Technical Basis Document No. 6: 
Waste Package and Drip Shield Corrosion 
Revision 1 
Prepared for: 
U.S. Department of Energy 
Office of Civilian Radioactive Waste Management 
Office of Repository Development 
1551 Hillshire Drive 
Las Vegas, Nevada 89134-6321 
Prepared by: 
Bechtel SAIC Company, LLC 
1180 Town Center Drive 
Las Vegas, Nevada 89144 
Under Contract Number 
DE-AC28-01RW12101 
QA: NA 
December 2003


1. INTRODUCTION 
This report provides a discussion of the degradation processes and long-term performance of 
waste packages and drip shields within the repository system. This report responds to Key 
Technical Issue (KTI) agreements (listed in Table 1-1) raised by the U.S. Nuclear Regulatory 
Commission (NRC) related to the adequacy of the representation of the testing environments and 
processes affecting the integrity of the waste package and drip shield during the repository 
regulatory period. It describes the degradation models, corrosion testing and treatment of 
uncertainty, appropriate for the environmental conditions predicted for the repository. 
This technical basis document provides a summary of the likely behavior of the waste package 
and drip shield in the repository after the permanent closure of the facility. This document is one 
in a series of technical basis documents prepared for each feature of the repository system 
relevant to predicting the likely postclosure performance of the repository. The relationship of 
waste package and drip shield performance (Process VI. Waste Package and Drip Shield 
Corrosion) to the other components is illustrated in Figure 1-1. 
Appendices to this technical basis document are intended to allow for a transparent and direct 
response to each KTI agreement. Each appendix addresses one or more of the KTI agreements. 
If agreements apply to similar aspects of the waste package or drip shield degradation process, 
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. 
A technical basis document on the in-drift chemical environment will provide the information on 
the evolution of the environment on the waste package and drip shield during the regulatory 
performance period. Based on the chemical environment described in that technical basis 
document, this report discusses the corrosion models developed, corrosion-related testing, and an 
assessment of other degradation processes. Section 1 discusses the materials selection and 
design and the integrated degradation model for the waste package and the drip shield. Section 2 
provides a description of the waste package and drip shield surface environments. The evolution 
of this local in-drift environment depends upon the interaction between warm waste packages, 
rocks in the drift wall, seepage water, dust deposits, and humidity. Section 2 describes these 
interactions and the resulting water compositions on the surface of the drip shield and waste 
package in sufficient detail to set the stage for subsequent sections. Sections 3 through 9 
describe the various degradation processes and the models that address them. Section 10 
addresses the integrity of the drip shield. Section 11 provides the conclusions derived from the 
individual models and their integration. Appendices A through H address specific KTIs 
associated with this technical basis document. 
December 2003 1-1 No. 6: Waste Package and Drip Shield Corrosion
Table 1-1. Key Technical Issue Agreements Addressed in this Report 
KTI/AIN/GEN Code 
CLST 1.07 
CLST 1.07 AIN-1 
CLST 1.12 
CLST 1.13 
CLST 1.14 
CLST 1.15 
CLST 1.16 
RDTME 3.18 
CLST 6.02 
No. 6: Waste Package and Drip Shield Corrosion 
Revision 1 
Wording 
Provide documentation for the alternative methods to measure corrosion rates of the 
waste package materials (e.g., ASTM G-102 testing) or provide justification for the 
current approach. DOE will document the alternative methods of corrosion 
measurement in the revision of Alloy 22 AMR (ANL-EBS-MD-000003), prior to license 
application. 
The use of appropriate standards should be adopted and exceptions should be 
properly justified. If a standard is mentioned but not used in its entirety, DOE should 
indicate specifically which parts of the code will be used (example, G1 is used for 
identification of equipment, G1 is used for data interval, etc.). Better justification for 
not using alternative techniques is needed. If such a justification cannot be provided, 
then DOE must provide details on the alternative techniques to be used for corrosion 
rate measurements. 
Provide the documentation for Alloy 22 and titanium for the path forward items listed 
on slides 34 and 35. DOE will provide the documentation in a revision to AMRs 
(ANL-EBS-MD-000005 and ANL-EBS-MD-000006) prior to LA. 
