolt capacity assumes free fall of 
unstable blocks due to seismic shaking (i.e., there is no credit taken for the frictional resistance 
along the sliding surface of the block and surrounding rock mass). Thus, the entire weight of the 
block is assumed to be carried by the bolt. The footprint of the blocks on the opening are 
ignored in the calculation (i.e., the blocks may be supported by more than one bolt for some of 
the cases with a larger footprint). 
The weight of the unstable blocks, the calculated bolt resistance capacity, and the calculated 
factor of safety are listed in Table F-2. The design capacity for the rock bolt is specified to be 
20 tons (BSC 2003, Section 6.4). Bolt capacity was adjusted based on the anchored length 
(shown in Figure F-10) and a discount of the capacity due to excavation-induced deformation. A 
discount of 2.25 tons is used based on the coupled numerical analysis reported in Ground 
Control for Emplacement Drifts for LA (BSC 2003, Figure 6-47). The weight of the unstable 
blocks listed in Table F-2 is conservatively estimated based on the 3DEC simulations. For most 
of the cases, the total weight of the unstable blocks predicted in the 25-m drift is used as the 
weight of the blocks in the localized area. The calculations show that the factor of safety for 
most of the cases is in the range of 3 to 5, with the lowest value of 1.6 for the worst case 
(Simulation 38). However, for Simulation 55 (which represents the median stability case among 
all 32 runs completed) the factor of safety is higher than 10. The results clearly indicate that the 
rock bolts alone will provide sufficient support for unstable blocks. 
Table F-2. Uncoupled Load-Capacity Calculation for Rock Bolts 
Anchored Bolt Length (m) Adjusted Bolt Capacity 
(tons) Simulation 
38 
Bolt 1 
11.73 
Bolt 1 
1.98 
13.53 2.29 39 
14.43 2.44 56 
14.43 2.44 25 
14.43 2.44 22 
16.23 2.74 55 
4.61 0.78 38, blocks rotated 
Bolt2 
NA 
NA 
2.44 
2.59 
NA 
NA 
NA 
NA 1.56 25, blocks rotated 
Source: BSC 2004a, Table 6.3-6. 
Consideration was also given to the possibility of the blocks slightly rotating to the most 
unfavorable orientation and the rock bolts penetrating through the peak of the unstable blocks, 
resulting in shorter anchored length (see Figure F-11). Simulations 25 and 38 were selected for 
this assessment because of the depth of the blocks. The calculation is included in Table F-2 with 
a relatively low safety factor outcome for these two rotated cases. Clearly, the adjusted single 
bolt capacity is not sufficient for Simulation 38. The capacity of the perforated steel sheets will 
be required to support the unstable blocks for this scenario. The remaining load required for the 
perforated sheet to support is approximately 2.7 tons. 
9.21 
F-17 No. 4: Mechanical Degradation 
Revision 1 
Factor of 
Safety 
1.6 
3.2 
3.5 
5.2 
3.5 
11.6 
0.6 
Weight of 
Unstable 
Blocks (tons) 
7.2 
4.2 
8.3 
5.7 
4.1 
1.4 
7.2 
5.7 
Bolt2 
NA 
NA 
14.43 
15.33 
NA 
NA 
NA 
NA 1.6 
June 2004 Revision 1 
Source: BSC 2004a, Figure 6.3-21. 
Figure F-11. Unstable Blocks Aligned along the Most Unfavorable Orientation 
To evaluate the performance of the perforated steel sheets against the unstable blocks for 
Simulation 38, the calculation procedure in Ground Control for Emplacement Drifts for LA (BSC 
2003, Section 6.4.3) was used. To simplify the calculation, the pyramid shape was assumed for 
F-18 June 2004 No. 4: Mechanical Degradation the assembly of the unstable blocks. The base dimension l can be calculated with known volume 
and height of the pyramid: 
3 V 
h 
w = W 
A 
= 
where 0.008896 is a constant for conversion from metric ton to meganewton (MN). 
l = 
where V and h is the volume and height of the pyramid, respectively. 
The volume of the block assemblage at the pyramid location is estimated to be 3 m3 based on the 
3DEC simulation results. With a volume of 3 m3 and a height of 2.22 m from Simulation 38 
rotated (Figure F-10), the base dimension l is calculated as 2 m. Figure F-12 shows that the 
actual footprint of the unstable block is larger than the base of the pyramid; hence, the 
calculation result is on the conservative side. 
If the remaining dead load (2.7 MT) is spread uniformly over the entire area of the perforated 
sheet, the unit load is 
. 0 * 7 . 2 008896 = 006 . 0 
22 
F-19 No. 4: Mechanical Degradation 
Revision 1 
(Eq. F-1) 
(Eq. F-2) MN / m2 
June 2004 Source: BSC 2004a, Figure 6.3-22. 
Figure F-12. Top View for the Footprints of the Unstable Blocks Assembly Predicted in 3DEC 
Simulation 38 and the Simplified Pyramid 
The thin-walled, corrugated stainless steel sheets installed in emplacement drifts may be 
conservatively assumed to behave like a flat plate, rigidly fixed on its edges, and subjected to a 
uniform load over its entire area, as stated in Ground Control for Emplacement Drifts for LA 
(BSC 2003, Section 6.4.3). The maximum stress at the center of the plate under a uniform load 
can be estimated using the following expression (Young 1989, p. 464, Case No. 8, Loading Case 
No. 8a): 
wl 2 
t 2 (Eq. F-3) 
June 2004 
= 
F-20 No. 4: Mechanical Degradation 
Revision 1 Revision 1 
where t is the thickness of the steel sheet (3 mm) and is a constant, equal to 0.1386 for the case 
considered. 
Stress in the stainless steel sheets is estimated to be approximately 370 MPa with the base 
dimension l as 2 m, corresponding to the factor of safety of 1.67 based on a tensile strength of 
620 MPa for stainless steel (BSC 2003, Table 3-6). The factor of safety is calculated against the 
stress associated with peak elastic strain in the steel (i.e., the stress state at the onset of inelastic 
hardening). However, it is not the rupture condition since this sheet will have considerable 
inelastic strain prior to rupture. The combination of the rock bolt and steel sheets is, therefore, 
considered adequate even for an extreme scenario with rotated blocks (shown in Figure F-11). 
F.4.5 Conclusions 
The sensitivity and uncertainty analyses for fracture patterns using the site-specific ground 
motion history with the discontinuum model are summarized in Sections F.4.3 and F.4.4. The 
design basis rock blocks have been generated from these analyses for evaluation of waste 
packages. The uncoupled ground support analysis shows that the ground support system is 
adequate with a relatively large margin of safety with consideration to sensitivity and uncertainty 
of the fracture patterns. 
F.5 REFERENCES 
F.5.1 Documents Cited 
BSC (Bechtel SAIC Company) 2003. Ground Control for Emplacement Drifts for LA. 
800-K0C-TEG0-00100-000-00A. Las Vegas, Nevada: Bechtel SAIC Company. 
ACC: ENG.20031016.0001 
BSC 2004a. Evaluation of Emplacement Drift Stability for KTI Resolutions. 800-KMC-SSE0- 
00200-000-00A. Las Vegas, NV: Bechtel SAIC Company. ACC: MOL.20040510.0199. 
BSC 2004b. Drift Degradation Analysis. ANL-EBS-MD-000027 REV 03. Las Vegas, NV: 
Bechtel SAIC Company. ACC: MOL.20040513.0081. 
BSC 2004c. Development of Earthquake Ground Motion Input for Preclosure Seismic Design 
and Postclosure Performance Assessment of a Geologic Repository at Yucca Mountain, NV. 
MDL-MGR-GS-000003 REV 00, with errata. Las Vegas, Nevada: Bechtel SAIC Company. 
ACC: DOC.20031201.0001; DOC.20040401.0004. 
CRWMS M&O (Civilian Radioactive Waste Management System Management and Operating 
Contractor) 2000. Ground Control for Emplacement Drifts for SR. ANL-EBS-GE-000002 REV 
00. Las Vegas, Nevada: CRWMS M&O. ACC: MOL.20000414.0875. 
Itasca Consulting Group 2002. Itasca SoftwareCutting Edge Tools for Computational 
Mechanics. Minneapolis, Minnesota: Itasca Consulting Group. TIC: 252592. 
June 2004 F-21 No. 4: Mechanical Degradation Revision 1 
NRC (U.S. Nuclear Regulatory Commission) 2002. Integrated Issue Resolution Status Report. 
NUREG-1762. Washington, D.C.: U.S. Nuclear Regulatory Commission, Office of Nuclear 
Material Safety and Safeguards. TIC: 253064. 
Reamer, C.W. and Williams, D.R. 2001. Summary Highlights of NRC/DOE Technical 
Exchange and Management Meeting on Repository Design and Thermal-Mechanical Effects. 
Meeting held February 6-8, 2001, Las Vegas, Nevada. Washington, D.C.: U.S. Nuclear 
Regulatory Commission. ACC: MOL.20010307.0511 through MOL.20010307.0521. 
Young, W.C. 1989. Roark's Formulas for Stress and Strain. 6th Edition. New York, 
New York: McGraw-Hill. TIC: 10191. 
F.5.2 Data, Listed by Data Tracking Number 
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. 
F-22 June 2004 No. 4: Mechanical Degradation Revision 1 
APPENDIX G 
ANALYSIS OF ROCK MOVEMENT IN INVERT 
(RESPONSE TO RDTME 3.09) 
June 2004 No. 4: Mechanical Degradation Revision 1 
Note Regarding the Status of Supporting Technical Information 
This document was prepared using the most current information available at the time of its development. This 
Technical Basis Document and its appendices providing Key Technical Issue Agreement responses that were 
prepared using preliminary or draft information reflect the status of the Yucca Mountain Projects scientific 
and design bases at the time of submittal. In some cases this involved the use of draft Analysis and Model 
Reports (AMRs) and other draft references whose contents may change with time. Information that evolves 
through subsequent revisions of the AMRs and other references will be reflected in the License Application 
(LA) as the approved analyses of record at the time of LA submittal. Consequently, the Project will not 
routinely update either this Technical Basis Document or its Key Technical Issue Agreement appendices to 
reflect changes in the supporting references prior to submittal of the LA. 
June 2004 No. 4: Mechanical Degradation Revision 1 
APPENDIX G 
ANALYSIS OF ROCK MOVEMENT IN INVERT 
(RESPONSE TO RDTME 3.09) 
This appendix provides a response for Key Technical Issue (KTI) agreement Repository Design 
and Thermal-Mechanical Effects (RDTME) 3.09. This agreement relates to concerns regarding 
the rock movement in the invert during the repository preclosure period. 
G.1 KEY TECHNICAL ISSUE AGREEMENT 
G.1.1 RDTME 3.09 
Agreement RDTME 3.09 was reached during the U.S. Nuclear Regulatory Commission 
(NRC)/U.S. Department of Energy (DOE) Technical Exchange and Management Meeting on 
Repository and Design Thermal-Mechanical Effects held February 6 to 8, 2001, in Las Vegas, 
Nevada (Reamer and Williams 2001). There has been no submittal to the NRC related to this 
KTI agreement. 
The wording of this agreement is as follows: 
RDTME 3.09 
Provide appropriate analysis that shows that rock movements in the invert are 
either controlled or otherwise remain within the range acceptable to provide for 
retrieval and other necessary operations within the disposal drifts. DOE will 
provide appropriate analysis that shows rock movements in the floor of the 
emplacement drift are within the range acceptable for preclosure operations. The 
analysis results will be provided in a revision to the Ground Control for 
Emplacement Drifts for SR, ANL-EBS-GE-000002 (or other document) 
supporting any potential license application. This is expected to be available to 
NRC in FY 2003. 
The agreement focuses on a concern regarding the rock movement in the invert during the 
repository preclosure period. Due to waste emplacement, emplacement drifts will experience 
elevated temperatures. Similar to the rock mass near the crown and the wall of an emplacement 
drift, the rock in the invert may also move inward due to displacement induced by thermal 
expansion. Excessive rock movement may complicate or preclude retrieval and other operations 
addressed in 10 CFR Part 63. 
G.1.2 Related Key Technical Issue Agreements 
Agreements RDTME 3.02 (Appendix D), RDTME 3.05 (Appendix A), RDTME 3.10 
(Appendix D), and RDTME 3.13 (Appendix D) are related to RDTME 3.09. 
 RDTME 3.02 addresses critical combination of in situ, thermal, and seismic stresses for 
ground support design. The loading combinations, based on the study described in 
resolution of RDTME 3.02 (Appendix D) are used for resolution of RDTME 3.09. 
G-1 June 2004 No. 4: Mechanical Degradation Revision 1 
 RDTME 3.05 addresses rock mass mechanical properties of lithophysal rock. The rock 
mass properties estimate, based on the approach described in resolution of RDTME 3.05 
(Appendix A), is used for resolution of RDTME 3.09. 
 RDTME 3.10 addresses justification for use of two-dimensional modeling for 
emplacement drifts. The two-dimensional models, based on the study described in 
resolution of RDTME 3.10 (Appendix D), are used for resolution of RDTME 3.09. 
 RDTME 3.13 addresses justification for boundary conditions used in modeling for 
emplacement drifts. The boundary conditions, based on the study described in 
resolution of RDTME 3.13 (Appendix D), are used for resolution of RDTME 3.09. 
G.3 RESPONSE 
To respond to agreement RDTME 3.09, an analysis was performed that shows that the rock 
movements in the invert of the emplacement drifts are within acceptable ranges for preclosure 
operations. A summary discussion of the analysis is provided within this appendix, with the 
detailed information covered in Evaluation of Emplacement Drift Stability for KTI Resolutions 
(BSC 2004a). 
The items covered in this appendix to respond to RDTME 3.09 include: 
 Acceptable range of rock deformation in the invert 
 The maximum allowable rock displacements considered in the design of the invert 
G.2 RELEVANCE TO REPOSITORY PERFORMANCE 
The preclosure safety analysis is to be used to demonstrate the safety of the proposed design and 
operations in the geologic repository operations area with regard to the overall preclosure 
performance objectives through a systematic examination of the site, design, potential hazards, 
initiating events and their resulting event sequences, and the potential radiological exposures to 
workers and the public (10 CFR 63.112). The emplacement drifts are an array of horizontal 
tunnels trending at 72 azimuth. Each drift will have a diameter of 5.5 m and will be separated 
from the adjacent drifts by a center-to-center distance of 81 m (BSC 2003; Williams 2002). The 
emplacement drifts provide the subsurface access and openings for the structures, systems, and 
components required for emplacement and retrieval operations. The emplacement area host rock 
provides shielding for the rest of the underground facilities from radiation emanating from the 
waste packages. 
The rock mass surrounding the emplacement drifts, including the invert, is subjected to loads 
induced by in situ stress, thermal loading, and seismic ground motions. Using the anticipated 
load conditions, the rock mass movement around the emplacement drifts during the preclosure 
period is analyzed to evaluate the drift stability and the potential need for invert ground control. 
The rock movement is evaluated to ensure that the retrieval operations are maintained. This 
appendix addresses the movement of the rock mass in the invert during the preclosure period 
only and does not evaluate or discuss rock movement in the postclosure period. 
G-2 June 2004 No. 4: Mechanical Degradation Revision 1 
 Design tolerances for gantry rails, which are related to retrieval operations 
 Methods for the prediction of rock deformation in the invert 
 Predicted rock deformation in the invert 
 Results of the predictions, including the rock deformations in the vertical, lateral, and 
longitudinal directions 
 Comparison of the predicted rock deformations and the design tolerances. 
The conclusions from the analysis (BSC 2004a), which utilizes a two-dimensional numerical 
method to predict the rock deformation in the invert, are: 
 Thermally induced rock expansion is the primary induced force regarding the rock 
deformation in the invert. 
 Preclosure seismically-induced ground motions have insignificant effect on the invert 
rock deformation. 
 Rock movement due to in situ stress release occurring from excavation will be 
equilibrated prior to the installation of the invert structures or gantry rails and will have 
no effect on repository operations. 
 The invert structure (i.e., the ballast material and ground support system) is not designed 
to control the rock deformation in the invert. 
 The predicted rock deformation in the invert is small compared to the design tolerances 
listed in Table G-1. 
Components 
Table G-1. Summary of Design Tolerances for Rock Deformation in Invert 
Laterala 
Verticalb 
Longitudinalc 
Design Tolerance 
(mm) 
28.6 (6.4) 
6.4 
12.7 
Source: BSC 2004a, Section 6.4.1. 
NOTE: aThe numbers in parentheses are the lateral relative displacements per 6 m 
c 
of drift. 
bThe numbers are the vertical relative displacements per 6 m of drift. 
The numbers are the longitudinal relative displacements per 12 m of drift. 
These design tolerances are based on acceptable clearance ranges and rail/gantry crane alignment 
to maintain proper emplacement equipment operations. 
G-3 June 2004 No. 4: Mechanical Degradation Revision 1 
The information in this report is responsive to agreement RDTME 3.09 made between the DOE 
and NRC. This report contains the information that the DOE considers necessary for NRC 
review for closure of this agreement. 
G.4 BASIS FOR THE RESPONSE 
The acceptable range of rock deformations in the invert was determined based on the information 
from the design of invert steel structure and gantry rails (BSC 2004b, pp. 69 and 70; CMAA 
70-2000, Section 1.4.2 and Table 1.4.2-1). The analysis for predicting the rock deformation in 
the invert is a two-dimensional model, as discussed in Section G.4.2, with the primary loading 
condition caused by the thermally induced rock movement in the invert after waste emplacement. 
G.4.1 Acceptable Range of Rock Deformation in Invert 
Rock deformation in the invert, specifically in the lateral, vertical, and longitudinal directions, is 
one of the criteria considered in the design of the invert structure and gantry rails and is reflected 
in the range of specified design tolerances. 
From the invert structure design perspective, the relative lateral and longitudinal displacements 
in the invert need to be predicted in order to appropriately define the design tolerances for the 
invert structure, or steel beams (BSC 2004b, Section 2). The steel invert structure will provide a 
framework, consisting of a series of steel beams bolted to the invert rock mass, that supports the 
emplacement pallets, waste packages, and the drip shields. It will also provide the support for 
the rails that support the gantry crane rails for emplacement and retrieval of waste packages and 
installation of the drip shield. The design tolerances for the steel invert structure are specified as 
follows (BSC 2004b, pp. 69 and 70): 
 Lateral displacement: 28.6 mm (1.13 in.) 
 Axial (drift) displacement: 12.7 mm (0.5 in.) every 12 m (39 ft) of steel beams. 
From the gantry rail design perspective, the predicted relative lateral and vertical displacements 
in the invert are also used to check or define the design tolerances. According to the 
Specifications for Top Running Bridge and Gantry Type Multiple Girder Electric Overhead 
Traveling Cranes (CMAA 70-2000, Section 1.4.2 and Table 1.4.2-1), the design tolerances for 
the gantry rails installed in emplacement drifts are specified as follows: 
 Lateral displacement: 6.4 mm (0.25 in.) every 6 m (20 ft) of rails 
 Vertical displacement: 6.4 mm (0.25 in.) every 6 m (20 ft) of rails. 
G.4.2 Methods for Prediction of Rock Deformation in Invert 
Figure G-1 shows the distributions of time-dependent drift wall temperatures along the 
emplacement drift (BSC 2004a, Figure 6.4-1). The temperatures were developed by the 
Ventilation Model and Analysis Report (BSC 2004c, DTN: MO0306MWDALAFV.000) using a 
continuous, preclosure forced ventilation airflow rate of 15 m3/s. The temperature distributions 
are nearly linear along the drift with respect to time. The maximum temperature difference over 
a 600-m-long drift is about 42C, occurring at 2 years after waste emplacement completion. Due 
G-4 June 2004 No. 4: Mechanical Degradation Revision 1 
to the confinement provided by the rock mass, thermally induced deformation in the longitudinal 
direction will likely be minimal. Using a conservative assumption of a plane strain condition in 
estimating the lateral and vertical deformations in the invert, a two-dimensional approach, in 
which the longitudinal strain is assumed to be zero, is appropriate to predict the rock deformation 
in the invert. 
Source: BSC 2004a, Figure 6.4-1. 
Figure G-1. Distributions of Time-Dependent Drift Wall Temperatures along Emplacement Drift: (a) as a 
Function of Location from the Drift Fresh Air Inlet; (b) as a Function of Time 
The 600-m-long emplacement drift is represented by seven cross-sectional planes, each with a 
thickness of 1 unit (meter) located at distances measured from the drift entrance: 0 m, 100 m, 
200 m, , 600 m. The behavior of these cross-sectional planes are assumed to be independent of 
each other and can be analyzed independently based on the rationale that: (1) two adjacent 
June 2004 G-5 No. 4: Mechanical Degradation Revision 1 
locations are separated by a distance of 100 m, and thermally induced deformations in the rock 
mass at one location will have limited impact on the behavior of the other location; and (2) the 
maximum difference in temperatures between two adjacent locations is only about 7C, as shown 
in Figure G-1, and this small thermal gradient is unlikely to significantly affect the behavior of 
each location. The thermal-mechanical response at locations between two adjacent stations 
(planes) is predicted using linear interpolation. 