Provide the data that characterizes the distribution of stresses due to laser peening 
and induction annealing of Alloy 22. DOE will provide the documentation in a revision 
to AMR (ANL-EBS-MD-000005) prior to LA. 
Provide the justification for not including the rockfall effect and deadwood from drift 
collapse on SCC of the waste package and drip shield. DOE will provide the 
documentation for the rockfall and dead-weight effects in the next revision of the SCC 
AMR (ANL-EBS-MD-000005) prior to LA. 
Provide the documentation for Alloy 22 and titanium for the path forward items listed 
on slide 39. DOE will provide documentation for Alloy 22 and Ti path forward items on 
slide 39 in a revision to the SCC and General and Localized Corrosion Analysis and 
Model Reports (ANL-EBS-MD-000003, ANL-EBS-MD-000004, ANL-EBS-MD-000005) 
by license application. 
Provide the documentation on the measured thermal profile of the waste package 
material due to induction annealing. DOE stated that the thermal profiles will be 
measured during induction annealing, and the results will be reported in the next 
stress corrosion cracking analysis model report (ANL-EBS-MD-000005) prior to LA. 
Provide a technical basis for a stress measure that can be used as the equivalent 
uniaxial stress for assessing the susceptibility of the various EBS materials to stress 
corrosion cracking. The proposed stress measure must be consistent and compatible 
with the methods proposed by the DOE to assess SCC of the containers in WAPDEG 
and in accordance with the agreements reached at the Container Life and Source 
Term Technical Exchange. DOE will include a detailed discussion of the stress 
measure used to determine nucleation of stress corrosion cracks in the calculations 
performed to evaluate waste package barriers and the drip shield against stress 
corrosion cracking criterion. DOE will include these descriptions in future revisions of 
the following: Design Analysis for UCF Waste Packages, ANL-UDC-MD-000001, 
Design Analysis for the Defense High-Level Waste Disposal Container, ANL-DDC-ME- 
000001, Design Analysis for the Naval SNF Waste Package, ANL-UDC-ME-000001, 
and Design Analysis for the Ex-Container Components, ANL-XCS-ME-000001. The 
stresses reported in these documents will be used in WAPDEG and will be consistent 
with the agreements and associated schedule made at the Container Life and Source 
Term Technical Exchange (Subissue 1, Agreement 14, Subissue 6, Agreement 1). 
Provide additional justification for the use of a 400 ppm hydrogen criterion or perform a 
sensitivity analysis using a lower value. DOE stated that additional justification will be 
found in the report “Review of Expected Behavior of Alpha Titanium Alloys under 
Yucca Mountain Condition” TDR-EBS-MD-000015, which is in preparation and will be 
available in January 2001. 
December 2003 1-2
Table 1-1. 
KTI/AIN/GEN Code 
CLST 6.03 
CLST 6.02 AIN-1 
CLST 6.03 AIN-1 
GEN 1.01 Comment 9 
GEN 1.01 (Comment 10) 
GEN 1.01 (Comment 119) 
GEN 1.01 (Comment 120) 
NOTE: AMR = analysis and modeling report; DOE = U.S. Department of Energy; LA = license application; EBS = 
engineered barrier system; SNF = spent nuclear fuel; SCC = stress corrosion cracking. 
No. 6: Waste Package and Drip Shield Corrosion 
Revision 1 
Key Technical Issue Agreements Addressed in this Report (Continued) 
Wording 
Provide the technical basis for the assumed fraction of hydrogen absorbed into 
titanium as a result of corrosion. DOE stated that additional justification will be found 
in the report “Review of Expected Behavior of Alpha Titanium Alloys under Yucca 
Mountain Condition” TDR-EBS-MD-000015, which is in preparation and will be 
available in January 2001. 
1. Provide better justification to verify the critical hydrogen concentration chosen is a 
realistic and representative value for the onset of cracking in Titanium Grade 7. An 
evaluation of the critical hydrogen concentration for Titanium Grade 24 should also be 
given. 
2. Provide an evaluation of possible detrimental effects of the fluoride on hydrogen 
uptake and its effects on the critical hydrogen concentration and subsequent cracking. 