Both in situ and seismic loading conditions are assumed to be the same for each of these 
cross-sectional planes with the thermal loading condition unique for each plane. The in situ 
loading condition is the geostatic state of stress existing prior to excavation and is related to the 
overburden thickness and horizontal stress ratios described in Section 2.3.3 of this technical basis 
document. Along an emplacement drift, the change of the overburden depth, and, therefore, the 
resulting change of the in situ stress state is assumed to be small. Use of the bounding in situ 
load in the analysis (BSC 2004a, Section 6.1.1.1) addresses the potential uncertainty associated 
with the variation in the in situ loading condition. At a drift or repository scale, seismically 
induced ground motions will have similar effect on the rock movement at different locations 
along an emplacement drift as long as the ground conditions at these locations are similar. 
Seven two-dimensional analyses were performed, each given an identical in situ or seismic 
condition but a unique temperature boundary condition that is associated with the specific 
location of the analysis. Since the seismic ground motions and temperature changes expected in 
the emplacement drifts are time-dependent, these two-dimensional analyses were assumed to be 
quasi-steady state and conducted using the FLAC computer code. The model configurations and 
boundary conditions used in these analyses are shown in Figure G-2. The rock mass mechanical 
properties considered are for lithophysal rock strength Categories 1 and 5, as listed in Evaluation 
of Emplacement Drift Stability for KTI Resolutions (BSC 2004a, Table 4-5a). The Category 1 
rock mass represents the weakest rock anticipated, whereas the Category 5 rock mass is for the 
strongest lithophysal rock. Detailed discussions on the rock mass strength categories and 
associated mechanical properties are provided in Sections 3.2, 4.2, and 5.3 and Appendix A of 
this document. 
June 2004 G-6 No. 4: Mechanical Degradation Source: BSC 2004a, Figure 6.1-4. 
Figure G-2. Geometry and Boundary Conditions for a Typical Two-Dimensional Model 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 G-7 Revision 1 
The two-dimensional analyses are limited to predicting the lateral and vertical deformations in 
the invert (the longitudinal deformation was assumed to be zero). The upper bound for total 
invert longitudinal displacement can be determined using an approach where an emplacement 
drift is modeled as a column that can expand freely in the longitudinal direction when subject to 
elevated temperatures. The total thermal expansion or elongation can then be estimated from the 
following equation: 
(Eq. G-1) 
where 
.L = . TL 
.L = thermally induced elongation, in meters 
 = coefficient of thermal expansion, per degree Celsius 
.T = increase in temperature, in degrees Celsius 
L = length of drift, in meters. 
G.4.3 Predicted Rock Deformation in the Invert 
From the perspective of the steel invert structure and gantry rail design, only the effects of 
thermal and seismic loading conditions on the rock deformation in the invert are important. The 
deformation induced by in situ loading condition is inconsequential because the deformation of 
the entire drift wall, including the invert region, has already occurred and equilibrated prior to 
the installation of the invert structure and rails. Consequently, the excavation-induced rock 
deformation will not have any effect on the installed invert structure or rails or the emplaced 
waste packages. Any offset in the invert structure or the gantry rails caused by the rock 
movement in the invert will be due primarily to either thermally induced expansion or seismic 
ground motions. Since in situ load is considered in every model, its effect on the rock 
deformation is excluded by subtracting the rock displacements induced by the in situ load from 
the total displacements caused by the other combined loads. 
Seven plane-strain FLAC analyses were completed for each ground condition or rock type and 
initial stress state (K0 value1), representing seven locations with different temperature conditions 
along an emplacement drift. With two bounding rock mass categories and two initial stress 
conditions, a total of 28 FLAC runs were performed. Detailed results from these runs assessing 
the impact on rock deformations induced by elevated temperatures in the invert are provided in 
Evaluation of Emplacement Drift Stability for KTI Resolutions (BSC 2004a, Section 6.4.3). A 
summary of the results is presented below. 
Vertical Displacement.The vertical relative displacements at the invert are illustrated in 
Figure G-3. These values are estimated between any location along the drift and the entrance 
(Station 0) from the predicted total vertical displacements under combined in situ and thermal 
loads. These vertical relative displacements are relevant to the criterion specified for the gantry 
rail design (see Section G.4.1). The maximum vertical relative displacement over a distance of 
600 m (between Station 6 and Station 0) is predicted to be about 4 mm and not sensitive to either 
the rock mass categories or the initial stress conditions (K0 values). Also, the vertical relative 
displacements vary linearly along the drift, indicating that a linear interpolation can be used to 
1 K0 is the ratio of horizontal to vertical stress. The vertical stress is the major principal stress component, with K0 
0 values are conservatively values ranging from 0.36 to 0.62 (Section 2.3.3 of this technical basis document). The K 
assumed to vary from 0.3 to 1.0 in these analyses. 
G-8 June 2004 No. 4: Mechanical Degradation Revision 1 
estimate the relative displacement between any two locations. The vertical relative 
displacements in the invert over a distance of 6 m for various rock mass categories and initial 
stress conditions are then estimated and listed in Table G-2. 
Source: BSC 2004a, Figures 6.4-2b, 6.4-5b, 6.4-8b, and 6.4-11b. 
Figure G-3. Vertical Relative Displacements in Invert as a Function of Location along Emplacement Drift 
Table G-2. Summary of Predicted Rock Vertical Relative Displacements in Invert under Thermal and 
Seismic Loading Conditions 
Relative Displacements 
under Thermal Loadinga 
Relative Displacements 
under Seismic Loading Rock Type and Initial 
Stress Condition (mm) 
0.0 
0.0 
0.0 
0.0 
(mm) 
0.04 
0.04 
0.04 
0.04 
Category 1 and K0 = 0.3 
Category 1 and K0 = 1.0 
Category 5 and K0 = 0.3 
Category 5 and K0 = 1.0 
Source: BSC 2004a, Table 6.4-1. 
Total Relative 
Displacements 
(mm) 
0.04 
0.04 
0.04 
0.04 
June 2004 
NOTE: aThe numbers are the vertical relative displacements per 6 m of drift. 
The preexisting conditions, whether a drift is heated or not, have practically no effect on the rock 
vertical relative displacements during seismic ground motions (BSC 2004a, Figures 6.1-6, 
6.1-11, 6.1-16, and 6.1-21). The invert of the entire drift, even with a thermal gradient, will 
move in translational mode like a rigid body during seismic ground motions. This observation 
indicates that the vertical relative displacements between any two locations along a drift during 
seismic ground motions are close to zero. The values are listed in Table G-2. 
G-9 No. 4: Mechanical Degradation Revision 1 
Lateral DisplacementDue to the thermal symmetry during heating, the maximum horizontal 
relative displacements within a drift occur between the walls at the springline. These relative 
displacements are defined as the drift horizontal closures (BSC 2004a, Section 6.1.4). A positive 
drift closure indicates a reduction in drift diameter. Use of the drift horizontal closures 
considered in the analyses to estimate the lateral displacements in the invert is very conservative. 
The predicted drift horizontal closures induced by heating are shown in Figure G-4. The 
maximum values are associated with Station 6 (600 m), and predicted to be about 2 mm. 
Horizontal displacements are only slightly sensitive to the rock mass strength categories. 
A summary of these values is provided in Table G-3. 
Source: BSC 2004a, Figures 6.4-3b, 6.4-6b, 6.4-9b, and 6.4-12b. 
Figure G-4. Time-Dependent Drift Horizontal Closures along Emplacement Drift Induced by Heating 
June 2004 G-10 No. 4: Mechanical Degradation Table G-3. Summary of Predicted Drift Total Horizontal Closures under Thermal and Seismic Loading 
Conditions 
Relative Displacements 
under Thermal Loadinga Rock Type and Initial 
Stress Condition 
Category 1 and K0 = 0.3 
Category 1 and K0 = 1.0 
Category 5 and K0 = 0.3 
Category 5 and K0 = 1.0 
(mm) 
1.68 (0.01) 
2.02 (0.02) 
1.17 (0.01) 
1.12 (0.01) 
Source: BSC 2004a, Table 6.4-2. 
NOTE: The numbers in this table are the horizontal closures and are conservatively used as the lateral 
displacements. 
a The numbers in parentheses are the lateral relative displacements per 6 m of drift. 
The horizontal relative closures between any station along the drift and Station 0 are shown in 
Figure G-5. They vary linearly along the drift, indicating that the linear interpolation can also be 
used for estimating the horizontal relative closure between any two locations. The difference in 
the horizontal closures between two locations separated by a distance of 6 m is then estimated, as 
listed in Table G-3. The difference in the horizontal closures between two locations separated by 
a distance of 6 m is practically negligible. 
Table G-3 also lists the horizontal closures induced by seismic ground motions. It can be seen 
that they are generally less than 2 mm, and sensitive to the rock mass modulus but not to the 
temperature conditions existing prior to shaking. 
Source: BSC 2004a, Figures 6.4-4b, 6.4-7b, 6.4-10b, and 6.4-13b. 
Figure G-5. Horizontal Relative Closures along Emplacement Drift Induced by Heating 
G-11 June 2004 No. 4: Mechanical Degradation 
Relative Displacements 
under Preclosure (10-4) 
Seismic Loading (mm) 
1.48 
1.64 
0.14 
0.14 
Revision 1 
Total Maximum Relative 
Displacementsa 
(mm) 
3.16 (0.01) 
3.66 (0.02) 
1.31 (0.01) 
1.26 (0.01) Revision 1 
Longitudinal DisplacementThe thermal expansion in the longitudinal direction is about 
4.0 mm over a distance of 12 m, which for a typical drift length of 600 m would be equal to 
0.2 m in total. This was estimated based on Equation G-1, using the maximum temperature 
difference of 42C over the 600-m-long drift (BSC 2004a, Table 4-1a) and a coefficient of 
thermal expansion of 8.02  10-6 per degree Celsius (BSC 2004a, Table 4-3). 
Comparison of the Predicted Rock Deformation and the Design TolerancesA summary of 
the predicted maximum relative displacements under thermal and seismic loading conditions is 
presented in Table G-4. It is evident that these relative displacements are well within the 
acceptable ranges of design tolerances for the steel invert structure and the gantry rails. 
Therefore, the rock movement in the invert during the repository preclosure period is not a 
concern from the perspectives of the steel invert structure performance and the retrieval 
operation. 
Table G-4. Summary of Predicted Maximum Rock Relative Displacements in Invert under Thermal and 
Seismic Loading Conditions 
Total Maximum Relative 
Displacements Components Acceptance 
Satisfied 
Satisfied 
Design Tolerance 
(mm) 
28.6 (6.4) 
6.4 
12.7 
(mm) 
3.66 (0.02) 
0.06 
4.0d 
G.4.4 Summary and Conclusions 
The rock movement in the invert of the emplacement drift is predicted to remain within the 
acceptable limits of design tolerances for the steel invert structure and the gantry rails. Both the 
invert structure performance and the retrieval operation during the preclosure period are expected 
to remain unaffected by the rock movement in the invert. 
G.5 REFERENCES 
G.5.1 Documents Cited 
BSC (Bechtel SAIC Company) 2003. Underground Layout Configuration. 800-P0C-MGR0- 
00100-000-00E. Las Vegas, Nevada: Bechtel SAIC Company. ACC: ENG.20031002.0007. 
BSC 2004a. Evaluation of Emplacement Drift Stability for KTI Resolutions. 800-KMC-SSE0- 
00200-000-00A. Las Vegas, NV: Bechtel SAIC Company. ACC: MOL.20040510.0199. 
BSC 2004b. Steel Invert Structure-Emplacement Drifts. 800-SSC-SSE0-00100-000-00A. 
Las Vegas, Nevada: Bechtel SAIC Company. ACC: ENG.20040119.0012. 
Laterala 
Verticalb 
Longitudinalc 
Source: BSC 2004a, Table 6.4-3. 
c 
NOTE: a
b 
The numbers in parentheses are the lateral relative displacements per 6 m of drift. 
The numbers are the vertical relative displacements per 6 m of drift. 
d 
The numbers are the longitudinal relative displacements per 12 m of drift. 
The number is estimated based on free expansion at 2 years per 12 m of drift. 
G-12 
Satisfied 
June 2004 No. 4: Mechanical Degradation Revision 1 
BSC 2004c. Ventilation Model and Analysis Report. ANL-EBS-MD-000030 REV 03 ICN 03, 
with errata. Las Vegas, Nevada: Bechtel SAIC Company. ACC: DOC.20031216.0002; 
DOC.20040202.0004; DOC.20040325.0003. 
Reamer, C.W. and Williams, D.R. 2001. Summary Highlights of NRC/DOE Technical 
Exchange and Management Meeting on Repository Design and Thermal-Mechanical Effects. 
Meeting held February 6-8, 2001, Las Vegas, Nevada. Washington, D.C.: U.S. Nuclear 
Regulatory Commission. ACC: MOL.20010307.0511 through MOL.20010307.0521. 
Williams, N.H. 2002. Thermal Inputs for Evaluations Supporting TSPA-LA. Interoffice 
memorandum from N.H. Williams (BSC) to Distribution, September 16, 2002, 0911024159, 
with enclosures. ACC: MOL.20021008.0141. 
G.5.2 Codes, Standards, and Regulations 
10 CFR Part 63. Energy: Disposal of High-Level Radioactive Wastes in a Geologic Repository 
at Yucca Mountain, Nevada. Readily available. 
CMAA 70-2000. Specifications for Top Running Bridge and Gantry Type Multiple Girder 
Electric Overhead Traveling Cranes. Charlotte, North Carolina: Crane Manufacturers 
Association of America. TIC: 249445. 
G.5.3 Data, Listed by Data Tracking Number 
MO0306MWDALAFV.000. ANSYS-La-Fine Ventilation. Submittal date: 06/23/2003. 
G-13 June 2004 No. 4: Mechanical Degradation INTENTIONALLY LEFT BLANK 
G-14 No. 4: Mechanical Degradation 
Revision 1 
June 2004 APPENDIX H 
CONTINUUM AND DISCONTINUUM ANALYSES OF 
GROUND SUPPORT SYSTEM PERFORMANCE 
(RESPONSE TO RDTME 3.11) 
No. 4: Mechanical Degradation 
Revision 1 
June 2004 Revision 1 
Note Regarding the Status of Supporting Technical Information 
This document was prepared using the most current information available at the time of its development. This 
Technical Basis Document and its appendices providing Key Technical Issue Agreement responses that were 
prepared using preliminary or draft information reflect the status of the Yucca Mountain Projects scientific 
and design bases at the time of submittal. In some cases this involved the use of draft Analysis and Model 
Reports (AMRs) and other draft references whose contents may change with time. Information that evolves 
through subsequent revisions of the AMRs and other references will be reflected in the License Application 
(LA) as the approved analyses of record at the time of LA submittal. Consequently, the Project will not 
routinely update either this Technical Basis Document or its Key Technical Issue Agreement appendices to 
reflect changes in the supporting references prior to submittal of the LA. 
June 2004 No. 4: Mechanical Degradation Revision 1 
APPENDIX H 
CONTINUUM AND DISCONTINUUM ANALYSES OF 
GROUND SUPPORT SYSTEM PERFORMANCE 
(RESPONSE TO RDTME 3.11) 
This appendix provides a response to Key Technical Issue (KTI) agreement Repository Design 
and Thermal-Mechanical Effects (RDTME) 3.11. The agreement relates to concerns regarding 
the continuum and discontinuum analyses of the ground support system performance that take 
into account long-term (preclosure period) degradation of the rock mass, joint-strength 
properties, and ground support components. 
H.1.1 RDTME 3.11 
Agreement RDTME 3.11 was reached during the U.S. Nuclear Regulatory Commission 
(NRC)/U.S. Department of Energy (DOE) Technical Exchange and Management Meeting on 
Repository and Design Thermal-Mechanical Effects held February 6 to 8, 2001, in Las Vegas, 
Nevada (Reamer and Williams 2001). There has been no submittal related to this KTI agreement 
to the NRC. 
The wording of this agreement is as follows: 
RDTME 3.11 
Provide continuum and discontinuum analyses of ground support system 
performance that take into account long-term degradation of rock mass and 
joint-strength properties. The DOE will justify the preclosure ground support 
system design (including the effects of long term degradation of rock mass and 
joint strength properties) in a revision to the Ground Control for Emplacement 
Drifts for SR, ANL-EBS-GE-000002 (or other document) supporting any 
potential license application. This is expected to be available to NRC in FY 2003. 
The agreement focuses on the time-dependent performance of the ground support system, taking 
into account the time-dependent degradation of rock mass and joint strength during preclosure. 
The role of ground support is to: (1) provide for personnel safety during construction and 
operation, (2) provide for operational clearances for equipment to facilitate repository operations 
during the preclosure period, (3) ensure that the regulatory requirement for waste retrievability is 
maintained, and (4) provide operational conditions facilitating installation of the drip shield at 
closure. 
The ground support is not introduced to protect waste packages from rockfall because the size 
and shape of the drift and potential rockfall particle sizes are small (see Appendix F for 
discussion of preclosure rockfall prediction). Ground support is considered to be not important 
to safety because the preclosure rockfall masses and velocities do not result in a credible scenario 
for breach of the waste package (BSC 2004a). Although ground support could remain functional 
for some additional period of time after closure, its functionality during the postclosure period is 
H.1 KEY TECHNICAL ISSUE AGREEMENT 
H-1 June 2004 No. 4: Mechanical Degradation Revision 1 
not considered in analyses of the postclosure performance of the repository. Analyses and the 
repository performance assessment during the postclosure period do not assign any credit to the 
installed ground control system, and the stability of the rock mass during that period is evaluated 
without considering ground control measures. 
The NRC concerns (NRC 2002, Section 2.1.7, p. 19) regarding the KTI agreement are as 
follows: 
 Time-dependent degradation of the repository host rock during preclosure, although not 
discussed in the DOE thermal-mechanical analyses for site recommendation (BSC 
2001), is potentially important because an operational life up to 100 years may be 
expected for the ground-support system (Wisenburg 2003). 
 A DOE expert panel on drift stability (Brekke et al. 1999) indicated that time-dependent 
degradation of the rock mass can be expected because of coupled thermal-hydrologicmechanical 
processes operating over a long period of time (100 years). Thermal, 
water-pressure, and rock-stress gradients occurring in the rock mass after nuclear waste 
emplacement would drive processes such as thermally induced fracture propagation, 
rock loosening, and cyclical evaporation and condensation of water. Such processes can 
be expected to cause degradation of the rock mass and ground support components. 
H.1.2 Related Key Technical Issue Agreements 
RDTME 3.04This KTI agreement involves providing a geotechnical parameters report that 
includes rock-mass property estimates of the lithophysal rocks of the Topopah Spring Tuff. 
RDTME 3.11 deals specifically with providing time-dependent characterization of the 
nonlithophysal and lithophysal rock-mass properties (Sections 3 and 4 of this technical basis 
document). Thus, the work performed to resolve RDTME 3.11 also forms a portion of the 
geotechnical parameters report that is used to resolve RDTME 3.04. 
RDTME 3.05This KTI agreement involves providing a methodology to account for the effect 
of lithophysae on ground control quality. The technical basis document discusses these load 
combinations in Section 4.2.3. 
RDTME 3.06This KTI agreement focuses on providing the design sensitivity and uncertainty 
analyses of the emplacement drift stability. The technical basis document discusses these load 
combinations in Section 5. 
TSPAI 2.02, Items 58 and 62TSPAI 2.02 Items 58 and 62 relate to the inclusion of rockfall 
and its potential mechanical impacts on engineered barriers (Item 58) and on the 
thermal-mechanical impacts of long-term rock-mass degradation on engineered barriers and 
potential hydrologic changes in the rock mass. RDTME 3.11 deals specifically with 
time-dependent rock-mass property estimates of the Topopah Spring Tuff nonlithophysal and 
lithophysal rocks and the development of rock-mass material and numerical models for 
representing rock damage, fracture development and propagation, and long-term degradation. 
The estimates feed long-term degradation and change of opening shape performance assessment 
H-2 June 2004 No. 4: Mechanical Degradation Revision 1 
studies of the engineered barriers. Section 4.2 of the technical basis document describes this 
integration in more detail. 