3. Provide the results for the Titanium Grade 24 structural drip shield components. 
Provide an evaluation of the possible detrimental effects of fluoride on possible 
hydrogen uptake rates, as well as enhanced corrosion resulting in higher than 
currently estimated hydrogen generation rates. 
Data supporting the residual stress calculations as a result of welding, after laser 
peening and after induction annealing are not provided. 
Basis: The distribution of residual stresses in the waste package final closure welds is 
based on Finite element modeling. Details of the Model are provided in the Stress 
Corrosion Cracking of the Drip Shield, the Waste Package Outer Barrier, and the 
Stainless Steel Structured Material AMR. The effects of induction annealing on the 
residual stresses in the final closure are detailed in the Residual Stress Minimization of 
Waste Packages from Induction Annealing AMR. Several assumptions are made in 
the models that are not supported by data. These include the assumed temperature 
profile during welding, the cooling rates during welding and the residual stress during 
induction annealing. 
The distribution of residual stress in the inner closure weld after laser peening is 
estimated in the SSPA using a shot-peened Incoloy 908 specimen. The technical 
basis for using a shot-peened specimen is not provided. Differences in the residual 
stress mitigation methods (i.e. mechanical shot-peening vs. laser peening) may result 
in significantly different stress distributions. 
The modified stress corrosion cracking parameters are based in recent tests that may 
not consider the range of possible environments and the effects of fabrication 
processes. 
Basis: The SSPA uses modified parameters for the stress corrosion cracking 
including the repassivation rate for the slip dissolution model and the minimum 
threshold stress for stress corrosion cracking. The SSPA indicates that these new 
parameters are based on recent data. The particular importance is the change in the 
minimum threshold stress which has been increased from 20-30 to 80-90 percent of 
the yield strength. The value of this parameter which is used in the model abstraction 
as the critical parameter for the occurrence of SCC is likely to be dependent on 
several factors that have not been investigated such as chemical composition of the 
environment and the effects of fabrication processes (only a limited number of cold 
worked and welded specimens has been evaluated). 
In p. 7-9 DOE claimed that NRC accepted the slip dissolution model. The DOE must 
supply the reference for this acceptance. 
Page 7-11, the use of the triangular distribution for the residual stress uncertainty 
dictates that the endpoints of the distribution are well known. Showing the data 
compared to the distribution would support the selection of a triangular distribution. 
December 2003 1-3
Revision 1 
Figure 1-1. Components of the Postclosure Technical Basis for the License Application 
1.1 WASTE PACKAGE AND DRIP SHIELD MATERIALS SELECTION, DESIGN, 
AND FABRICATION 
Spent nuclear fuel and high-level radioactive waste (70,000 MTHM) will be placed in 
approximately 11,000 waste packages (Figure 1-2). The design for the waste package is based 
upon double-wall construction (Plinski 2001). The inner wall serves as a structural support and 
is constructed from Type 316 stainless steel with controls on carbon and nitrogen levels. The 
corrosion-resistant waste package outer shell is constructed from Alloy 22 (UNS N06022), a 
high-performance nickel-based alloy. Solution heat treatment of the as-fabricated disposal 
container is used to remove fabrication-related residual tensile stress (BSC 2003a, Section 8.1). 
After the waste package is filled with high-level radioactive waste or spent nuclear fuel, at the 
site hot cell facility, the stainless steel cylinder is sealed and two closure lids made of Alloy 22 
will be welded to the container outer shell by gas tungsten arc welding. Because residual weld 
stress at the closure welds might initiate stress corrosion cracking (SCC), a postweld stress 
mitigation process will be used to reduce the tensile stress in the outer surface, thereby 
eliminating the stress initiator. The baseline stress mitigation process is laser-shock peening 
(also called laser peening). The titanium alloy drip shield is used to prevent seepage water, as 
well as rocks, from directly impinging on the waste package. They will be installed in the stressrelieved 
condition, following emplacement of the waste packages in the drifts. The stress relief 
treatment is used to reduce fabrication-related tensile stress below the SCC initiation threshold. 