H.3 RESPONSE 
SummaryThe approach to analyzing the ground support system considering continuum and 
discontinuum analyses of nonlithophysal and lithophysal tuff units and long-term degradation are 
summarized in this appendix. 
A considerable effort was extended to evaluate appropriateness of utilizing continuous and 
discontinuous numerical models for representing lithophysal and nonlithophysal rock. In 
general, the three-dimensional discontinuum approach is considered appropriate for analyzing 
degradation and rockfall in the hard, jointed, and blocky nonlithophysal tuff units. 
A two-dimensional discontinuum approach was considered more applicable for analyzing 
degradation and rockfall in the lithophysal tuff units, where the tuff behavior is governed by the 
distribution and dimensions of lithophysae (voids) in the rock mass structure (see Sections 4.1 
and 4.2 of this technical basis document). Stability of the drifts in the preclosure time frame 
have been examined using parametric studies based on a continuum based modeling approach, in 
which rock mass properties have been derived from large-scale laboratory testing (lithophysal 
H.2 RELEVANCE TO REPOSITORY PERFORMANCE 
As planned, the repository excavations will be developed in two major types of lithostratigraphic 
volcanic tuff: the blocky nonlithophysal units (the middle nonlithophysal unit (Tptpmn) and the 
lower nonlithophysal unit (Tptpln)) and the lithophysal units (the upper lithophysal unit (Tptpul) 
and the lower lithophysal unit (Tptpll)). The lithophysal rock units constitute approximately 
85% of the repository emplacement area, whereas the nonlithophysal units (particularly the 
Tptpmn) make up the remaining 15% of emplacement area (see Section 2.2 of this technical 
basis document). Section 2.3 of the technical basis document describes the rock mass structural 
characteristics of the nonlithophysal and lithophysal units that control the rock mass mechanical 
properties. Section 3 of this technical basis document provides a description of the resulting rock 
mass and fracture properties of these units, while Section 4 describes the development of 
modeling capabilities to represent their mechanical response. Characterization of 
time-dependent effects for these two different rock types requires an application of numerical 
methods capable of accounting for the different structures that control mechanical response. 
Analyses involve the use of numerical models representing the rock strata as either an equivalent 
continuum or discontinuum media (see Section 4 of this technical basis document). 
A demonstration of the appropriateness of using these approaches for characterizing the tuff 
strata performance, considering in particular the long-term degradation of rock and ground 
support system, was described in Section 5 of this technical basis document. 
Estimates of rock mass mechanical properties (including time-dependent strength properties) are 
used as direct input to numerical models that predict the emplacement drift behavior when 
subjected to in situ, thermal, and seismic loading, as well as evaluating time-dependent changes 
in mechanical properties of rock (see Section 5 of this technical basis document). The results 
from these calculations are used to evaluate the impact that long-term change of rock strength 
may have on the stability and functionality of emplacement drifts during the preclosure period. 
H-3 June 2004 No. 4: Mechanical Degradation Revision 1 
rock) and empirically-estimated rock mass mechanical properties (nonlithophysal rock). These 
analyses are reviewed in Appendix B. 
The response to RDTME 3.11 is derived from the following analyses: 
 Three-dimensional 3DEC modeling of drift degradation and rockfall in nonlithophysal 
rocks associated with in situ, thermal, and seismic loading, as described in Sections 5.3.1 
and 5.3.2.1 of this technical basis document, was performed. The fracture geometries in 
the 3DEC model were derived from a stochastically-defined fracture network developed 
from the FracMan program, and were based on field mapping of fractures in the ESF 
and ECRB (Section 4.1.1). 
 Two-dimensional UDEC modeling of drift degradation and rockfall in lithophysal rocks 
associated with in situ, thermal, and seismic loading were completed, as described in 
Sections 5.3.1 and 5.3.2.2. 
 Two-dimensional UDEC modeling of long-term, time-dependent degradation of 
emplacement drifts in lithophysal rocks is based on the use of a stress corrosion model 
for representing time-dependent strength loss. The static fatigue laboratory data and 
time-related strength loss algorithm is described in Section 5.3.2.2.4. 
 Ground support techniques and materials have been chosen to ensure longevity through 
the preclosure time frame (BSC 2003a). 
The results of these analyses are described below: 
 A bounding analysis of potential effects of long-term degradation of fracture properties 
was conducted for nonlithophysal rock using the 3DEC program. It is assumed that the 
ultimate impact of long-term degradation of fracture properties would occur by shearing 
off surface asperities, leaving a smooth and cohesionless surface whose friction angle 
would equal the assumed residual friction angle of 30 (Section 5.3.2.1.8). These 
analyses indicate stable, unsupported emplacement drifts during the preclosure period. 
This analysis is very conservative in that shearing of fracture surfaces under the low 
preclosure stress conditions is not expected to result in time-dependent shearing of 
asperities. 
 The two-dimensional UDEC model of emplacement drifts in lithophysal rock was used 
to examine drift stability for bounding ranges of rock mass mechanical and thermal 
properties, as well as loading from in situ, thermal, and seismic stresses. Additionally, 
time-dependent strength degradation modeling was performed by reduction of the rock 
mass shear strength (cohesion) and tensile strength based on extrapolations from static 
fatigue testing. These analyses all indicate that unsupported drifts in the lithophysal 
rock are expected to be stable during the preclosure period. 
 Ground support methods, consisting of rock bolts and thin perforated steel sheeting over 
the entire exposed periphery of the emplacement drifts, were developed to allow air 
circulation to the rock surface, while minimizing any potential loosening or raveling of 
June 2004 H-4 No. 4: Mechanical Degradation Revision 1 
small rock fragments from the tunnel surface. Ground support materials are to be 
constructed from stainless steel to provide certainty of preclosure longevity. 
The above summary indicates a high degree of conservatism in the analysis of preclosure timedependency 
and in the ground support methods developed to ensure their function over the 
preclosure period. 
The information in this report is responsive to agreement RDTME 3.11 made between the DOE 
and NRC. This report contains the information that DOE considers necessary for NRC review 
for closure of this agreement. 
H.4 BASIS FOR THE RESPONSE 
This section describes the strategy, technical basis, and approach for resolving agreement 
RDTME 3.11. 
H.4.1 Overview of Resolution Strategy 
As a general approach for resolving the geomechanical issues related to the RDTME KTIs and 
addressing the associated NRC and DOE agreements, a resolution strategy was outlined in 
Resolution Strategy for Geomechanically-Related Repository Design and Thermal-Mechanical 
Effects (RDTME) (Board 2003), as discussed in Section 6.1. Documents that include analyses of 
aspects associated with issues required for the resolution of the RDTME 3.11 are listed below. 
H.4.3 Underground Environment 
Excavation of the repository drifts will result in concentration of the in situ stress around the 
openings (Section 1.2.2.1). The magnitude of in situ stresses at the repository horizon, 
considering the upper bound case of 400-m depth (Section 2.3.3) (BSC 2003b, Section 4.1) is 
equal to 9.2 MPa and is low in comparison to the uniaxial strength of tuff, whether 
nonlithophysal or lithophysal (BSC 2003c, Figure 8-22). Because the in situ stresses are 
relatively small in comparison to the rock mass strength, little, if any, yield of the rock mass is 
expected, even for the case of unsupported openings. Thus, the openings will undergo primarily 
elastic deformation, which equilibrates within a short distance (about two tunnel diameters) 
behind the advancing tunnel boring machine. Because permanent ground support is placed after 
the emplacement drift is completed and the tunnel boring machine withdrawn, it will be 
subjected only to deformation and loading that may occur from transient effects, such as thermal 
and seismic loading (Section 1.2.2.1 of this technical basis document). 
During the preclosure period, approximately 100 years from initial excavation of the 
emplacement drifts, forced ventilation will be used to remove approximately 90% of the heat 
generated by the waste packages (BSC 2004b). This heat removal will keep drift wall 
H.4.2 Basis for Resolution of RDTME 3.11 
The emphasis of this agreement is to determine the adequacy of the DOE approach to applying 
continuum and discontinuum analyses of ground support system performance that take into 
account long-term degradation of rock mass and joint-strength properties. 
H-5 June 2004 No. 4: Mechanical Degradation Revision 1 
temperatures below about 75C for 600-m drifts and 85C for 800-m drifts (Section 1.2.2.1) 
(BSC 2004b, Figure 6.6) and will result in small thermally related rock mass stress changes. 
These temperature ranges are also too low to cause significant impact on the ground support 
hardware. As discussed in Section 5 and Appendix B, little yielding is expected in either 
lithophysal or nonlithophysal rock due to preclosure loading. 
H.4.4 Ground Support System 
The role of ground support in nonlithophysal and lithophysal tuff is somewhat different. 
However, the element of providing confinement to the surface of tunnel walls and preventing 
loosening that would lead to detachment of blocks of nonlithophysal tuff or raveling of 
lithophysal tuff is common to both units. The ground support will be installed in two 
steps: (1) initial support, installed immediately behind the advancing tunnel boring machine for 
safety and personnel protection, and (2) the permanent support installed after the excavation is 
completed. 
During the preclosure period, the emplacement drifts will be supported by rock bolts and slotted 
stainless steel sheeting (BSC 2003a, Section 7.4), thus, minimizing, if not eliminating, 
mechanical degradation of the excavations. In particular, ground support consists of 3-m-long 
friction rock bolts placed on a square pattern on 1.25-m spacing and 3-mm-thick (approximately) 
perforated stainless steel sheets installed in a 270 arc around the drift periphery. Analyses, 
presented in Appendix B, indicate stability of the openings, even without ground support. Rock 
bolts and sheeting are expected to perform satisfactorily with acceptable factors of safety (BSC 
2003b, Section 7). 
As pointed out in Ground Control for Emplacement Drifts for LA (BSC 2003b), the ground 
support methods are selected based on the requirements of function, performance, and service 
life of the ground support system. The stability analyses described in this report employ the 
two-dimensional (FLAC) and three-dimensional (FLAC3D) programs, where rock mass is 
modeled as a continuum medium and is used to represent both nonlithophysal and lithophysal 
rock via input of different rock mass properties. Performance of the rock strata is assessed 
considering the opening response to thermal and seismic loads. The stability of emplacement 
drifts is assessed in two stages. Stage 1 provides assessment of unsupported drifts. Stage 2 
includes evaluation of emplacement drifts with ground support installed. 
Analysis of ground reinforcement performance in discontinuous strata is also reported in 
Evaluation of Emplacement Drift Stability for KTI Resolution (BSC 2004c, Section 6.3.5). The 
uncoupled ground support calculation (see Appendix F) represents the ground reinforcement 
installation identical to the one discussed above for fractured nonlithophysal rock. Rock strata 
represented within the three-dimensional discontinuum code, 3DEC, contain a network of joints 
subjected to in situ, thermal, and seismic loading. The performance of the ground support 
system subjected to dynamic loads is tested for the most severe case of preclosure in situ, 
thermal, and seismic loads. The modeling results indicate that the ground support performs well 
with an ample factor of safety for the extreme scenario assuming occurrence of unfavorable 
block orientations (BSC 2004c, Section 6.3.6, Appendix F). 
June 2004 H-6 No. 4: Mechanical Degradation Revision 1 
The performance of ground support during preclosure is also determined by the ability of its 
materials and components to remain functional without excessive deterioration. Longevity of 
Emplacement Drift Ground Support Materials for LA (BSC 2003a) reports on a range of factors 
potentially affecting the performance of ground support under preclosure environment in 
emplacement drifts. These factors include temperature and humidity within the drifts and rock 
bolt holes. Various forms of corrosion were examined, leading to the choice of stainless steel for 
ground support materials. Conclusions indicate that under these conditions the recommended 
materials and ground support hardware will perform satisfactorily for the preclosure period, 
taken to be nominally 100 years (BSC 2003a, Sections 7.3 and 7.4). 
H.4.5 Long-Term Degradation of Nonlithophysal Tuff 
Extensive work has been performed to enhance the understanding of the behavior and long-term 
performance of the hard, blocky, fractured nonlithophysal tuff. A three-dimensional 
discontinuum model (3DEC) for nonlithophysal rocks that accounts for joint structure and 
formation of the wedge-type blocks has been developed. Extensive discussion on this subject is 
presented in Drift Degradation Analysis (BSC 2004d) and Section 5.3.2.1 of this technical basis 
document. Appendices C and F provide a summary on the approach to the analysis of seismic, 
thermal, and time-dependent effects on rockfall and drift degradation. 
H.4.6 Long-Term Degradation of Fractures 
Fractures are an integral part of nonlithophysal tuff structure. Fracture surfaces have some 
degree of roughness due to undulations and asperities. When the fracture is subjected to shearing 
stress, as is typical around an excavation, the asperities will become points of stress 
concentration. These asperities may eventually yield due to stress corrosion cracking,1 with the 
resulting effect of shearing off the surface roughness. The ultimate (and extremely conservative) 
time-degradation effect of long-term loading of fracture surfaces would, therefore, be to shear all 
surface roughness, resulting in no cohesion or dilation of the surfaces, and with friction angle 
reduced to its residual value, taken to be approximately 30 (Section 5.3.2.1.8). Analyses of the 
impact of long-term joint deterioration, conducted using the 3DEC program, conservatively 
assume that the fractures have no cohesive strength, no dilation, and are at their residual angle of 
friction. The procedure used to determine the magnitude and range of stress conditions applied 
in the process of examining impact of long-term deterioration of joints is discussed in Drift 
Degradation Analysis (BSC 2004d, Section 6.3.1.5) and are summarized in Appendix C, 
Section C.4.5, Part 3d. Analyses performed using these parameters are presented in Evaluation 
of Emplacement Drift Stability for KTI Resolution (BSC 2004c, Section 6.3). 
H.4.7 Long-Term Degradation of Lithophysal Tuff 
The time-dependent change in strength of lithophysal rock is discussed at length in 
Section 5.3.2.2.4 of this technical basis document, in Drift Degradation Analysis (BSC 2004d, 
Appendix S), and in Subsurface Geotechnical Parameters Report (BSC 2003c, Section 9). Static 
fatigue laboratory experiments (essentially, creep testing of hard rock) have been conducted on 
nonlithophysal tuff to determine the time-to-failure of samples as a function of the ratio of 
1 A description of stress corrosion cracking as a mechanism for time-dependency in tuff is given in Section 5.3.2.2.4 
of this technical basis document. 
H-7 June 2004 No. 4: Mechanical Degradation Revision 1 
applied stress to the unconfined compressive strength. The PFC model was calibrated to 
reproduce the time-dependency of the nonlithophysal rock. The model was then used to examine 
the impact of lithophysal porosity on the time-to-failure. The result of these analyses is a 
relationship of time-dependency for lithophysal tuff as a function of lithophysal porosity. An 
engineering model was developed in which the strength of the lithophysal rock (in terms of 
cohesion) as a function of time for given levels of applied stress was developed. This abstraction 
provides a means of directly incorporating time-dependency into the UDEC discontinuum model 
that has been developed for examination of rockfall in lithophysal rock. This model has been 
used to examine the combined effect of in situ, thermal, and seismic loading, in addition to 
time-dependent strength degradation, on drift stability and collapse in lithophysal rocks. It was 
found that significant time-dependency during the postclosure period was restricted to the lowest 
strength category for the lithophysal rock, and little time-dependent strength loss is expected for 
the mean strength categories that represent the majority of the Tptpll. Based on these results, 
little time-dependent failure response is expected in the lithophysal rock over the preclosure time 
frame. Therefore, the existing ground support design for the preclosure period need not consider 
time-dependency in the lithophysal rock mass. 
H.4.8 Modeling of Time-Dependent Deterioration of Lithophysal and Nonlithophysal 
RockSummary 
Time-dependent analysis of emplacement drifts in lithophysal and nonlithophysal rock is 
reviewed in Drift Degradation Analysis (BSC 2004d) and Section 5 of this technical basis 
document. The approach taken to time-dependency is as follows: 
 Nonlithophysal RockTime-dependent effects are assumed to take the form of 
stress-corrosion failure of stress asperities on joint surfaces, with the ultimate result of 
joints achieving their residual strength condition with 0 cohesive strength, 0 dilation, and 
residual angle of friction (assumed to be 30). Even with these very conservative 
assumptions, 3DEC discontinuum analysis of stability of the emplacement drifts 
indicates stable conditions in the preclosure for unsupported tunnels (see 
Section 5.3.2.1.8). A minor amount of rockfall is predicted for unsupported tunnels in 
the event of seismic shaking. 
 Lithophysal RockAs described above, time-dependent strength properties of the 
lithophysal rock mass were developed based on laboratory static fatigue testing of 
nonlithophysal rock and extrapolation using the PFC2D program. Use of the resulting 
time-dependent strength properties within two-dimensional discontinuum modeling (see 
Section 5.3.2.2.4) shows that little time-dependent degradation of the emplacement drifts 
is predicted over the postclosure time frame, with little or no time-dependency expected 
during the preclosure time period of 100 years (Figures 5-41 to 5-47 of this technical 
basis document). 
Based on these discontinuum numerical analyses, preclosure time-dependency effects are 
estimated to be minimal, and the ground support, consisting of rock bolts and slotted steel 
surface sheeting, is considered to be sufficient to maintain drift stability. 
June 2004 H-8 No. 4: Mechanical Degradation Revision 1 
H.4.9 Conclusions 
The current approach to assessing ground support system performance employs routinely 
numerical codes in which rock strata are represented as continuum or discontinuum. These 
two-dimensional and three-dimensional codes are used to analyze ground support performance 
considering the time-dependent properties of rock and ground support under a range of loading 
conditions, which satisfies the requirements of RDTME 3.11. 
H.5 REFERENCES 
Board, M. 2003. Resolution Strategy for Geomechanically-Related Repository Design and 
Thermal-Mechanical Effects (RDTME). REV 00. Las Vegas, Nevada: Bechtel SAIC Company. 
ACC: MOL.20030708.0153. 
Brekke, T.L.; Cording, E.J.; Daemen, J.; Hart, R.D.; Hudson, J.A.; Kaiser, P.K.; and Pelizza, S. 
1999. Panel Report on the Drift Stability Workshop, Las Vegas, Nevada, December 9-11, 1998. 
Las Vegas, Nevada: Management and Technical Support Services. ACC: MOL.19990331.0102. 
BSC (Bechtel SAIC Company) 2001. Ground Control for Emplacement Drifts for SR. 
ANL-EBS-GE-000002 REV 00 ICN 01. Las Vegas, Nevada: Bechtel SAIC Company. 
ACC: MOL.20010627.0028. 
BSC 2003a. Longevity of Emplacement Drift Ground Support Materials for LA. 800-K0C- 
TEG0-01200-000-00A. Las Vegas, Nevada: Bechtel SAIC Company. 
ACC: ENG.20030922.0004. 
BSC 2003b. Ground Control for Emplacement Drifts for LA. 800-K0C-TEG0-00100-000-00A. 
Las Vegas, Nevada: Bechtel SAIC Company. ACC: ENG.20031016.0001. 
BSC 2003c. Subsurface Geotechnical Parameters Report. 800-K0C-WIS0-00400-000-00A. 
Las Vegas, Nevada: Bechtel SAIC Company. ACC: ENG.20040108.0001. 
BSC 2004a. Bounding Characteristics of Credible Rockfalls of Preclosure Period. 800-00C- 
MGR0-00200-000-00A. Las Vegas, Nevada: Bechtel SAIC Company. 
ACC: ENG.20040315.0009. 
BSC 2004b. Ventilation Model and Analysis Report. ANL-EBS-MD-000030 REV 03 ICN 03, 
with errata. Las Vegas, Nevada: Bechtel SAIC Company. ACC: DOC.20031216.0002; 
DOC.20040202.0004; DOC.20040325.0003. 
BSC 2004c. Evaluation of Emplacement Drift Stability for KTI Resolution. 800-KMC-SSE0- 
00200-000-00A. Las Vegas, Nevada. Bechtel SAIC Company. ACC: MOL.20040510.0199. 
BSC 2004d. Drift Degradation Analysis. ANL-EBS-MD-000027 REV 03A. Las Vegas, 
Nevada. Bechtel SAIC Company. ACC: MOL.20040513.0081. 
June 2004 H-9 No. 4: Mechanical Degradation Revision 1 
NRC (U.S. Nuclear Regulatory Commission) 2002. Integrated Issue Resolution Status Report. 
NUREG-1762. Washington, D.C.: U.S. Nuclear Regulatory Commission, Office of Nuclear 
Material Safety and Safeguards. TIC: 253064. 
Reamer, C.W. and Williams, D.R. 2001. Summary Highlights of NRC/DOE Technical 
Exchange and Management Meeting on Repository Design and Thermal-Mechanical Effects. 
Meeting held February 6-8, 2001, Las Vegas, Nevada. Washington, D.C.: U.S. Nuclear 
Regulatory Commission. ACC: MOL.20010307.0511 through MOL.20010307.0521. 
Wisenburg, M. 2003. Licensing Position on Duration of the Preclosure Period. E-mail from 
M. Wisenburg to T. Dunn and M. Board, July 1, 2003, with attachment. 
ACC: MOL.20030922.0234. 
June 2004 H-10 No. 4: Mechanical Degradation THE EFFECT OF SUSTAINED LOADING ON INTACT ROCK STRENGTH 
(RESPONSE TO RDTME 3.07) 
No. 4: Mechanical Degradation 
APPENDIX I 
Revision 1 
July 2004 Revision 1 
Note Regarding the Status of Supporting Technical Information 
This document was prepared using the most current information available at the time of its development. This 
Technical Basis Document and its appendices providing Key Technical Issue Agreement responses that were 
prepared using preliminary or draft information reflect the status of the Yucca Mountain Projects scientific 
and design bases at the time of submittal. In some cases this involved the use of draft Analysis and Model 
Reports (AMRs) and other draft references whose contents may change with time. Information that evolves 
through subsequent revisions of the AMRs and other references will be reflected in the License Application 
(LA) as the approved analyses of record at the time of LA submittal. Consequently, the Project will not 
routinely update either this Technical Basis Document or its Key Technical Issue Agreement appendices to 
reflect changes in the supporting references prior to submittal of the LA. 
July 2004 No. 4: Mechanical Degradation Revision 1 
APPENDIX I 
I.1.1 RDTME 3.07 
Agreement RDTME 3.07 was reached during the U.S. Nuclear Regulatory Commission 
(NRC)/U.S. Department of Energy (DOE) Technical Exchange and Management Meeting on 
Repository Design Thermal-Mechanical Effects held February 6 to 8, 2001, in Las Vegas, 
Nevada (Reamer and Williams 2001). This KTI agreement response is Appendix I of Technical 
Basis Document No. 4: Mechanical Degradation and Seismic Effects, Revision 1. The technical 
basis document, including Appendices A to H, was previously submitted to the NRC on June 17, 
2004. 
The wording of the agreement is as follows: 
RDTME 3.07 
The DOE should account for the effect of sustained loading on intact rock 
strength or provide justification for not accounting for it. The DOE will assess 
the effects of sustained loading on intact rock strength. The DOE will provide the 
results of this assessment in a design parameters analysis report (or other 
document), expected to be available to NRC in FY 2002. 
The agreement is concerned with the estimation of the impacts of time-dependent strength 
changes in repository host rocks due to sustained stresses from in situ and thermal loading and 
from environmental influences, such as wetting and drying. The impact of these strength 
changes on emplacement drift stability needs to be accounted for in the assessment of 
postclosure performance. The concern is further elaborated in Integrated Issue Resolution Status 
Report (NRC 2002, Section 2.1.7.3.3.2, pp. 2.1.7-19 and 2.1.7-20). The NRC concerns are 
paraphrased as follows: 
 Emplacement drift degradation over time can be expected due to coupled 
thermal-hydrologic-mechanical processes (Brekke et al. 1999) and needs to be accounted 
for in analysis of time-dependent degradation modes. Thermally induced fracturing, rock 
loosening, and cyclical evaporation and condensation of water may drive degradation 
mechanisms. 
THE EFFECT OF SUSTAINED LOADING ON INTACT ROCK STRENGTH 
(RESPONSE TO RDTME 3.07) 
This appendix provides a response to Key Technical Issue (KTI) agreement Repository Design 
and Thermal-Mechanical Effects (RDTME) 3.07. The agreement relates to the impact of 
potential long-term strength changes of the repository host horizon on stability of emplacement 
drifts. The drifts will be subjected to sustained in situ and thermal stress, and related time 
dependency of rock strength could result in drift degradation over time. 
July 2004 
I.1 KEY TECHNICAL ISSUE AGREEMENT 
I-1 No. 4: Mechanical Degradation Revision 1 
 Geochemical alteration of the rock mass resulting from elevated temperatures in the 
presence of saturated rock conditions could impact mineral chemical stability, driving 
rock weathering alteration processes that could lead to degradation. This effect could be 
most pronounced for fracture and lithophysae coatings, affecting the shear strength of 
joints or strength of the lithophysal rock mass. 
Appendix D to Technical Basis Document No. 4: Mechanical Degradation and Seismic Effects. 
I.1.2 Related Key Technical Issue Agreements 
RDTME 3.02This KTI agreement response was submitted to the NRC on June 17, 2004, as 
Appendix D to Technical Basis Document No. 4: Mechanical Degradation and Seismic Effects. 
This KTI agreement requires that drift degradation and ground support analyses be conducted for 
critical combinations of in situ, thermal, and seismic stresses. 
RDTME 3.04This KTI agreement response was submitted to the NRC on June 17, 2004, as 
Appendix E to Technical Basis Document No. 4: Mechanical Degradation and Seismic Effects. 
This KTI agreement involves providing a geotechnical parameters report that includes rock-mass 
property estimates of the subunits of the Topopah Spring Formation. RDTME 3.07 deals 
specifically with providing estimates of time-dependent strength changes of repository host 
horizon rocks, which is included in the geotechnical parameters report. 
RDTME 3.05This KTI agreement response was submitted to the NRC on June 17, 2004, as 
Appendix A to Technical Basis Document No. 4: Mechanical Degradation and Seismic Effects. 
This KTI agreement requires that the methodology for estimation of rock mass properties for 
lithophysal rock be described. 
RDTME 3.10This KTI agreement response was submitted to the NRC on June 17, 2004, as 
Appendix D to Technical Basis Document No. 4: Mechanical Degradation and Seismic Effects. 
This KTI agreement requires verification of the adequacy of the use of two-dimensional models 
for analysis of drift degradation. Sections 4.2.2.2 and 5.2.3.1.2 of this technical basis document 
discuss the use of two- and three-dimensional models for drift degradation analysis. 
RDTME 3.11This KTI agreement response was submitted to the NRC on June 17, 2004, as 
Appendix H to Technical Basis Document No. 4: Mechanical Degradation and Seismic Effects. 
This KTI agreement requires examination of the long-term degradation of the rock mass in 
lithophysal and nonlithophysal rocks, particularly as it affects ground support. Section 5.2.3.2.4 
of the technical basis document summarizes the specific approach to accounting for long-term 
strength degradation of material properties via use of static fatigue testing of tuffs. 
RDTME 3.12This KTI agreement response was submitted to the NRC on June 17, 2004, as 
Appendix F to Technical Basis Document No. 4: Mechanical Degradation and Seismic Effects. 
This KTI agreement requires a dynamic analysis of ground support systems during the 
preclosure phase using site-specific ground motions and discontinuum numerical modeling. The 
technical basis document centers on the postclosure dynamic analysis of lithophysal and 
nonlithophysal rocks using discontinuum methods. 
RDTME 3.13This KTI agreement response was submitted to the NRC on June 17, 2004, as 
I-2 July 2004 No. 4: Mechanical Degradation Revision 1 
This KTI agreement requires technical justification for boundary conditions for models used in 
drift degradation and ground support analyses. Section 5 of this technical basis document 
discusses static and dynamic mechanical and thermal boundary and initial conditions. 
TSPAI 2.02, Items 58 and 62TSPAI 2.02, Items 58 and 62 are related to the inclusion of 
rockfall and its potential mechanical impacts on engineered barriers and on the 
thermal-mechanical impacts of long-term rock-mass degradation on engineered barriers and 
potential hydrologic changes in the rock mass. RDTME 3.07 deals specifically with the 
estimation of time-dependent degradation of emplacement drifts subjected to in situ and thermal 
loading. The estimates made for long-term degradation and change of opening shape feed 
performance assessment studies of the engineered barriers. Section 1 of this technical basis 
document describes this integration in more detail. 
I.2 RELEVANCE TO REPOSITORY PERFORMANCE 
Agreement RDTME 3.07 deals with examination of time-dependent effects on emplacement drift 
degradation. Time-dependent degradation, as opposed to possible drift instability induced by a 
seismic event, is envisioned to be a relatively slow process involving sequential dislodging of 
blocks from the excavation periphery, which then fall into the drift and build up around the drip 
shield. This process is referred to as a raveling mode of rock failure. The root causes of 
time-related degradation are either changes in loading conditions around the drifts, geochemical 
alteration of the rock matrix or fracture fillings, or time-dependent changes in rock or fracture 
strength resulting from coupled thermal-hydraulic-mechanical effects. 
Time-dependent degradation of emplacement drifts may have an important impact on repository 
performance. Appendix H (the response to RDTME 3.11) addresses the issue of the effect of 
time-dependent degradation in the preclosure (i.e., 100 year) time frame and the impact this 
could have on ground support. It was concluded that time dependency in fracture or rock-mass 
strength is insignificant in the preclosure period, and that ground support is properly designed to 
account for potential rock-mass loosening and raveling mechanisms and uses materials that have 
sufficient longevity to function over this time period. In the postclosure period, ground support 
will eventually degrade, leaving unsupported drifts. The impact of drift degradation modes, 
which could lead to partial or complete drift collapse, could be important to performance in the 
following areas: 
1. Partial or complete drift collapse would result in rubble within the drift, which could 
create an insulating blanket surrounding the drip shield and waste package. This 
could, in turn, affect in-drift temperature and humidity conditions. The importance of 
this effect is dependent to some extent on the timing of the degradation (i.e., the 
number of years after waste emplacement). If it happens more than a few hundred 
years after closure, drift collapse has a relatively small effect on waste package peak 
temperature. Drift collapse is also accompanied by a decrease in drift relative 
humidity as a result of the enlargement of the emplacement drift, the overall 
temperature rise in the drift, and the increased temperature gradient from the waste 
package to the solid drift wall. 
I-3 July 2004 No. 4: Mechanical Degradation Revision 1 
2. Drift degradation could result in an enlargement of the emplacement drift over time. 
Change in drift footprint (i.e., plan view dimension), shape, and the presence of rubble 
could impact seepage to the drift in the postboiling time frame. 
3. Rubble accumulation in the drift will result in static loading to the drip shield. These 
potential loads are required as input to the evaluation of drip shield performance. 
This appendix summarizes the results of analyses of time-related drift degradation in the 
nonlithophysal and lithophysal rocks that comprise the repository host horizon. A detailed 
description of the analyses can be found in Section 5.3 of this technical basis document and in 
Drift Degradation Analysis (BSC 2004a, Appendix S). Drift degradation resulting from seismic 
events is also described in Section 5.3 of this technical basis document. 
I.3 RESPONSE 
Long-term response of excavations depends on a number of factors, including: 
1. Degradation due to time-dependent fracture development in the rock matrix or along 
joint (or existing fracture) surfaces in the presence of water vapor and driven by 
mechanical and thermal stresses. Large-span excavations initially created with low 
factor of safety against collapse and using poorly controlled excavation methods 
increase degradation potential. 
2. Degradation due to time-dependent alteration of rock matrix or joint-filling materials 
due to rock mass thermal and moisture conditions. Water sensitive minerals, such as 
clays, increase degradation potential. 
General understanding of the time dependency of tunnels in hard rocks is subject to a high 
degree of uncertainty because most engineered excavations are designed for a relatively short 
life-span or are typically heavily supported to provide safe access for personnel. Although it is 
well known that many unsupported mining excavations deteriorate within years of initial 
excavation, there are also many examples of man-made excavations and natural underground 
openings hundreds to millions of years in age that have suffered little deterioration. A detailed 
database of the long-term stability of unsupported man-made excavations and natural 
underground openings is not readily available. Therefore, it is difficult to develop general rules 
for time-dependent degradation based purely on empirical observations. The approach used here 
for estimation of time-dependent degradation effects is based on development of a mechanistic 
understanding of the time dependency of strength change of Yucca Mountain site-specific rocks. 
The approach is based on development of a database of laboratory test results defining timerelated 
strength changes of tuff, followed by application of these data within numerical models 
of drift stability. This analysis accounts for the site-specific, time-evolving stress conditions 
over the postclosure time period, as well as environmental effects on rock strength change. 
Uncertainty in the rock mass properties is accounted for through use of appropriate ranges of 
estimated rock mass strength and time-dependent variability. 
Due to the general lack of water-sensitive in-filling materials on natural fractures or within the 
rock matrix, the impact of environmental effects on rock strength as a function of time is 
expected to be inconsequential. The primary driving mechanism for time dependency is 
July 2004 I-4 No. 4: Mechanical Degradation Revision 1 
expected to be strength loss of the matrix material due to subcritical microcrack growth resulting 
from a stress corrosion mechanism (e.g., Potyondy and Cundall 2001). This mechanism results 
in time-dependent crack growth due to stress, temperature, and the partial pressure of water at 
the tips of microcracks. These microcracks may begin as flaws in the rock or grain-boundary 
contacts. A crack grows either through grains or along grain boundaries due to the hydration and 
breaking of siliconoxygen bonds at the tip of the crack. The rate at which the crack grows is 
controlled by the diffusion of water to the crack tip, which is, in turn, dependent on the crack 
geometry, which is stress-dependent. The overall impact of these factors is a logarithmic form of 
crack growth as a function of time and stress. This mechanism is commonly termed stress 
corrosion cracking. Experimental data on single crystals of quartz, as well as in rocks, have 
validated these mechanisms (e.g., Martin 1972; Kranz 1979). Laboratory testing has been 
performed to define the rate at which this stress corrosion mechanism occurs in Yucca Mountain 
tuff under saturated and heated conditions (Martin et al. 1997; DTN: SN0406L0212303.002). 
The results of this testing have been used to develop a time-dependent strength model that has 
been implemented within a drift-scale stability model. This model was then used to examine 
time-dependent drift degradation for the postclosure in situ and thermal stress history. 
The analyses described in the technical basis document (Section 5.3.2.2.4), and summarized 
here, indicate that in situ and thermal stressing result in only minor damage and degradation 
around the tunnels and that time-dependency of the strength change in tuff provides a more 
important mechanism, particularly in high porosity lithophysal rock. A range in drift 
degradation conditions is expected to occur in lithophysal rocks as a result of variability of rock 
porosity, which, in turn, affects rock-mass strength. In general, the analyses indicated that: 
 A small percentage of emplacement driftsthose located in the highest porosity/lowest 
quality areas of the Tptpllare expected to show significant degradation over time. 
These constitute less than about 10% of the emplacement drift length based on the 
mapped lithophysal cavity data described in Section 4.2.2.1. 
 In the average quality rock conditions, little drift deterioration is expected from 
time-dependent effects. 
 Analysis of drip shield stability, based on a conservative assumption of rubble loading 
from complete collapse of the emplacement drifts, indicates that the drip shield is 
structurally stable. 
 Analysis of the effects of drift degradation on in-drift temperature and relative humidity 
conditions shows that little impact on maximum waste package temperatures are 
expected for partial or complete collapse, if it occurs more than a few hundred years after 
repository closure. Significant decrease in drift humidity can occur as a result of collapse 
due to the enlargement of the emplacement drift, the overall temperature rise in the drift, 
and the increased temperature gradient from the waste package to the solid drift wall. 
 Collapse and accompanying increase in drift plan-view footprint can result in impact to 
seepage flux into the tunnel. 
July 2004 I-5 No. 4: Mechanical Degradation Revision 1 
Finally, evidence from observations of drift collapse in mining is used to verify that 
emplacement drifts in lithophysal and nonlithophysal rocks have dimensions that are well below 
those known to induce collapse. The field observations are obtained from experience in caving 
mines in which the span of excavations is increased to induce collapse. Additionally, the 
emplacement drift tunnels are developed with a stable circular shape and are driven using a 
nonrock-damaging mining technique using tunnel boring machines. Observations of tunnels 
in the existing ESF and ECRB Cross-Drift show stable conditions with no observed time-related 
damage in the repository host horizon units. Evidence of large, unsupported tunnels in similar 
rock at Hoover Dam is presented. These tunnels, which have been in continuous service for 
worker access for over 70 years, are stable and have shown little deterioration. 
The information in this report is responsive to agreement RDTME 3.07 made between DOE and 
NRC. This report contains the information that DOE considers necessary for NRC review for 
closure of this agreement. 
I.4 BASIS FOR THE RESPONSE 
I.4.1 Introduction 
As demonstrated by the ESF and ECRB Cross-Drift, excavations in the repository host horizon 
are currently stable under the existing conditions of in situ stresses and rock mass strength with 
only light1 ground support. However, it is expected that the ground support will completely lose 
its integrity during the 10,000-year regulatory period and, to some extent, drift degradation will 
occur due to increased stress from thermal loading and due to strength decay of the rock mass. 
Drift degradation is an important issue for repository design and performance because drifts must 
remain stable during the preclosure operational period and, eventually, rubble resulting from 
degradation could impact in-drift environmental conditions, seepage, and the integrity and 
performance of the drip shields. Estimation of the rate of drift degradation for the duration of the 
10,000-year regulatory period is, therefore, required. 
I.4.2 Empirical Observations of Degradation and Collapse of Excavations 
I.4.2.1 Unsupported Excavation Spans and Stand-Up Time 
There is currently no accepted methodology for estimating the time-dependent degradation 
behavior of tunnels in hard rocks. However, a number of empirical correlations have been 
developed for providing a means of estimating maximum stable spans of unsupported 
excavations and the length of time that an unsupported excavation may remain open and still 
provide a safe working environment2 (termed stand-up time) (Bieniawski 1989). Because these 
correlations are often used in estimating the time of instability of excavations and could be used 
for estimation of long-term degradation response, this appendix discusses their applicability to 
the postclosure stability of emplacement drifts at Yucca Mountain. 
1 Current ground support in 5 m and 7.62 m diameter tunnels in the repository host horizon includes friction rock 
bolts on nominal 1.25 m centers and light wire mesh on the crown of the excavations. In some locations, bolting 
and meshing of the sidewalls was also performed. 
2 A safe working environment here means that there is minimal danger of rockfall from loose pieces of surface rock 
created by blast damage or unstable blocks formed by fracture planes. 
I-6 July 2004 No. 4: Mechanical Degradation Revision 1 
Emplacement Drifts Act As Isolated ExcavationsEmplacement drifts for the repository, 
5.5 m in diameter, are separated from one another by a center-to-center distance of 81 m and are 
located at a nominal depth of 300 m below ground surface. The large ratio of drift diameter to 
spacing and diameter to depth means that the emplacement drifts are mechanically isolated from 
one another and from the effects of the ground surface. There is, eventually, thermal interaction 
between tunnels, but the local mechanical effects related to isolated tunnels is of greatest 
importance to stability. As discussed in Section I.4.2.2, the span of the emplacement drifts is 
well below the critical span for collapse of unsupported excavations as observed in mining 
practice. 
Empirical Correlations for Unsupported Spans and Standup TimeAs the span of an 
excavation is made larger, there eventually comes a point when instability and unsafe working 
conditions result. Empirical correlations (e.g., Hoek and Brown 1980, p. 287) of the maximum 
safe unsupported span and the time that this span may remain unsupported have been developed 
to assist tunneling engineers in planning excavations that provide a safe working environment for 
tunnel construction workers. These correlations, which are typically based on some measure of 
rock mass quality, are inherently conservative in nature due to their primary purpose, which is to 
ensure the safety of construction personnel from rockfall hazard (not necessarily from general 
collapse). It is important to note that these correlations are not based on case histories of actual 
collapse. 
Stand-up time curves (e.g., Hoek and Brown 1980, Figure 6) provide an estimate for potential of 
safe unsupported tunnels and are expressed in terms of span versus stand-up time for various 
rock-mass qualities (as determined by a geotechnical rock mass quality rating). For a given span 
of a tunnel, the stand-up time decreases as the quality of the rock mass becomes poorer. The 
stand-up time is a crude measure of unsupported span stability in that it does not account for 
opening shape, tunneling method, or, most importantly, the stress state around an excavation 
(e.g., depth of the excavation). For excavations whose span is well below the critical collapse 
span, the ratio of the stress concentration around the excavation to the short-term strength of the 
rock mass is one of the most important factors controlling stability and its evolution as a function 
of time. Stand-up time curves were developed based on empirical evidence of instability of 
excavations in particular stress conditions (e.g., deep South African mines), and their application 
to completely different conditions is questionable. 
Stand-up time is typically projected to be on the order of hours or days, even for good quality 
rock masses and may be only on the order of years. The long-term stability of existing lightlysupported 
or unsupported tunnels in mechanically-similar tuffs and rhyolite at the Yucca 
Mountain site and Hoover Dam (described in Section I.4.2.3) is evidence that stand-up time is a 
worker-safety-related indicator and is not relevant to predictions of degradation or true collapse 
time. These remarks indicate that construction-related span and stand-up time correlations have 
only limited application to postclosure predictions of drift degradation at the Yucca Mountain 
site. Therefore, as described in Section I.4.3.1, an analytical approach, based on the mechanics 
of time-dependent fracture growth in brittle rocks, is used for estimation of long-term 
degradation response rather than reliance on existing empirical estimations. 