December 2003 1-4 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
and High-Level Radioactive Waste 
Figure 1-2. Typical Waste Package Types That Will Be Used for the Storage of Spent Nuclear Fuel 
Materials Selection for the Waste Package Outer Shell–Alloy 22 (UNS N06022) was selected 
for its excellent corrosion resistance in brines and will be used for construction of the outer shell 
of the waste package. It has already been used for the construction of mockups. It consists of 
20.0 to 22.5 percent chromium, 12.5 to 14.5 percent molybdenum, 2.0 to 6.0 percent iron, 2.5 to 
3.5 percent tungsten, 2.5 percent (max.) cobalt, 0.015 percent (max.) carbon, and balance nickel 
(BSC 2003b). Other elements present include phosphorus, silicon, sulfur, and manganese 
(Treseder et al. 1991). The localized corrosion resistance of Alloy 22 is due to the additions of 
molybdenum and tungsten, both of which stabilize the passive film at low pH (Hack 1983). The 
oxides of these elements are insoluble at low pH. Consequently, Alloy 22 is highly resistant to 
localized attack. Very high repassivation potentials have been observed by some (Gruss et al. 
1998), while others have found very low corrosion rates in simulated crevice solutions 
containing 10 weight percent FeCl3 (Haynes International 1997). Furthermore, no localized 
attack of Alloy 22 has been seen in crevices exposed to water compositions representative of 
December 2003 1-5 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
most of those expected in the repository. Such tests have been conducted in the Long Term 
Corrosion Test Facility (LTCTF) (see Section 5). 
Drip Shield Materials–The drip shields will be fabricated using titanium alloys containing small 
palladium additions to provide corrosion resistance in the anticipated postclosure drift 
environment. Titanium Grade 7 (UNS R524000), an alpha-phase alloy, was selected for the 
plate material, which will be supported using higher strength welded ribs and bulkheads made 
from Titanium Grade 24 (UNS R56405), an alpha-beta alloy. Similar to Alloy 22, based on 
laboratory test results, literature, and model predictions, these titanium alloys are expected to 
provide a high level of resistance to the various potential corrosion-related degradation modes, 
including localized corrosion and hydrogen-induced cracking. 
1.2 INTEGRATED MODEL FOR WASTE PACKAGE AND DRIP SHIELD 
DEGRADATION 
Integrated Model–Systematic interactions between the in-drift environment and the drip shield 
and waste package outer shell result in the occurrence of various degradation modes, which are 
shown schematically in Figure 1-3. This schematic accounts for a wide range of degradation 
modes, each operable to varying degrees, depending upon the temperature regime. In the 
preclosure regime, the walls of the drifts will be kept dry by ventilation air, and no significant 
degradation of the waste packages is expected. In the postclosure dryout regime (when 
temperature is greater than or equal to the boiling point of seepage water), potentially relevant 
high-temperature modes of degradation include dry oxidation and corrosion underneath 
deliquescent brines (formed by the adsorption of airborne water by dust deposits). In the 
transition regime (temperature is approximately the boiling point range of seepage water), 
seepage waters could concentrate on the waste package through evaporation if the drip shield 
were to fail. The potential modes of attack in these concentrated brines include uniform general 
corrosion, localized corrosion, and SCC. In the low temperature regime, seepage waters can 
enter the drift, but the thermal driving force for localized corrosion would be less. The possible 
modes of attack include general corrosion, localized corrosion, and SCC. Microbially influenced 
corrosion (MIC) may also occur. These modes of failure are dealt with in this technical basis 
document, and are illustrated in Figure 1-3. As described in subsequent sections, not every 
degradation mode represented in Figure 1-3 may occur on either the waste package or the drip 
shield in the repository (e.g., the drip shield material is not subject to localized corrosion or 
MIC); however, Figure 1-3 is an illustration of the conceptual model used to consider the 
possible degradation modes. In addition to these scenarios, other possible events that could lead 
to failure include seismic and volcanic activity. These scenarios are examined in the technical 
basis documents on seismic events and igneous events. In addition, effects of combined features, 
events, and processes are not discussed here. 
December 2003 1-6 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
NOTE: TTT = time-temperature-transformation. 