July 2004 I-7 No. 4: Mechanical Degradation Revision 1 
I.4.2.2 Evidence of Maximum Unsupported Span from Mining Case Examples 
The mining industry, on the other hand, routinely drives unsupported excavations to large spans 
with the express purpose of inducing collapse for caving of ore bodies. Therefore, empirical 
evidence of maximum unsupported spans from mining case examples is more relevant to the 
estimation of the actual spans of excavations that induce degradation and ultimate collapse. The 
span of an excavation required to induce caving and collapse has been the topic of extensive 
study in the mining industry because certain types of mining methods (e.g., block and panel 
caving) are predicated to induce caving when a rock mass is undercut. Brown (2003) provides a 
summary of worldwide caving operations and the spans of excavations necessary to induce 
collapse as a function of the rock mass quality (i.e., strength). Figure I-1 provides a summary of 
worldwide experience in excavation span and collapse potential for cave mining. The plot 
provides an experience-based correlation between rock mass quality rating (in terms of rock 
mass rating) and the ratio of the plan view area of the excavation to its perimeter (A/P)3 for 
excavations that produce: (1) stable spans, (2) transitional excavations in which instability may 
begin but caving is not yet occurring, and (3) caving. As the rock mass quality rating decreases, 
the spans at which collapse occurs also decrease. To apply these data to emplacement tunnels at 
Yucca Mountain, an estimate of the rock mass quality is required. The case of lithophysal rock 
is presented here because it is the weakest of the host horizon units. As described in Subsurface 
Geotechnical Parameters Report (BSC 2003, Section 9.2.5), the estimated rock mass quality 
rating in terms of Geologic Strength Index or Rock Mass Rating (GSI or RMR) for the 
lithophysal rocks at Yucca Mountain is in the range of approximately 50 to 60. For this range of 
RMR, the hydraulic radius of a flat-roofed excavation required for caving is in the range of 
approximately 25 to 35 m. A transitional zone between stable excavations and caving for the 
lithophysal rock RMR range occurs for A/P ratios of 15 to 20 m. For the proposed 5.5-m 
diameter emplacement drifts, the hydraulic radius is simply equal to the radius (2.75 m). 
Therefore, field practice indicates that the stable to transitional cavingstable state is 
characterized by a hydraulic radius for lithophysal tuff in the range of 15 to 20 m. The boundary 
between transitional stable/caving to a caving state is characterized by an A/P ratio in the range 
of approximately 25 to 35 m. Note that the typical mining undercut has a flat roof, which 
promotes instability, while the shape of the emplacement drifts is circular, which promotes 
stability. The emplacement drift A/P ratio is, therefore, approximately 5 to 7 times below what 
would be considered to be a transitional state between stable and caving conditions. Based on 
field experience in caving, the A/P ratio of the emplacement drifts is well below the level of the 
A/P ratio estimated to be in a transitional state of collapse and even further below the level for 
assurance of complete collapse. Therefore, it is not surprising that the existing ESF and ECRB 
Cross-Drift (7.62 and 5 m diameter, respectively) are stable and show no evidence of degradation 
nearly a decade after excavation, even though the ground support is minimal. 
3 The plan view area of an excavation divided by its perimeter is termed hydraulic radius in the mining industry, and 
is typically used as a parameter for estimation of caving potential or roof instability because it takes into account 
the span and shape of excavation. This term is not to be confused with the hydraulic diameter, which typically is 
used in fluid mechanics calculations and is 4 times the cross-sectional area of a conduit divided by the perimeter. 
For a long tunnel of constant cross section, such as an emplacement drift, the hydraulic radius is simply equal to the 
tunnel radius. 
I-8 July 2004 No. 4: Mechanical Degradation Revision 1 
I.4.2.3 Observations of Yucca Mountain Tunnels and Excavations at Hoover Dam 
Unsupported or lightly supported tunnels (although perhaps not considered safe from a personnel 
standpoint) can stand in a stable condition for long time periods, particularly in good quality rock 
masses. For example, the ESF (7.62 m diameter) and ECRB Cross-Drift (5 m diameter) tunnels 
at the Yucca Mountain site were constructed in 1995 to 1997 and in 1998, respectively. 
Although the ESF main loop is located largely in the Tptpmn, the ECRB Cross-Drift cuts 
through and exposes all of the repository host horizon units. The tunnels are, in general, lightly 
supported with friction rock bolts and light wire mesh in the tunnel roof, with occasional friction 
bolts in the tunnel walls. There is no evidence of significant deterioration or degradation of the 
rock mass, and no significant episodes of rockfall have occurred. 
Source: Brown 2003. 
NOTE: Caving potential is expressed in terms of the modified rock mass rating and the ratio of the excavations plan 
view area to its perimeter. Modified rock mass rating is equivalent to rock mass rating in the case of Yucca 
Mountain excavations. Stable and caving regions are separated by a transition zone. 
Figure I-1. Excavation Dimensions Required for Caving Gained from Field Experience in Block and 
Panel Caving Mines 
July 2004 I-9 No. 4: Mechanical Degradation Revision 1 
An external review panel convened to examine Yucca Mountain drift stability (Brekke et al. 
1999) found that excavations of the north ramp through the upper lithophysal zone and the 
ECRB Cross-Drift through the lower lithophysal zone show that both zones have properties that 
are favorable for stability with minimum ground support. The panel also found that rock 
conditions in the lower lithophysal zone in the ECRB Cross-Drift were similar to those observed 
in the upper lithophysal zone in the north ramp; that continuous joints were not apparent, and 
there was almost no overbreak or loosening of the slabs or blocks; and that zones with more 
frequent short fractures were present and could be described as fracture zones, but even in these 
areas, overbreak and block loosening were largely absent. Tunnel deformation measurements 
have been regularly monitored since excavation, showing stable conditions. The conclusion is 
that the tunnels in both the lithophysal and nonlithophysal rock masses are obviously in a stable 
and self-supporting mode with no obvious deterioration in 5 to 8 years. Additionally, the Drift 
Scale Test involved heating a representative repository-scale tunnel in the nonlithophysal rock 
mass, first, to postclosure temperature distributions, followed by a thermal overdrive experiment 
to test rock strength limits. The experiment, now well into its cooldown phase, showed stable 
and predictable conditions at expected repository peak temperature conditions. Overdrive to 
approximately 200C drift-wall temperatures showed predictable, minor spalling of a small 
portion of the center of the crown of the drift (BSC 2004a, Section 7.7.7.5.3). Cooldown has 
showed no observable loosening or instability of the tunnel. This experiment confirmed 
modeling estimates of stable drift conditions for expected repository temperature and combined 
in situ and thermal stress conditions. 
The Hoover Dam, with abutments excavated in Tertiary pyroclastic flows, was completed by the 
Bureau of Reclamation in 1936. Along with the construction of the dam itself are a series of 
tunnels and adits (Figure I-2) that were excavated to accommodate the various penstocks, valves, 
access ways, spillways, and river bypasses. With the exception of the visitor center elevator 
shaft (completed in the 1990s), all the excavations were completed with simple drill and blast 
methods (simple meaning here that no smooth-wall blasting techniques were used). Some of 
the larger openings, generally those more than 6 m (20 ft) high, were excavated using headingand-
bench methods (i.e., the tunnel is mined in two passes with the upper portion excavated first, 
followed by bench blasting of the lower portion) to develop the full size of the openings. Many 
of the tunnels and adits were excavated to greater than 12 m (40 ft) in diameter. While some of 
the penstock and spillway tunnels were lined with concrete, many of the adits4 and access ways 
remain unlined. 
The rock at the site is the tuff of Hoover Dam, a fairly localized unit composed of andesitic to 
dacitic pyroclastic flows and breccias. At the dam, the volcanics are slightly to densely welded, 
and slightly weathered to unweathered. At the penstock adits, the rock is moderately hard to 
hard and contains abundant lithic fragments and occasional corroded pumice fragments. The 
rock is slightly to moderately fractured, with most fractures devoid of fracture filling. Many of 
the discontinuities exposed in the adits are frequently shears and small faults displaying distinct 
4 An adit is a term for a tunnel with one end that daylights. 
July 2004 I-10 No. 4: Mechanical Degradation Revision 1 
slickensides5. Where the adits extend below the phreatic surface, occasional calcium carbonate 
precipitate is present adjacent to active or old seeps. 
The adits were excavated downstream of the power plant to allow insertion of large, steel 
penstock sections into tunnels that paralleled the canyon walls. The adits are still in use, housing 
the sewage treatment system and other support utilities necessary to the function of the dam. The 
adits are approximately 40 ft (12 m) high by 35 ft (10.7 m) wide (Figure I-3) becoming slightly 
taller with depth. After the drill and blast excavation, the adits were left unlined and 
unsupported, and continue to be unsupported to the present time. Rock fall in the adits has been 
limited to very occasional centimeter-size fragments, even without ground support. 
5 A slickenside is a striated, polished fracture surface resulting from shear displacement. Typically, the slickensided 
surface has low friction angle. 
July 2004 I-11 No. 4: Mechanical Degradation NOTE: Construction adit, shown at right, is approximately 40 ft high. 
Figure I-2. Excavation for Nevada Canyon Wall Outlet Works (Top) Showing Construction Adit in 1933 
and (Bottom) in 2004 
I-12 No. 4: Mechanical Degradation 
Revision 1 
July 2004 Revision 1 
NOTE: Irregular tunnel walls resulting from drill-and-blast excavations (top) and close-up view of the tunnel crown 
showing evidence of drill half-barrels (bottom). 
Figure I-3. 2004 Photographs of Unsupported Construction Adit at Hoover Dam, Excavated in 1931 
July 2004 I-13 No. 4: Mechanical Degradation Revision 1 
Additionally, there are numerous access ways throughout the lower canyon walls in and around 
the Hoover Dam power plant. These smaller tunnels, 6 to 15 ft (1.8 to 4.6 m) in diameter, allow 
access by personnel and tourists to various areas of the power plant and penstocks. Few of these 
tunnels are supported either by rock bolts or mesh. No steel supports are visible in the Hoover 
excavations. As with the adits, rock in the access ways and tunnels has been limited to rare 
centimeter-size fragments that are removed by the janitorial staff. 
I.4.2.4 Summary of Empirical Observations 
The safety-related empirical correlations for maximum unsupported span and stand-up time of 
excavations developed for the tunneling industry are not relevant for prediction of long-term 
response of repository excavations. The span of the emplacement drifts is significantly less than 
that required to initiate collapse, as indicated by practice in the mining industry. The excavations 
will be developed using nonblasting methods (i.e., using a tunnel boring machine) and with a 
circular shape that minimizes overbreak and promotes stability. Observations of existing tunnels 
in the repository host horizon at the Yucca Mountain site as well as in similar rock at Hoover 
Dam show stable conditions with minimal or no ground support. 
I.4.3 Potential Mechanisms of Time-Related Degradation of Emplacement Drifts 
During the postclosure period, the following mechanisms (excluding seismic events) of timedependent 
degradation could be important: 
1. Degradation due to time-dependent fracture development in the rock matrix or along 
joint surfaces in the presence of water vapor and driven by mechanical and thermal 
stresses. Large-span excavations, initially created with a low factor of safety against 
collapse using poorly controlled excavation methods, increase degradation potential. 
2. Degradation due to time-dependent alteration of rock matrix or joint filling materials 
due to rock mass thermal and moisture conditions. Water sensitive minerals, such as 
clays, increase degradation potential. 
Each of these issues is addressed below. 
I.4.3.1 Degradation Due to In Situ, Thermal, and Thermal-Hydraulic Stress Conditions 
and Time-Dependent Fracture Development 
I.4.3.1.1 Emplacement Drift Degradation from In Situ and Thermal Stress Change 
The nominal preclosure time period is assumed to be 100 years and will be followed by closure 
of the facility. Closure of the repository will involve emplacement of drip shield structures over 
the waste packages, cessation of forced ventilation of emplacement drifts, and backfilling of the 
ramps, shafts, access mains, and exhaust mains with crushed tuff. With the cessation of forced 
ventilation, the emplacement drifts and surrounding rock mass temperature will rise, as shown 
nominally in Figure I-4 (BSC 2004b). Here, the average drift wall and waste package 
temperature is shown as a function of time for a range of waste package types that spans the 
temperature range for locations within an emplacement drift in the lower lithophysal unit 
(designated Tptpll) emplacement area. The drift temperature rapidly rises, resulting in a peak 
July 2004 I-14 No. 4: Mechanical Degradation Revision 1 
21-PWR waste package temperature of approximately 170C approximately 30 years after 
cessation of forced ventilation. The temperature then undergoes a slow cooling over time, with 
the waste package surface remaining above the boiling point for approximately 1,000 years. In 
this nominal scenario, it is assumed that forced ventilation of the waste packages has occurred 
for 50 years. 
Source: BSC 2004b, Figure 6.3-14. 
NOTE: These waste packages bracket the range of temperature within a centrally located emplacement drift. 
Case 
This temperature history will result in transient, thermally induced stresses in the emplacement 
drift walls that are superimposed on the preexisting excavation-induced stresses to obtain the 
total wall stress. A complete discussion of the transient, thermally induced stresses around 
emplacement drifts in nonlithophysal and lithophysal rocks is given in Drift Degradation 
Analysis (BSC 2004a, Sections 6.3.1.3 and 6.4.2.3). The combined in situ and thermally induced 
stresses result in no predicted yield in the stronger, nonlithophysal rock mass (in agreement with 
observations from the Drift Scale Test) and result in only minor yield at the springline 
(sidewalls) of emplacement drifts in lithophysal rock. Figure I-5 shows that less than 0.5 m of 
the immediate springline area of the drift is expected to yield in any rock mechanical properties 
quality category6 of the lithophysal rock mass. The predicted sidewall shear failure would be 
observed as spalling of the sidewalls of the tunnels. In summary, parametric analyses of the 
6 A discussion of the rock mass mechanical property ranges for lithophysal rocks can be found in Section 4.2.2 of 
this technical basis document or in Appendix A. The lithophysal rock mechanical properties have been subdivided 
into a series of five strength categories that span the range of expected in situ porosity conditions, with category 1 
being the lowest quality, highest porosity case, and category 5 being the highest quality, lowest porosity condition. 
Figure I-4. Drift-Wall Temperature (a) and Waste Package Temperature (b) as a Function of Time for 
Three Waste Package Types, Emplacement Drift in the Tptpll, Mean Seepage Infiltration 
July 2004 I-15 No. 4: Mechanical Degradation Revision 1 
effect of combined in situ and thermally induced stresses (in the absence of time-dependent 
effects on material properties) are predicted to result in only minor drift degradation. 
Source: BSC 2004a, Figure 6-142. 
NOTE: 80 years (i.e., 30 years after cessation of nominal 50 years of forced ventilation) is the time of maximum drift 
wall temperature and thermally induced stress. Yield at springline can be seen (c) as that region that has 
failed and unloaded (destressed) due to shear fracture development. 
Figure I-5. Thermally Induced Rockfall and Stresses after 80 and 10,000 Years of Heating in Rock 
Mass for Poorest Quality Tptpll Strength 
July 2004 I-16 No. 4: Mechanical Degradation Revision 1 
I.4.3.1.2 Impact of Thermal-Hydrologic Effects on Emplacement Drift Degradation 
Estimates of the coupled thermal-hydrologic behavior of the rock mass in the near vicinity of the 
drifts is examined in Multiscale Thermohydrologic Model (BSC 2004b). Numerical parameter 
analyses of the impact of waste heating on fluid pressures on fractures and infiltration into the 
emplacement drifts were conducted using the multiscale thermal-hydrologic model. For the 
range of hydrologic properties of the four host-rock units, the fracture permeability is sufficiently 
large and fractures are sufficiently well connected to allow gravity-driven drainage of water to 
occur in an unrestricted fashion. Thus, percolation flux, not fracture permeability, is the ratelimiting 
quantity governing the magnitude of gravity-driven liquid-phase flow to the boiling 
dryout zone. Additionally, the analyses show that potential pressure buildup along fractures due 
to vapor pressure from boiling of water is also negligible due to the free-draining nature of the 
fractured rock mass. From a mechanical stability standpoint, this means that fluid pressure on 
fractures during the drying and rewetting phases of the postclosure period has a negligible effect 
on drift stability. These predictions are borne out by the Drift Scale Test in which no mechanical 
drift instabilities occurred for thermal conditions representative of repository postclosure 
conditions. 
In summary, the conjecture in the Drift Stability Panel Report that postclosure temperature 
coupled with hydrologic effects in fractures will inevitably lead to collapse of some of the drifts 
(Brekke et al. 1999, p. 3-13) is not supported by either the results of the Drift Scale Test or the 
coupled thermal-hydrologic or coupled thermal-mechanical analyses. 
I.4.3.2 
I.4.3.2.1 
Degradation Due to Time-Dependent Fracture Development in the Rock Matrix 
in the Presence of Water Vapor and Driven by In Situ and Thermal Stresses 
As stated in Section I.3, time-dependent strength loss in the rock mass was examined using a 
mechanics-based approach in which laboratory testing is combined with numerical model 
extrapolations for prediction of the drift stability. The following section summarizes this work. 
A detailed description can be found in Drift Degradation Analysis (BSC 2004a, Appendix S). 
July 2004 
Background on Time Dependent Characteristics of Brittle Rocks 
One of the most striking characteristics of brittle rocks is that at temperatures well below the 
melting point, a rock subjected to a constant load exhibits a continuous increase in strain with 
time. This time-dependent deformation is termed creep. Studies on creep indicate that the 
observed strain depends upon the applied stress, the temperature, the partial pressure of water, 
and the confining pressure (e.g., Martin 1972). Moreover, the same mechanism responsible for 
the strain of brittle rocks in constant strain-rate tests is also operative in creep. That is, cracking, 
both along grain boundaries and through individual grains, produces the observed strain (e.g., 
Brace et al. 1966). Above approximately one half to two-thirds of the compressive strength, the 
dominant mode of deformation for brittle rocks is the opening and growth of cracks parallel to 
the major principal stress direction or axial orientation in unconfined compression. It is typically 
assumed that the strain rate of hard rocks in creep is related to the time-dependent growth of 
these cracks. 
I-17 No. 4: Mechanical Degradation Revision 1 
Verification of the relationship between time-dependent crack growth and creep strain rate in 
brittle rock is performed through laboratory testing. Experimental results indicate that stable 
time-dependent crack growth at a constant compressive load or at a constant stress intensity 
factor occurs in quartz and glass in the presence of water vapor. Moreover, the rate of crack 
growth depends on the applied stress, the temperature, and the partial pressure of water in the 
atmosphere surrounding the crack. The relative weakening of quartz or silicate glass, reflected 
by an increase in the rate of crack growth with an increase in any of the three variables, is 
consistent with the general theory of stress corrosion in silicates proposed by Charles (1959). He 
postulated that the velocity of a slowly propagating crack with a high tensile stress at the crack 
tip is proportional to the rate of the hydration reaction at the crack tip. The following equation 
(Martin 1972) quantifies the general relationship for environment-sensitive crack growth. 
+ (Eq. I-1) 0 RT 
 =   Pn exp 
. . F . V *  Vm . 
RT . . .
. .. 
where  is the rate of crack growth,  0 
is the initial flaw size, P is the partial pressure of water, 
. F is the activation energy for the process, T is temperature, R is the universal gas constant, V* 
is the activation volume,  is stress,  is the surface energy of the solid, Vm is the molar volume 
of the solid,  is the radius of curvature of the crack tip, and  and n are constants. 
If the partial pressure of water, the temperature, and the applied stress are constant, a constant 
crack propagation velocity will be observed. When any one of the thermodynamic variables is 
increased, the crack velocity increases. This expression has been verified with experimental 
studies (Wiederhorn 1968; Martin 1972; Scholz 1972). 
The validation of Equation I-1 is extremely important. First, it establishes a rate-dependent 
process for the propagation of cracks in quartz and silicate glass. If the same behavior is 
observed in rocks then it implies that time can be scaled from very short times to extremely long 
times in the absence of other competing mechanisms. Specifically, if moisture-assisted stable 
crack growth is the primary mechanism of creep in brittle rocks, measurements made at 
laboratory scales of up to 106 seconds can be extrapolated to much longer scales on the order of 
1011 to 1014 seconds. Presently there are no other independent data that suggest other competing 
mechanisms for time dependent deformation in brittle rocks at temperatures below 300 C. 
Based on these results, there is confidence that Equation I-1 accurately represents the behavior of 
the rate of crack growth at the crack tip for brittle silicate materials at temperatures below 300 C. 
Next, the behavior of brittle rocks can be examined during creep and compared to the 
observations of stable, time-dependent crack growth gained from tests on quartz and glass. A 
creep test is conducted by rapid application of uniaxial or triaxial load to a rock sample to a 
given differential stress, followed by holding the load constant while monitoring the longitudinal 
and lateral strains. Typically, creep is reported in terms of three distinct phases: (1) primary or 
transient creep, (2) secondary or steady-state creep, and (3) tertiary or accelerating creep 
(Figure I-6). 
I-18 July 2004 No. 4: Mechanical Degradation Revision 1 
Source: Martin et al. 1997, Figure 4. 
NOTE: Specimen failed at tertiary creep phase. 
Figure I-6. Example of Creep Strain Plotted as a Function of Time for a Static Fatigue Test Conducted 
on a Sample of Topopah Spring Tuff at a Constant Differential Stress of 132.8 MPa, a 
Confining Pressure of 5.0 MPa, a Pore Pressure of 1 MPa, and a Temperature of 150C. 
Transient creep has been reported for a variety of rock types over a wide range of temperatures 
and pressures. The strain in this region decelerates rapidly under static load and is often reported 
as proportional to the logarithm of time. Moreover, both the lateral and the longitudinal strains 
exhibit this logarithmic time dependence. 
At low stresses, transient creep may account for the observed strain. However, at high stresses, 
secondary creep is often observed. Generally, in secondary creep, often called steady-state 
creep, the strain is proportional to time. The total strain caused by both primary and secondary 
creep is often represented by an equation of the form 
(Eq. I-2)  = A + B log t + Ct 
where  is strain, t is time, and A, B, and C are constants. 
If secondary creep is allowed to continue, eventually the strain rate increases (tertiary creep) and 
the rock fails. All three stages of creep have been observed in granite, quartzite, and tuff (Martin 
1972; Martin et al. 1997). A typical creep curve for a specimen of welded tuff from the middle 
nonlithophysal unit of the Topopah Spring (Tptpmn) is shown in Figure I-6. All three stages of 
creep are clearly evident. 
July 2004 I-19 No. 4: Mechanical Degradation Revision 1 
Stable crack growth in quartz reported by Martin (1972) and Martin and Durham (1975), 
illustrated specific characteristics that are related to creep deformation. In these studies, the 
specimens were loaded to a fixed compressive stress and the growth of a crack parallel to the 
applied load was observed. Each specimen was tested in a controlled environment and the 
change in crack length was noted as a function of time. A typical data set obtained on a single 
specimen of quartz tested at a temperature of 241C and a partial pressure of water of 
4.5  10-2 kPa is shown in Figure I-7. The test specimen geometry is shown in the upper left 
portion of the graph. At a stress of 66 MPa, the change in crack length with time is very similar 
to that observed in the creep of brittle crystalline rocks. The crack exhibits an initial period of 
rapidly decelerating growth followed by a quasi-linear or secondary segment. After 6.3  104 
seconds, the stress was increased to 74 MPa. Immediately, the rate of crack growth increased. 
The same characteristics observed at the lower stress were exhibited for the 74 MPa segment. 
There was a strong transient followed by a secondary or quasi-linear crack growth segment. At 
approximately 8  104 seconds, the stress was increased to 83 MPa. The rate of crack growth 
increased dramatically; and the experiment was terminated when the crack length reached 
3.7 mm. These data are consistent with Equation I-1; that is, the rate of crack growth increased 
with increasing stress and nearly vanishes at low stresses. Additional experiments showed that 
increasing either the partial pressure of water surrounding the crack or the temperature also 
results in an increase in the rate of crack growth. 
Source: Martin et. al. 1997, Figure 1. 
NOTE: The experiment was conducted at 241C and a partial pressure of water of 4.05  10-2 kPa. 
Figure I-7. Crack Length as a Function of Time for an Axial Crack Growth Experiment in Single Crystal 
Quartz 
July 2004 I-20 No. 4: Mechanical Degradation Revision 1 
The above discussion points out that creep experiments with complex, silicate rocks display the 
same basic time-dependent response as demonstrated by crack-growth studies in single crystals 
of quartz and glass. From a practical standpoint, it is advantageous to define the ultimate timeto-
failure in terms of the stress, temperature and partial pressure of water, rather than in terms of 
crack growth. Time-to-failure is typically defined using the creep test to determine the static 
fatigue of a material. Static fatigue refers to the failure time of a rock or single crystal at 
constant stress, temperature, confining pressure, and partial pressure of water without regard to 
the strain history. Scholz (1972) conducted a series of static fatigue tests in compression on 
single crystal quartz. He observed that the mean time to failure, .t., depended on the partial 
pressure of water, P, stress, ; the activation energy, F; and temperature, T, according to: 
(Eq. I-3) 0 . RT 
where a and K are constants. 
t = t P .a exp. . 
.F . K' .. 
. 
The foregoing discussion demonstrates that: 
. The strength of brittle silicate rocks such as tuff is not a single-valued function of any 
parameter, but is a complex continuum that depends on the state of stress, the saturation 
(pore pressure), the temperature, and the time (including strain rate). 
. Studies of the basic growth of single fractures and the creep strain resulting from 
microcrack growth in complex silicate rocks demonstrate that the same basic stress 
corrosion mechanism is responsible for time-dependent crack growth and the ultimate 
time-to-failure of the material. 
. The stress corrosion mechanism gives rise to a logarithmic relationship of time-to-failure 
as a function of the state of stress, the temperature, and the pore pressure. 
. As a result of the basic understanding of the static fatigue mechanism in brittle rocks, it 
is possible to extrapolate long-term failure response from relatively short-term static 
fatigue experiments in the laboratory. 
Since the effects of time-dependent fracture development on weakening of tuff and its impact on 
drift degradation may be important in the postclosure repository environment, creep experiments 
on tuff samples have been conducted to determine its static fatigue response under appropriate 
environmental conditions. 
I.4.3.2.2 Static Fatigue Testing to Define Time-Dependent Behavior of Welded Tuff 
The typical way to define the time-dependent strength of rock is to establish the time required for 
failure of heated, saturated rock samples that are subjected to an applied constant axial stress. 
Creep test experiments are conducted to determine the static fatigue strength of the rock 
associated with tertiary creep rupture. These tests, typically conducted in uniaxial or triaxial 
compression, involve rapidly increasing the applied axial stress to a given percentage of the 
estimated compressive strength of the same size rock samples. The stress is then held constant 
July 2004 I-21 No. 4: Mechanical Degradation Revision 1 
until the sample spontaneously fails due to time-dependent rupture. A plot of the logarithm of 
the time-to-failure versus the ratio of the applied stress to the unconfined compressive strength is 
developed. The plot is typically linear, reflecting the basic mechanisms of stress corrosion as 
described above. Rock samples subjected to stress levels that are small in comparison to the 
compressive strength (i.e., below about 50% to 60%) result in excessively long times to failure 
and cannot be tested practically in the lab due to the long test duration. However, these loading 
conditions are not of interest for drift stability in the postclosure time frame of hundreds to 
thousands of years. The loading conditions of interest to time-dependent degradation at Yucca 
Mountain are those in which the applied stresses from in situ and thermal loading in the drift 
wall periphery are a high percentage of the rock strength (e.g., greater than approximately 60% 
to 70%). Here, the time to failure may result in significant degradation in hundreds to thousands 
of years. In this case, relatively short-term laboratory experiments (on the order of days to 
weeks) can supply time constants capable of describing the stress corrosion process. 
A series of triaxial static fatigue experiments were conducted on heated and saturated 
nonlithophysal cores from the Tptpmn in 1997 (Martin et al. 1997) and in 2004 (DTN: 
SN0406L0212303.002). Triaxial experiments on 50.8 mm diameter cores with a confining 
pressure of 5 MPa and pore water pressure of 4.5 MPa were conducted so that pore water of the 
saturated samples would remain in a liquid state as the temperatures were increased over boiling 
(125C and 150C). The resulting effective stress (the confining pressure minus the pore water 
pressure) was approximately 0.5 MPa, or essentially a state of uniaxial compression. This 
procedure was used to ensure a conservative state in which saturated samples were maintained at 
postclosure rock temperatures. Figure I-8 shows a typical specimen ready for testing. 
Figure I-8. Triaxial Static Fatigue Experimental Setup and Posttest Sample for Heated, Saturated, 
50.8 mm Diameter Samples of Tptpmn 
The results of the testing on nonlithophysal cores of Tptpmn, as well as those from similar 
testing of Lac du Bonnet granite performed for the Canadian high-level radioactive waste 
July 2004 I-22 No. 4: Mechanical Degradation Revision 1 
program (Schmidtke and Lajtai 1985; Lau et al. 2000) are given in Figure I-9. Granite results are 
included as a means of comparing the effects of rock type and for demonstrating the similarity in 
the general nature of the time-to-failure data for different rock types. Scatter in the data is due to 
sample heterogeneity, as well as the fact that the driving stress ratio (the horizontal axis) uses an 
estimated value for the unconfined compressive strength (adjusted for sample porosity) for 
normalizing the applied stress level. 
Since there is significant variability in the unconfined compressive strength of each sample, there 
will be a scatter in the resulting plot of time-to-failure versus driving stress ratio. As seen in 
Figure I-7, the welded tuff has a significantly slower time static fatigue failure than granite, as 
evidenced by the steeper slope of the linear fit to the data. This slower time-to-failure is 
presumably a result of the relatively homogeneous, fine-grained, high silica content nature of the 
tuff, as opposed to the heterogeneous nature of the grain structure of granite. 
Source: BSC 2004a, Figure S-1 (Lac du Bonnet data); Martin et al. 1997 (tuff 1997 data); DTN: 
SN0406L0212303.002 (tuff 2004 data). 
NOTE: Tests of Lac du Bonnet granite were conducted at 25C. The driving stress ratio is defined as the ratio of 
applied constant test stress to the estimated unconfined compressive strength. 1997 tuff tests were 
conducted at 150C, 2004 tuff tests were conducted at 125C. LdB = Lac du Bonnet. Linear fits to 1997 
Lac du Bonnet only and 1997 and 2004 tuff tests are shown. Samples that did not fail are also shown but 
not used in developing linear fits to data. 
Figure I-9. Static-Fatigue Data for Unconfined and Triaxial Compression of Heated, Saturated Welded 
Tuff and Lac du Bonnet Granite 
Linear fits to the unconfined compression data of Lac du Bonnet granite and to the 1997 only 
and to all (1997 and 2004) welded tuff data are given. The fits to both sets of welded tuff data 
are given because the analysis of drift degradation presented in the following section was 
conducted based on fits to only the 1997 data. After these analyses were completed, the 
July 2004 I-23 No. 4: Mechanical Degradation Revision 1 
additional 2004 data were collected. The linear fits to the data sets show the general consistency 
of the overall slope of the fits, although there is considerably more scatter in the 2004 test results. 
Due to data uncertainty, a lower bound for the slope of the time-to-failure curve based on the Lac 
du Bonnet data, was also used in numerical modeling estimates. The static fatigue testing was 
performed on saturated tuff cores at elevated temperature, ensuring that the impact of water on 
time-dependent yielding was accounted for in the estimation of time-dependent effects on drift 
stability. 
I.4.3.2.3 Development of a Mechanical Model for Simulating Time-Dependency in 
Nonlithophysal and Lithophysal Rock 
After an estimate of the relationship of stress level to time-to-failure for nonlithophysal tuff has 
been defined through testing, it is necessary to establish the impact of lithophysal porosity on the 
time dependence and to generalize the results into a time-dependent strength model that can be 
used to estimate drift stability. The methodology for development of a mechanical model for 
representing time-dependent degradation effects in welded tuff is described below. 
The particle flow code (PFC2D (BSC 2002a) and PFC3D (BSC 2002b)) discontinuum numerical 
modeling tools were used for understanding the impact of lithophysal porosity on time-to-failure 
as a function of applied stress. The particle flow code represents rock as a packed assemblage of 
small, rigid, grains that are bonded together at their contact points with shear and tensile strength. 
Porosity is developed naturally in the model by control of the shape and size of void space 
between chains of bonded grains. The macroscopic deformability of the assemblage is governed 
by the normal and shear stiffness at the contact points. Shear or tensile fracture between grains 
can form in a physically realistic manner as dictated by applied stresses, resulting in deformation 
and fracture of the rock mass. Quasi-static or dynamic stresses may be applied to the simulated 
rock for the solution of general boundary value problems. Details on the mechanics of the 
particle flow code program are provided in Itasca SoftwareCutting Edge Tools for 
Computational Mechanics (Itasca Consulting Group 2002). A modification to the basic particle 
flow code program was developed for simulation of time-dependent, stress corrosion cracking of 
rock. This model, termed the particle flow code stress corrosion model, was used for simulating 
time-dependent tunnel fracturing of Lac du Bonnet granite for the Canadian waste disposal 
research program. In this model, time-dependent intergranular bond fracture strength was 
developed based on the general concept of a stress corrosion mechanism. The long-term 
behavior is controlled by three particle flow code stress corrosion model parameters, 1, 2, and 
a. The terms 1 and 2 (rate constants) and a (microactivation stress) do not affect short-term 
material properties. These material parameters are derived from calibration against the time-tofailure 
data supplied by static fatigue testing (e.g., Figure I-9). The particle flow code stress 
corrosion model has been extensively documented and calibrated against static fatigue testing of 
Lac du Bonnet granite and validated against time-dependent tunnel breakout observed at the 
Underground Research Laboratory in Manitoba, Canada (Potyondy and Cundall 2001). Details 
of the stress corrosion model in particle flow code and its calibration can be found in Drift 
Degradation Analysis (BSC 2004a). 
Prior to representing time dependency, it was necessary to demonstrate that the particle flow 
code model can reproduce the basic, non-time-dependent mechanical behavior of nonlithophysal 
and lithophysal tuff. The calibration of the model against laboratory compression data from 
July 2004 I-24 No. 4: Mechanical Degradation Revision 1 
small cores of nonlithophysal tuff and from large cores of lithophysal tuff from the Tptpul and 
Tptpll was described in Section 4.2.2.3 of the technical basis document and in Appendix A. The 
same basic particle flow code model, incorporating a time-dependent particle bonding strength, 
was then calibrated to reproduce the time-to-failure dependent response of nonlithophysal tuff 
determined from the static fatigue testing. The calibration of the model is carried out by 
conducting simulated creep tests on nonlithophysal samples in exactly the same way they are 
performed in the laboratory. Figure I-10 presents a typical simulated creep test in which axial 
load is applied to the particle flow code sample and held constant at 80% of its unconfined 
compressive strength. Tensile fractures (cracks) develop spontaneously in the model as a 
function of time based on the time-dependent bond strength of the constituent grains. The plot 
shows the development of a network of tensile stress corrosion cracks that accumulate and 
propagate within the sample until a macroscopic shear failure mechanism develops with resulting 
brittle rupture during the tertiary creep stage. The simulated creep test shows all three stages of 
creep: transient, secondary and tertiary, and reproduces the typical response of creep experiments 
in tuff (e.g., Figure I-6). 
Source: BSC 2004a, Figure S-12. 
NOTE: Numerical simulation of creep test run by holding applied axial stress constant at 0.8 times the unconfined 
compressive strength. The damage, in terms of new crack growth, is displayed at various times along the 
creep curve. Brittle failure of the sample occurs when sufficient time-dependent crack growth results in 
failure mechanism. 
Figure I-10. Example of Simulated Creep Curve and Brittle Rupture Calibration for Nonlithophysal Tuff, 
(in This Case, Providing a Lower-Bound Estimate by Using Lac du Bonnet Granite Time-to- 
Failure Curve) Static-Fatigue Test at Driving-Stress Ratio (Ratio of Applied Stress to 
Unconfined Compression Strength) of 0.8 
A large number of particle flow code simulations of static fatigue tests of nonlithophysal rock 
was run at a wide range of driving stress ratios, and two particle flow code stress corrosion 
parameters were calibrated, so that the model was able to reproduce the basic time-to-failure fits 
July 2004 I-25 No. 4: Mechanical Degradation Revision 1 
shown in Figure I-9 for tuff and granite. The third stress corrosion parameter, the activation 
stress, was conservatively assumed to be 0. The consequence of this assumption is that the longterm 
strength of the particle flow code synthetic material is 0. It is well known that real rocks 
have long-term (true) strength that is on the order of 50% of the short-term strength. In other 
words, if the load is less than long-term strength, the rock will never fail, irrespective of duration 
of the load. The model was then used to investigate the impact of lithophysal porosity on the 
rate of time dependence. It is assumed that time-dependent behavior of the matrix is the same 
for both lithophysal and nonlithophysal rocks. A series of simulated creep experiments for 
lithophysal porosities of 11% and 20% were conducted, resulting in the generation of a set of 
time-to-failure versus driving stress ratio plots for various levels of lithophysal porosity 
(Figure I-11). 
I.4.3.2.4 Drift-Scale Model for Simulation of Time-Degradation of Emplacement Drifts 
The particle flow code model for time dependency of lithophysal rock, shown in Figure I-11, is 
computationally very large and is difficult to apply on the scale of a complete emplacement drift. 
To overcome these computational limitations, a drift-scale model using the UDEC discontinuum 
program was developed for investigation of non-time-dependent drift degradation analyses in 
lithophysal rocks (see Section 4.2.7 of the technical basis document). The UDEC model is 
described below. The same model is used for time-dependent drift-scale analyses with the 
exception that the strength properties of the model are adjusted as a function of time. The 
relationships for time-to-failure as a function of lithophysal porosity developed from the 
laboratory testing and particle flow code extrapolations are used to define time-dependent 
strength properties for the drift-scale UDEC model. The approach to definition of timedependent 
effects on strength in UDEC is simplistic in that it relates rock mass damage resulting 
from stress corrosion cracking directly to a loss of cohesion and tensile strength of the rock mass. 
The degree of strength loss was determined by (1) conducting a series of particle flow code 
numerical creep tests at different values of driving stress ratio and (2) interrupting the particle 
flow code creep test simulations at various times during a simulated test and conducting 
numerical compression and tensile strength tests on the damaged sample. For example, Figure I10 
shows four snapshots of the crack-damaged state of a simulated rock sample at various 
times along the creep curve. The strength properties of these damaged states were determined, 
and the resulting cohesion and tensile strength defined as a function of time for a given driving 
stress ratio. The strength loss was generalized into a damage coefficient that varies from 0 to 1 
(0 indicates no strength loss, while 1 is complete strength loss). The cohesion and tensile 
strength of the rock is multiplied by this coefficient to derive the strength properties of the rock 
mass as a function of time. Essentially, this approach relates the reduction in strength properties 
(shear and tensile strength) to the increase in fracture density or damage to the rock mass. 
Figure I-12a shows the form of the damage coefficient as a function of time for various driving 
stress ratios as derived from the particle flow code stress corrosion model for a nonlithophysal 
simulation. As seen in this plot, damage occurs in a brittle fashion with abrupt failure near the 
peak strength. The amount of damage accumulated prior to the abrupt failure (as shown by the 
damage coefficient) is less than 10% for high driving stress ratios (e.g., greater than 0.6), 
whereas damage accumulation is significantly larger for low driving stress ratios (e.g., less than 
0.6). A simplified representation of the damage coefficient evolution in terms of time is shown 
in Figure I-12b. 
July 2004 I-26 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figures S-5 and S-16. 
NOTE: The time dependency has approximately the same slope for all void porosities for the straight-line fit. 
Lithophysae are simulated as circles with a diameter of 90 mm. MS50 is the designation for the material 
properties of the matrix derived from nonlithophysal calibrations. 
Figure I-11. (a) Example Particle Flow Code Specimens with Void Porosities of 0.107 and 0.204 and, 
(b) Effect of Void Porosity on Time-to-Failure Response for Lithophysal Tuff Material (0% to 
20% Void Porosity, nv) 
July 2004 I-27 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figures S-21 and S-28. 
NOTE: Each curve has a vertical asymptote at a time-to-failure for a given driving-stress ratio, which is provided by 
the best fit to tuff given in Figure I-7. 
Figure I-12. Time Evolution of Damage Due to Strength Degradation Coefficient for Nonlithophysal Tuff 
Material during Static-Fatigue Tests at Driving-Stress Ratios Ranging from 0.2 to 0.8 (a) 
and Idealized Damage Coefficient as a Function of Time for a Range of Applied Stress 
Conditions (b) 
July 2004 I-28 No. 4: Mechanical Degradation Revision 1 
As discussed in Section 4.2.7 of the technical basis document, the UDEC drift-scale model is 
composed of many small elastic blocks that are bonded across incipient,7 ubiquitous fractures 
with shear and tensile strength components (Figure I-13). While the non-time-dependent UDEC 
drift-scale model assumes constant strength properties for the incipient fractures, the 
time-dependent damage model is transient in nature and assigns cohesion and tensile strength 
based on the damage coefficient as a function of time after excavation (i.e., Figure I-12b). 
are dependent on time based on the damage coefficients given in Figure I-12. 
NOTE: Model composed of irregular blocks joined across ubiquitous, incipient fractures. Blocks are bonded by 
shear and tensile strengths calibrated to rock mass strength estimates. These shear and tensile strengths 
July 2004 
Figure I-13. UDEC Model of Emplacement Drift in Lithophysal Rock Mass 
7 The term incipient here refers to the fact that the contact planes between blocks are initially bonded with the 
deformability and strength properties of the rock mass, and, thus, prior to failure, the rock mass acts as a 
continuum. If forces dictate, shear or tensile failure may occur along any of the incipient fractures, with the result 
of propagation along the incipient fracture network. The random nature of the incipient fracture network allows 
directional freedom of fracture growth. 
I-29 No. 4: Mechanical Degradation Revision 1 
Time-Dependent Drift Degradation Analyses I.4.3.2.5 
I.4.3.2.5.1 Effect of In Situ Stress Only in Lithophysal Rocks 
A series of parametric drift degradation simulations were conducted for the range of potential 
lithophysal rock mass strength categories in the Tptpll. The response to RDTME 3.05 
(Appendix A of this technical basis document) describes the subdivision of the expected range of 
mechanical properties of the lithophysal rock into five mechanical properties categories, each 
characterized by a Youngs Modulus and unconfined compressive strength. Each of these 
categories was related to the degree of lithophysal porosity, which, in turn, can be related to the 
level of abundance of that porosity in the Tptpll. Figure I-14 provides a histogram of 
approximate abundance of lithophysal porosity in the Tptpll obtained from geologic mapping in 
the ECRB Cross-Drift (BSC 2004a, Appendix O). An approximate relationship between 
lithophysal porosity and the rock mass mechanical properties category (1 is the weakest rock, 
highest porosity and 5 is the strongest rock, lowest porosity) is also given in this figure. 
Thermal-mechanical, time-dependent drift degradation analyses for the postclosure time frame 
were conducted for each of these strength categories as follows. Each of the incipient fractures 
in the drift-scale model (Figure I-13) is assigned stiffness, cohesion, tensile strength, and a 
friction angle corresponding to a particular rock mass mechanical properties category. 
Calibration of the model shows that these properties provide the proper overall rock mass 
deformability and strength properties corresponding to the particular mechanical properties 
category. When performing a time-dependent drift degradation analysis, the cohesion and tensile 
strength at each incipient fracture location is adjusted as a function of time, with yield assumed 
to occur in a brittle fashion when the time-to-failure is reached (Figure I-12). Thus, as the drift is 
excavated and as transient thermal stresses develop, time-dependent yield and fracture can occur 
around the excavation, resulting in redistribution of stress and possible propagation of drift 
breakout, collapse, and rockfall. 
The model was run by first excavating the emplacement drift under in situ stresses only, 
followed by application of the transient rock mass temperature conditions. The time-dependent 
fracture state and drift stability was examined at 1, 5, 10, 100, 1,000, and 10,000 years during the 
heating and cooling phases of the postclosure period. Additional analyses were conducted to 
examine the effect of seismic loading with the application of the 10-4 annual exceedance level 
ground motion time histories (approx. 0.5 m/s peak ground velocity; see Section 5.3.2.1.3 of this 
technical basis document). This ground motion was applied to the time-degraded model at the 
time of peak thermal stress (approximately 30 years after closure), at 1,000 years, and at 10,000 
years. The 10-4 event was chosen to determine if the added seismic stress and shaking would be 
sufficient to dislodge time-degraded, fractured, and loosened rock that may still be in place on 
tunnel walls. 
Figures I-15 to I-18 show the resulting time-dependent drift degradation estimates for 
mechanical property categories 1, 2, 3, and 5 for loading by in situ stresses only for 1 to 10,000 
years. As seen in these plots, the results for the lowest rock qualities (categories 1 and 2) show 
significant deterioration of the drifts would be expected soon after excavation as a result of in 
situ stress loading only. Observations in the ECRB Cross-Drift and the ESF, which have been 
excavated for 6 or more years, show no progressive raveling or overbreak (e.g., Brekke et al. 
I-30 July 2004 No. 4: Mechanical Degradation Revision 1 
1999, p. 2-5). This observation holds even in those areas of high lithophysal porosity found near 
the top of the Tptpll (see Section 2.3.2 of this technical basis document). Categories 3 and 5 
show little time-dependent effect from in situ stressing only. These results point out that the best 
estimate of tuff time dependence, coupled with the assumptions of brittle rock failure and 
constant, homogeneous properties within a given model cross section produces conservative 
damage estimates. Category 3 is considered to represent an average condition of lithophysal 
porosity for the Tptpll and shows little overbreak with time. Analyses presented in Drift 
Degradation Analysis (BSC 2004a, Appendix S) examine the impact of spatially variable rock 
properties within a given model section based on the mapped variability of lithophysal porosity 
in the ECRB Cross-Drift. The spatially variable model has a range of rock categories distributed 
throughout the cross section, and results of analyses show time-dependent drift degradation 
response similar to that shown for category 3 mechanical properties. An important conclusion 
from these initial in-situ-stress-only analyses is that best estimate time-dependent fracture growth 
within the tuff matrix is not expected to lead to collapse modes and significant drift degradation. 
I.4.3.2.5.2 Combined Thermal and Time-Dependent Effect in Lithophysal Units 
Throughout the regulatory period of 10,000 years, the emplacement drifts and surrounding rock 
mass will be subject to a heating cycle. Time-dependent strength degradation will happen 
concurrently with transient, thermally induced stress changes. Increased stresses around the 
excavation will accelerate the process of strength degradation. The results of numerical 
simulation of drift degradation as a result of these two processes are shown in Figures I-19 to I21. 
Time-dependent strength degradation is assessed using the tuff best-fit static-fatigue line. 
As expected, most rockfall occurs in category 2 rock mass, as shown in Figure I-19. Initially, 
most of the rockfall comes from the walls, which are loaded almost to a yielding state for this 
rock mass category under in situ stress conditions only. Strength degradation combined with a 
temperature increase, which at early times increases the hoop stress in the walls (not only in the 
roof), results in some rockfall from the wall at 5 and 10 years after emplacement of the waste. 
The large increase in the temperature and, consequently, in the stresses after the forced 
ventilation stops causes additional rockfall (at 80 years). At this stage, stress increase is 
predominantly in the roof. Therefore, some rockfall comes from the roof. It is counterintuitive 
that more rockfall is predicted in category 5 (Figure I-21) than in category 3 (Figure I-20). 
However, a large stiffness of category 5 lithophysal rock mass causes a large (larger than in 
category 3) increase in the hoop stress and yielding in the roof, even assuming the short-term 
yield strength of the rock mass. 
I-31 July 2004 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figures E-9 and E-10. 
NOTE: Lithophysal porosity data are from ECRB Cross-Drift station 14+44 to 23+26. Examples of approximate rock 
strength category levels taken from 1 by 3 m panel maps: (a) Category 3 with lithophysal porosity of 
approximately 19%; (b) Category 4 with lithophysal porosity of 13.3%; and (c) Category 5 with lithophysal 
porosity of 8.5%. 
Figure I-14. Distribution of Lithophysal Porosity and Estimated Rock Properties Categories for the Tptpll 
in the ECRB Cross-Drift 
July 2004 I-32 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure S-36. 
NOTE: Block colors correspond to their displacement (in meters) as shown in scale at right. Block color helps 
visualize those blocks that have moved. 
Figure I-15. Predicted Evolution of Damage Due to Strength Degradation for Category 1Tuff Best-Fit 
Static-Fatigue Curve, Applied In Situ Stress Only 
July 2004 I-33 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure S-37. 
NOTE: Block colors correspond to their displacement (in meters) as shown in scale at right. Block color helps 
visualize those blocks that have moved. 
Figure I-16. Predicted Evolution of Damage Due to Strength Degradation for Category 2Tuff Best-Fit 
Static-Fatigue Curve, Applied In Situ Stress Only 
July 2004 I-34 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure S-38. 
NOTE: Block colors correspond to their displacement (in meters) as shown in scale at right. Block color helps 
visualize those blocks that have moved. 
Figure I-17. Predicted Evolution of Damage Due to Strength Degradation for Category 3Tuff Best-Fit 
Static-Fatigue Curve, Applied In Situ Stress Only 
July 2004 I-35 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure S-39. 
NOTE: Block colors correspond to their displacement (in meters) as shown in scale at right. Block color helps 
visualize those blocks that have moved. 
Figure I-18. Predicted Evolution of Damage Due to Strength Degradation for Category 5Tuff Best-Fit 
Static-Fatigue Curve, Applied In Situ Stress Only 
July 2004 I-36 No. 4: Mechanical Degradation Source: BSC 2004a, Figure S-41. 
NOTE: Block colors correspond to their displacement (meters). 
Figure I-19. Evolution of Damage Due to Strength Degradation Resulting from In Situ and Thermal 
Load for Category 2 - Tuff Best-Fit Static-Fatigue Curve 
I-37 No. 4: Mechanical Degradation 
Revision 1 
July 2004 Source: BSC 2004a, Figure S-42. 
NOTE: Block colors correspond to their displacement (meters). 
Figure I-20. Evolution of Damage Due to Strength Degradation Resulting from In Situ and Thermal 
Load for Category 3 - Tuff Best-Fit Static-Fatigue Curve 
I-38 No. 4: Mechanical Degradation 
Revision 1 
July 2004 Source: BSC 2004a, Figure S-43. 
NOTE: Block colors correspond to their displacement (meters). 
Figure I-21. Evolution of Damage Due to Strength Degradation Resulting from In Situ and Thermal 
Load for Category 5 - Tuff Best-Fit Static-Fatigue Curve 
I-39 No. 4: Mechanical Degradation 
Revision 1 
July 2004 Revision 1 
It should be noted that static-fatigue curves are temperature dependent. This dependence is not 
explicitly included in the analysis. However, the tuff data are obtained from tests conducted at 
150C, which is higher than the maximum temperature of the rock mass anticipated throughout 
the repository for postclosure. Consequently, the results obtained in this analysis, although for 
isothermal static-fatigue curves, are conservative. 
I.4.3.2.5.3 Combined Seismic, Thermal, and Time-Dependent Effect in Lithophysal 
Units 
The 10-4 ground motion was applied to category 2 and 5 rock mechanical property cases at two 
time periods: (1) at the time of the peak thermal stress state (nominally at 80 years after 
emplacement, or 30 years after cessation of 50 years of forced ventilation) and (2) at 
10,000 years at the completion of the postclosure heating and cooling cycle. Figures I-22 and 
I-23 show the resulting predicted degraded drift states for these cases. Essentially, the 
application of the 10-4 ground motion (peak ground velocity of approximately 0.5 m/s) dislodges 
any fractured and loosened rock created by the thermal stress and time dependency. As seen in 
these figures, the impact of time dependency of strength properties is to widen the diameter of 
the emplacement drifts in the lower quality (category 2) rock due to progressive shear failure at 
the sidewalls. In the highest quality (category 5) rock, additional progressive yield of the roof is 
evident. Roof yield in higher quality rock is due to the higher modulus of the rock mass and, 
thus, higher roof-parallel thermally induced stresses during the heating cycle. This is similar to 
the roof crown spalling effect observed in the Drift Scale Test. 
July 2004 I-40 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figures S-44 and S-46. 
NOTE: Left figure shows state for in situ plus thermal stress; right figure shows state for in situ plus thermal plus 
seismic stress. Blocks colored by total displacement to allow visual recognition of blocks detached from 
surrounding intact mass. Block colors correspond to their displacement (meters) as shown in scale at right. 
Figure I-22. Effect of 10-4 Ground Motion after (Top) 80 Years and (Bottom) 10,000 Years of the 
Heating and Cooling Cycle in Category 2 Rock Mass 
July 2004 I-41 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figures S-45 and S-47. 
NOTE: Left figure shows state for in situ plus thermal stress; right figure shows state for in situ plus thermal plus 
seismic stress. Blocks colored by total displacement to allow visual recognition of blocks detached from 
surrounding intact mass. Block colors correspond to their displacement (meters) as shown in scale at right. 
Heating and Cooling Cycle in Category 5 Rock Mass 
I.4.3.2.5.4 
Figure I-23. Effect of 10-4 Ground Motion after (Top) 80 Years and (Bottom) 10,000 Years of the 
July 2004 
Time-Dependent Degradation of Nonlithophysal Rock 
The time-to-failure for intact nonlithophysal rock blocks shows significantly less time 
dependency than for the lithophysal rock described in the previous section (see Figure I-11). 
Therefore, insignificant time-related fracture growth is expected in the intact rock blocks. 
However, the potential exists for time-dependent yield of roughness (asperities) on fracture 
surfaces subjected to long-term shear stress. As described in Section 5.3.2.1 of this technical 
basis document, natural fractures in the nonlithophysal rock tend to control the drift stability 
response. There are four sets of fractures in the nonlithophysal rock: two subvertical cooling 
fracture sets, and one set of low dip-angle vapor phase partings and one set of random fractures 
I-42 No. 4: Mechanical Degradation Revision 1 
with dip averaging around 45 (see Section 4.1.2). The vertical cooling fractures are smooth and 
planar with low cohesion, friction, and dilation angle, while the vapor phase partings are rough 
and filled with vapor phase mineralization. The vertical cooling fractures consequently have 
higher cohesion, friction, and dilation angles (see Table 3-5). Stability of the emplacement drifts 
in the nonlithophysal rock has been examined for applied in situ, thermal, and seismic stressing 
(summarized in Section 5.3.2 of this technical basis document). The potential effect of 
time-dependent degradation of the fracture surfaces was examined in an impact study described 
in Section 5.3.2.1.8. A conservative approach to accounting for time-dependent effects was 
taken by assuming that: 
 The fractures were smooth, with roughness completely sheared off. Surface properties 
were set to a dilation angle of 0 and a residual friction angle of 30. The reduced friction 
angle is a typical value for a smooth joint reported by Goodman (1980, p. 158) and is 
consistent with the direct shear test results described in Section 3.2.4.1 of this technical 
basis document (DTN: GS030283114222.001). 
 The fractures have no cohesion. 
Stress analyses under in situ, thermal, and seismic stressing show that these highly conservative 
assumptions result in only a minor change in drift degradation. Increased thermal loading tends 
to increase stability of the fractured rock mass as thermally induced tangential stress around the 
excavation tends to increase the clamping forces on fractures. It is concluded that fracture 
geometry and the potential for formation of kinematically unstable blocks is the greatest 
influence on drift stability. 
I.4.3.2.6 Summary of Time-Dependent Strength Degradation Effects 
A mechanics-based model of time-dependent degradation of nonlithophysal and lithophysal tuff 
was developed using time-to-failure data generated from laboratory triaxial compression of 
heated, saturated samples of the Tptpmn. Using a similar approach to that developed for the 
Canadian high-level radioactive waste storage program, the laboratory data were used to 
calibrate a time-dependent fracture model for nonlithophysal rock within the particle flow code 
discontinuum numerical model. Lithophysal voids were added to the particle flow code model to 
estimate the impact of lithophysal porosity on the creep and time-to-failure response of tuff. 
This behavior was then embedded into the strength properties for the drift-scale model of 
emplacement drift stability in lithophysal rocks. In turn, this model was used for conducting 
parametric studies of drift stability subjected to in situ, thermal, and seismic stressing over the 
postclosure time period. The results indicate that some drift sidewall failure and breakout can be 
expected for the lowest mechanical property categories of lithophysal tuff (i.e., categories 1 and 
2) that represent approximately 10% of the rock mass. For the average and highest mechanical 
property categories (i.e., categories 3, 4, and 5), representing approximately 90% of the 
lithophysal rock mass, little time-dependent degradation is expected. Time-related degradation 
effects in fractured nonlithophysal rock were examined through three-dimensional discontinuum 
stress analyses in which the surface roughness and cohesion were assumed to be completely 
destroyed. Even with these highly conservative assumptions, it was found that fracture 
geometry, not surface properties, control the degree of degradation of the tunnels in 
nonlithophysal rock. 
I-43 July 2004 No. 4: Mechanical Degradation Revision 1 
I.4.3.2.7 Bounding Case of Drift Degradation Using Lac du Bonnet Granite Time-to- 
Failure Data 
The analyses presented in Section I.4.3.2.6 were performed assuming the slope of the best fit 
time-to-failure data for tuff. As shown in Figure I-9, the slope of the time-to-failure data fit 
obtained for Lac du Bonnet granite provides a lower bounding case to that for Topopah Spring 
Tuff. In other words, the granite exhibits significantly greater time dependency and strength loss 
than tuff. A series of time-dependent drift degradation analyses similar to those summarized in 
Figures I-15 to I-18 were performed for lithophysal rock categories 1, 2, and 5 but using the 
slope of the time-to-failure best fit for Lac du Bonnet granite in unconfined compression as 
shown in Figure I-9. These analyses are performed to investigate drift degradation extent as a 
function of time for a bounding case of time-dependent properties. The results, described in 
detail in Drift Degradation Analysis (BSC 2004a, Appendix S), show: 
 Using the granite time-dependency behavior for Tptpll strength category 1 (represents 
local conditions of less than about 5% of the Tptpll) shows significant drift breakouts at 
the sidewalls of the tunnel within 1 year of excavation and complete drift collapse in 
hundreds to thousands years. The conservatism in these properties is obvious because it 
would indicate extensive drift damage at early times, which is not observed in existing 
ECRB Cross-Drift and ESF tunnels. 
 Using the granite time-dependency behavior for Tptpll strength category 2 (represents 
local conditions, combined with category 1 that account for less than 10% of the Tptpll) 
results in breakouts at the sidewalls of the drift and rubble accumulation that does not 
cover the drip shield after thousands of years. Again, these analyses show significant 
early (1 to 5 years after excavation) breakouts on drift sidewalls not observed in existing 
tunnels at the site. 
 Using the granite time-dependency behavior for Tptpll strength category 5 results in 
minor breakouts at the springlines of the drift with little rubble accumulation. 
The above analyses use a highly conservative rate of time dependency evidenced by predictions 
of breakout at drift sidewalls and significant raveling at early times for the lower mechanical 
property categories 1 and 2. As discussed previously, there are no observations in either the 
7.62-m-diameter ESF or 5-m-diameter ECRB Cross-Drift that support such observations. 
However, even with these conservative calculations, significant breakout and raveling is 
restricted to those categories representing local conditions of high lithophysal porosity and less 
than about 10% of the Tptpll. Therefore, the overall conclusion that drift degradation resulting 
from time-dependent fracture and strength loss based on calculations using the best-fit, 
site-specific time-to-failure data for Topopah Spring Tuff is reasonable and represents typical 
conditions to be expected in the Tptpll. 
I.4.4 Degradation Due to Time-Dependent Alteration of Rock Matrix or Joint Filling 
Materials from Postclosure Thermal and Moisture Conditions 
One of the mechanical degradation issues raised in Integrated Issue Resolution Status Report 
(NRC 2002, Section 2.1.7.3.3.2, pp. 2.1.7-19 and 2.1.7-20) is the potential for geochemical 
July 2004 I-44 No. 4: Mechanical Degradation Revision 1 
alteration of the rock mass resulting from elevated temperatures in the presence of saturated rock 
conditions. The concern is that alteration of minerals in the postclosure environment, either 
along or filling rock fractures or within the rock matrix, could impact rock or fracture strength 
and lead to enhanced degradation processes. Time-related degradation of tunnels due to wetting 
and drying cycles is an important factor for drift stability in some rock types. These rock types 
are typically those in which highly-altered rock, or rock with significant swelling clay along 
fracture planes occurs. In the case of the host repository horizon at Yucca Mountain, these issues 
are not particularly important. Clay is not a common mineral in the crystallized rocks of the 
repository host horizon, nor are clay minerals a volumetrically significant fracture-coating 
material. Four types of data support this observation: 
A number of detailed geologic studies have been conducted at the Yucca Mountain site to define 
the basic mineralogy of the rocks and the petrologic and geochemical processes that occurred 
during the formation of the Topopah Spring Tuff and have continued from that time. These 
studies included a detailed description of the mineralogy of the repository host horizon from 
samples and observations developed from surface-based core holes through the repository block, 
as well as from the ESF and ECRB Cross-Drift. From the standpoint of mechanical degradation 
of the rock mass, these studies show: 
 The Topopah Spring Tuff is largely composed of fine-grained feldspars and silicatebased 
rocks that formed during the cooling of the rock mass shortly after deposition. 
Clay-forming minerals were typically not formed during the petrogenesis of the 
repository host horizon. 
 Clays typically do not form significant fracture-fill materials in the crystallized rocks of 
the Topopah Spring Tuff. 
 There are limitations on the environmental conditions needed to form clays and indicate 
the minimum likelihood that clays might form along fractures in the near field of a 
repository. Therefore, mineral alteration during the postclosure period is considered 
negligible. 
The following discussion summarizes the data obtained from a number of studies that support the 
above points. 
1. Mineralization in the Repository Host HorizonThe Topopah Spring Tuff is a 
compositionally zoned pyroclastic flow deposit. High-silica rhyolite forms 
approximately the lower two-thirds of the deposit, and trachyte (or quartz latite) forms 
the upper one-third of the deposit. The crystallized rocks that formed during the 
cooling of the deposit consist primarily of very fine-grained intergrowths of feldspar 
and silica polymorphs of quartz and cristobalite. These minerals typically crystallized 
from volcanic glass at temperatures of approximately 800C. X-ray diffraction 
analyses of 444 core samples from 13 boreholes that penetrate 3.5 km of the 
crystallized Topopah Spring Tuff indicate that there are different ratios of feldspar and 
silica polymorphs in the crystal-rich rocks (DTNs LADB831321AN98.002, 
LADV831321AQ97.001, LADV831321AQ99.001, LA000000000086.002). In the 
crystal-rich member, 93 samples have mean percentages of feldspar (69.1%), quartz 
July 2004 I-45 No. 4: Mechanical Degradation Revision 1 
(2.4%), cristobalite (13.3%), and tridymite (11.2%), and in the crystal-poor member, 
351 samples have mean percentages of feldspar (55.7%), quartz (15.8%), cristobalite 
(19.3%), and tridymite (5.1%). The total of the percentages cited for these two 
members do not sum to 100% because there are minor amounts of other minerals in 
many of these rocks. Smectite values vary from trace amounts to 15% with a mean 
value of 3.1%. These compositions of rocks form the walls of fractures where the 
minerals are very fine-grained and form various textures; however, the minerals are 
relatively uniformly distributed at scales of millimeters to centimeters. These data 
indicate that concentrations of smectite minerals along fractures do not occur as a 
result of crystallization during cooling. 
2. Clay Infillings Are Rarely Observed along Fractures in the ESF and ECRB 
Cross-DriftDetailed line survey data collected in the ECRB Cross-Drift indicate that 
of the 1,816 fractures in the 2.66-km-long tunnel, only 10 (or 0.4%) of the 
discontinuities, such as fractures, shears, and faults, have some amount of clay (DTNs: 
GS990408314224.001, GS990408314224.002). These 10 discontinuities are filled (or 
partially filled) with clay and broken or crushed rock or sand. Detailed studies of the 
clast textures, structures, and architectures of the broken or crushed rock or sand fill 
materials have not been completed, but general observations indicate many of these 
features do not show evidence of mechanical degradation and shear. Buesch and Lung 
(2003) describe volcaniclastic tuffaceous sandstone and claystone as fracture-fill 
material in the crystallized rocks of the Tiva Canyon and Topopah Spring Tuff. They 
proposed that volcanic glass particles settled by gravity along open fractures from the 
superjacent nonwelded bedded tuffs, and the clay formed in place (possibly millions of 
years later) by water seeping along the fractures and reacting with the glass. This 
mechanism is entirely consistent with occurrences of clays in the ECRB Cross-Drift 
detailed line survey data. Detailed line survey data from the crystallized Topopah 
Spring Tuff in the ESF main drift have not been recently reexamined, but only 4% of 
the fractures recorded have clay as filling material. 
Integrated Issue Resolution Status Report (NRC 2002) specifically discussed the work 
by Carlos et al. (1995) in regard to fracture fillings: 
Mineral-alteration products currently occur at Yucca Mountain mostly as 
fracture coating and as lithophysal-cavity deposits (Carlos, et al., 1995). 
The mineralogy, thickness, and amount and uniformity of coverage of 
fracture coatings are highly variable and uncertain (Thoma, et al., 1992). 
The coatings consist mainly of zeolites, manganese oxide minerals, silica 
phases, carbonates (mostly calcite), and clay minerals (mostly smectite but 
occasionally illite). Smectite is fairly ubiquitous in fractures throughout 
the volcanic sequence (Carlos, et al., 1995). 
Carlos et al. (1995) described the qualitative amount of materials in the fractures only, 
but typically they did not describe the amount or thickness of the mineral coatings. 
For example, if a mineral, such as clay (smectite), is listed as having major abundance, 
then greater than 20% of the minerals in the fractures are clay. However, the mineral 
coating might be less than 1 mm thick. Only for mordenite in the crystallized 
July 2004 I-46 No. 4: Mechanical Degradation Revision 1 
Topopah Spring Tuff did Carlos et al. (1995) describe about 1% of the fractures with 
this mineral and the amount increased with depth to more than 20% (and more than 
50% in some boreholes). The relevant point for this discussion is that selective 
fractures were sampled, and the intention was not to quantify the relative proportion of 
fractures with specific minerals or to quantify the thickness or continuity of the 
mineral coatings. However, there is one detailed study of mineral coatings on fracture 
walls in core with descriptions of the percentages of the amount of mineral coatings 
and thickness of the coatings (DTN: LA9912SL831151.001). Borehole ESF-HDTEMP-
2 is a 60 m (200 ft) long, horizontal borehole in the Drift Scale Test, and the 
rocks are in the middle nonlithophysal (Tptpmn) zone of the Topopah Spring Tuff. 
Only fractures in two 1.3 m (4.3 ft) long segments of core were described (7.6 to 8.9 m 
(25.0 to 29.3 ft) and 18.9 to 20.3 m (62.15 to 66.45 ft)). Clay forms localized deposits 
on the fracture surfaces that typically are less than 5% of the fracture surface area with 
thickness typically about 0.1 mm and a few deposits as much as 0.5 mm thick. These 
descriptions demonstrate that clay does not form continuous coatings on fracture walls 
and that the coats are very thin. 
3. Alteration of Repository Host Rocks Is Not Expected in the Repository 
EnvironmentConditions such as temperature, chemistry of water, and amount of 
water needed for the alteration of feldspars in the crystallized host rock to form clay 
and other sheet silicate minerals (sericite) are not considered to have been present 
within the Topopah Spring Tuff since its formation. Sericite is a general term for very 
fine-grained sheet silicates (illites) that form with other alteration products in 
hydrothermal systems at temperatures near or above 400C, typically in acidic 
aqueous solutions (Jackson 1997). As pointed out by the predicted temperature 
conditions expected in the repository environment (e.g., Figure I-24), these conditions 
are not anticipated in the near-field environment of the repository. Even in samples 
from the Drift Scale Test boreholes, the clays on fractures appeared to have been there 
prior to the test and were not affected by the experimental conditions. 
In summary, these petrologic and empirical relations indicate that clays typically do not form 
significant fracture-fill materials in the crystallized rocks of the Topopah Spring Tuff. They also 
indicate that the expected repository environmental conditions are not conducive to formation of 
clays during the postclosure period. Therefore, the impact of geochemical alteration within the 
postclosure environment is expected to have a negligible impact on drift degradation processes. 
July 2004 
I.4.5 Effect of Drift Collapse on the In-Drift Environment 
A parameter study was conducted to examine the impact of drift collapse on in-drift thermalhydrologic 
parameters (BSC 2004b). The multiscale thermal-hydrologic model was used to 
examine the effect of a rubble-filled drift on waste package and invert temperature and relative 
humidity at the waste package and invert. The drift is assumed to collapse (instantaneously) to 
twice the initial diameter (i.e., 11 m collapsed diameter) and is filled with rubble with a bulking 
factor of 0.231. The thermal conductivity of the rubble (Kth) is defined as the intact rock thermal 
conductivity of the Tptpll multiplied by the factor (1/(1 + bulking factor)). Two thermal 
conductivity values (a high case calculated as defined above for a bulking factor of 0.231, and a 
I-47 No. 4: Mechanical Degradation Revision 1 
low case, which is taken to be one-half the high case value) of the dry and wet rubble thermal 
conductivity were used in the analyses, as shown in Table I-1. 
Table I-1. Thermal Conductivity of Rubble 
Intact Host-Rock 
Property Value 
Basis for Rubble Property 
Value 
Intact Value  1/(1 + BF) 
(High-Kth rubble value)/2 
Host-Rock Rubble Property Value 
1.28 W/mK 1 W/mK (High-Kth case)a 
0.5 W/mK (Low-Kth case) 
Property 
Bulk dry thermal 
conductivity 
1.89 W/mK Bulk wet thermal 
conductivity 
Intact Value  1/(1 + BF) 
(High-Kth rubble value)/2 
1.515 W/mK (High-Kth case)b 
0.7575 W/mK (Low-Kth case) 
Source: BSC 2004b, Table 6.2-3. 
NOTE: a
b
This value is rounded down slightly 
Value is close to, but slightly less than, the value obtained from the Intact Value x 1/(1 + BF), in order to be 
consistent with the slight reduction made to the dry Kth value, which was rounded down. 
Figure I-24 shows the in-drift thermal-hydrologic parameters as functions of time from 
repository closure for the case of the hottest waste package, which is the 21-PWR Absorber Plate 
waste package. These plots show three cases: (1) an open, noncollapsed drift, (2) a collapsed, 
rubble-filled drift with high Kth for the rubble, and (3) a collapsed, rubble-filled drift with low 
Kth. The temperature (or any of the other environmental parameters plotted) will follow the 
intact drift curve until the assumed time of collapse. At that point, the temperature (or other 
parameters) will translate vertically to one of the other curves, depending on the assumed 
thermal conductivity of the rubble. 
th 
Examination of the waste package temperature curve (Figure I-24a) shows that significant 
impact to peak waste package temperature results only if drift collapse occurs within the first 100 
to 200 years after closure. After that time, the waste package temperature will always be below 
the peak temperature for the intact drift case, which occurs within about 20 to 30 years after 
closure. The total time at which the waste package surface remains above boiling for the hottest 
waste package case is approximately 1,000 years for the intact drift, 1,500 years for the high K 
case, and 2,000 years for the low Kth case. The relative humidity at the waste package decreases 
significantly for collapsed cases. 
July 2004 I-48 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004b, Figure 6.3-1. 
NOTE: The cases are: (1) intact-drift (nominal) case, (2) low-probability-seismic collapsed drift with high-Kth hostrock 
rubble, and (3) low-probability-seismic collapsed drift with low-Kth rubble. The plotted variables are (a) 
waste package temperature, (b) invert temperature, (c) waste package relative humidity, (d) invert liquidphase 
saturation, and (e) matrix liquid-phase saturation of the rubble surrounding the drip shield. 
Figure I-24. Thermal-Hydrologic Variables for the Hottest Waste Package (21-PWR Absorber Plate 
Waste Package) at the P2WR5C10 Location in the Tptpll (tsw35) Unit for the Mean 
Infiltration Flux Case 
July 2004 I-49 No. 4: Mechanical Degradation Revision 1 
I.4.6 Effect of Drift Collapse on Drip Shield Stability 
The drip shield design was developed to withstand the static load supplied by the weight of the 
collapsed drift. The methodology for estimating the rubble loading to the drip shield was 
described in detail in Section 5.3.2.3 of this technical basis document. The UDEC drift-scale 
model shown in Figure I-13 was used for determination of rubble loading. Six realizations of 
randomly shaped, 0.2 m side-length bonded blocks are used to describe the initial state of the 
rock mass. The shear and tensile strength of the rock mass was reduced from the initial 
calibrated rock mass strength until collapse was induced and was allowed to continue until the 
bulked rubble provided a back-pressure to the surrounding intact rock and effectively stopped 
further collapse (Figure I-25). The bulking factor of the rubble ranges from about 18% to 25%, 
depending on the particular realization (see Figure 5-58 of the technical basis document), and 
drift diameter expands to roughly 1.5 to 2 times its original size. 
Source: BSC 2004a, Appendix V, Figure V-6. 
Figure I-25. Quasi-Static Drift Degradation, 0.2 m Block Size: Equilibrium State for Deformable Drip 
Shield with Arched Top, Footings Free to Slide or Detach from the Invert 
The resulting pressure distribution around the drip shield was determined for each of the six 
realizations. The results show a highly nonuniform load distribution around the drip shield due 
to point load contacts when the small blocks become wedged against the drip shield 
(Figure I-26). 
July 2004 I-50 No. 4: Mechanical Degradation Revision 1 
Source: BSC 2004a, Figure 6-166. 
NOTE: Average pressure on each segment is shown for all six realizations. Segment numbering starts at 1 at the 
right footing and continues counterclockwise to the left footing. Those elements on the right, top, and left 
sides of the drip shield are shown. 
Figure I-26. Quasi-Static Pressure on Drip Shield Segments for Six Realizations for Random, 0.2-m 
Block Geometries 
The pressure distribution for each realization was used as input to a three-dimensional, nonlinear 
finite element structural analysis of the drip shield (BSC 2004c). The nonuniform pressure 
distribution was applied to the structure, followed by solution of the stress state and deformation. 
The analysis examines all of the potential failure modes including excessive deformation of the 
surface plates or load-bearing structure, development of yielded areas due to stress corrosion 
crack development, and potential buckling of the legs. It was found that the drip shield is stable 
under all quasi-static loading distributions. 
July 2004 I-51 No. 4: Mechanical Degradation Revision 1 
I.4.7 Conclusions 
RDTME 3.07 deals with the time-dependent degradation of emplacement drifts. This appendix 
presented discussion on five potential mechanisms or factors that could be associated with timedependent 
degradation of the rock mass. These factors are listed below, with a summary of the 
conclusions. 
Degradation due to the in situ stress state combined with the transient thermal and hydrologic 
stress conditions: 
 Numerical analyses were performed to examine the effects of the combined in situ and 
thermally induced stresses on the stability of the excavations for both nonlithophysal and 
lithophysal rock. A range of fracture geometries and mechanical properties spanning the 
expected range of conditions in the repository host horizon provided the basis for these 
analyses. The analyses show that yielding of the rock mass surrounding the excavations 
due to these sources of loading is minor for the entire postclosure period. 
 The potential impact of fluid vapor pressures along fractures in destabilizing the rock 
mass around the tunnel was investigated using the multiscale thermal-hydrologic model. 
The conclusion is that the fracture permeability is great enough that fluid or vapor 
pressure change is negligible during the entire heating and cooling cycle of the 
postclosure period. 
Degradation due to time-dependent fracture development in the rock matrix in the presence of 
water vapor and driven by mechanical and thermal stresses: 
 Time-dependent fracture development in the matrix in the presence of water vapor and 
in situ and thermal stresses was examined using discontinuum numerical modeling 
approaches. Laboratory creep experiments were conducted to define the static fatigue 
strength and rate of strength loss of Topopah Spring Tuff. Tests were conducted using 
water-saturated samples at temperatures at 125C or 150C. The samples were confined 
to prevent drying of the samples through boiling of water, thus ensuring water vapor was 
present to promote stress corrosion crack growth. Time-dependent results for granite 
determined for the Canadian high-level waste program were used as a means of 
comparison to the tuff data. 
 A mechanics-based, discontinuum model (particle flow code) was developed to simulate 
time-dependent fracture development in tuff. The model was calibrated to reproduce the 
time-to-failure behavior of tuff for varying applied constant stress levels. This model 
was then used to simulate the effect of lithophysal porosity on time-to-failure. 
 The UDEC drift-scale discontinuum model, which is used for drift stability simulation in 
lithophysal rocks, was then used to examine drift degradation due to time-dependent 
fracturing. The time-to-failure response generated from the laboratory testing of 
nonlithophysal rocks and the extrapolations for lithophysal rock was embedded into the 
drift-scale model through adjustment of rock properties. 
July 2004 I-52 No. 4: Mechanical Degradation Revision 1 
 Analyses of drift stability for the expected range of lithophysal rock mechanical 
properties variability indicate that relatively minor levels of time-dependent drift 
degradation are expected over the postclosure regulatory time frame. 
Degradation due to time-dependent fracture development and yield of asperities (roughness) on 
fracture surfaces: 
 Potential time-dependent degradation of surface roughness and cohesion was 
conservatively represented in three-dimensional discontinuum stability models by 
setting the cohesion and dilation angle to 0. In other words, a bounding case was 
assumed in which all surface roughness was assumed to have been destroyed due to 
time-dependent failure of the asperities. 
 Analyses of drift stability show that this assumption leads to only minor changes in 
rockfall because the stability of nonlithophysal rock is controlled primarily by fracture 
geometry. 
Degradation due to time-dependent alteration of rock matrix or joint filling materials due to rock 
mass thermal and moisture conditions: 
 The rock joints and rock matrix in the repository host horizon units are composed of 
minerals that will not undergo chemical alteration over postclosure temperature, 
moisture, and time frames. Fractures are characterized by rock wall contact with only 
minor amounts of clay minerals. 
 Time-dependent chemical alteration of rock joints or matrix has a minor impact on drift 
degradation. 
Degradation due to large excavation dimensions, unfavorable tunnel shape, and construction 
methods: 
 Repository excavations at Yucca Mountain have small dimensions in comparison to 
unsupported spans known to result in collapse from practice in the mining industry. 
 Emplacement drifts are to be excavated with a circular cross section using tunnel boring 
machines. The circular shape and nonexplosive excavation method are favorable to 
stability. 
I.5 REFERENCES 
I.5.1 Documents Cited 
Bieniawski, Z.T. 1989. Engineering Rock Mass Classifications. New York, New York: John 
Wiley & Sons. TIC: 226350. 
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Crystalline Rocks. Journal of Geophysical Research, 71, (16), 39393953. Washington, D.C.: 
American Geophysical Union. TIC: 226778. 
July 2004 I-53 No. 4: Mechanical Degradation Revision 1 
Brekke, T.L.; Cording, E.J.; Daemen, J.; Hart, R.D.; Hudson, J.A.; Kaiser, P.K.; and Pelizza, S. 
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2000/NT 4.0. 10828-2.0-00. 
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Vegas, Nevada: Bechtel SAIC Company. ACC: ENG.20040108.0001. 
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Nevada: Bechtel SAIC Company. ACC: MOL.20040513.0081. 
BSC 2004b. Multiscale Thermohydrologic Model. ANL-EBS-MD-000049 REV 02 DRAFT A. 
Las Vegas, Nevada: Bechtel SAIC Company. ACC: MOL.20040630.0537. 
BSC 2004c. Mechanical Assessment of the Drip Shield Subject to Vibratory Motion and 
Dynamic and Static Rock Loading. CAL-WIS-AC-000002 REV 00. Las Vegas, Nevada: 
Bechtel SAIC Company. ACC: MOL.20040630.0540. 
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International High-Level Radioactive Waste Management Conference (IHLRWM), March 30 
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Lining Minerals at Yucca Mountain, Nevada. LA-12977-MS. Los Alamos, New Mexico: Los 
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Sons. TIC: 218828. 
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July 2004 I-55 No. 4: Mechanical Degradation Revision 1 
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I.5.2 Data, Listed by Data Tracking Number 
GS030283114222.001. Direct Shear Data from Selected Samples of the Topopah Spring Tuff. 
Submittal date: 02/20/2003. 
GS990408314224.001. Detailed Line Survey Data for Stations 00+00.89 to 14+95.18, ECRB 
Cross Drift. Submittal date: 09/09/1999. 
GS990408314224.002. Detailed Line Survey Data for Stations 15+00.85 to 26+63.85, ECRB 
Cross Drift. Submittal date: 09/09/1999. 
LA000000000086.002. Mineralogic Variation in Drill Core UE-25 UZ#16 Yucca Mountain, 
Nevada. Submittal date: 03/28/1995. 
LA9912SL831151.001. Fracture Mineralogy of Drill Core ESF-HD-TEMP-2. Submittal date: 
01/05/2000. 
LADB831321AN98.002. Revised Mineralogic Summary of Yucca Mountain, Nevada. 
Submittal date: 05/26/1998. 
LADV831321AQ97.001. Mineralogic Variation in Drill Holes. Submittal date: 05/28/1997. 
LADV831321AQ99.001. Quantitative XRD Results for the USW SD-6 and USW WT-24 Drill 
Core Samples. Submittal date: 04/16/1999. 
SN0406L0212303.002. Static Fatigue Data from New England Research (NER) (Test Dates: 
1/2/2004 through 2/25/2004). Submittal date: 06/08/2004. 
I-56 July 2004 No. 4: Mechanical Degradation