Figure 1-3. Systematic Interactions between the In-Drift Environment and the Drip Shield and Waste 
Package Outer Shell Resulting in the Occurrence of Various Degradation Modes 
Phase Instability–The Titanium Grade 7 drip shield material is a stable alpha (á) phase alloy 
and possesses outstanding phase stability (BSC 2003c, Section 6.5.3). While Titanium Grade 7 
does contain small amounts of alloying elements, most notably palladium, it is essentially pure 
titanium, which will not form intermetallic compounds under the thermal exposure conditions in 
the repository. On this basis, Titanium Grade 7 is considered immune to the effects of phase 
instability under repository exposure conditions. 
Alloy 22 is a metastable austenitic alloy, where secondary complex phases can be precipitated at 
high temperature (BSC 2003d). These intermetallic phases (P, ó, and µ) are rich in those 
elements responsible for the exceptional corrosion resistance of Alloy 22, and can therefore 
cause depletion of the passivating elements in close proximity to the precipitates. Such depleted, 
localized areas may be more susceptible to corrosion than areas where no precipitation has 
occurred. Furthermore, these precipitates may cause degradation of the mechanical properties. 
In addition to the higher temperature complex phases, long-range ordering can occur at 
somewhat lower temperatures and could potentially negatively affect SCC susceptibility. 
Multicomponent phase diagrams have enabled prediction of the phases that are possible as a 
function of elemental composition and temperature. Time-temperature-transformation diagrams 
have been predicted, compared with experimental kinetic measurements, and used to calculate 
the rate at which deleterious phases (precipitates and long-range ordering) form at timetemperature 
combinations relevant to the repository. As will be discussed in the Section 3 
summary, such predictions indicate insignificant precipitation and long-range ordering for the 
December 2003 1-7 No. 6: Waste Package and Drip Shield Corrosion
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mill-annealed and for the as-welded conditions during the 10,000-year simulation period, as long 
as the temperature is kept below 200°C (or 300°C for shorter times). 
Dry Oxidation–Dry oxidation is another high-temperature degradation mode (BSC 2003b; 
CRWMS M&O 2000a). The reaction of oxygen with Alloy 22, or Titanium Grade 7, can cause a 
uniform thickening of the oxide layer on the surface. The surface oxide can consist of any of the 
alloying elements or constituents, although for Alloy 22 a chromium oxide layer has been 
assumed as the basis of calculations. The rate of metal loss due to dry oxidation at the highest 
temperature in the repository is insignificant compared to other possible modes of attack. Dry 
oxidation is not a performance limiting process of the waste package outer shell under the 
exposure condition expected in the repository and is not considered in the waste package 
performance analysis for the repository (BSC 2003b, Section 8). Additional details are presented 
in Section 4. 
General Corrosion–General corrosion of the waste package outer shell and the drip shield 
occurs when the relative humidity at the waste package surface is equal to or greater than the 
relative humidity threshold (RHthreshold) for corrosion initiation. The general corrosion rate for 
the drip shield is temperature independent and does not decrease with exposure time. The 
general corrosion rate of the waste package outer shell is a function of temperature, expressed 
with an activation energy using a modified Arrhenius relationship. Because of the very low 
general corrosion rates of the waste package outer shell for the conditions expected in the 
repository, waste package performance is not limited by general corrosion during the regulatory 
period. Additional details are provided in Section 5. 
MIC: The drip shield is immune to MIC under the repository exposure conditions (BSC 2003c, 
Section 6.5.2). The waste package outer shell is potentially subject to MIC when the relative 
humidity at the waste package surface is equal to or greater than 90 percent. The effect of MIC 
is represented by an enhancement to the abiotic general corrosion rate of the waste package outer 
shell (BSC 2003b, Section 8). 
Corrosion Underneath Deliquescent Brines–As the temperature begins to decrease during 
cooling, the possibility exists of concentrated electrolytes developing through the formation of 
deliquescent brines and the evaporative concentration of seepage waters. Modes of attack 
experienced in aqueous electrolytes include general corrosion, localized corrosion, and SCC. In 
the case of general corrosion, the rate of dissolution is uniform over all surfaces and is due to the 
transport of cations from the metal-oxide interface to the oxide-electrolyte interface, where the 
formation of soluble metal-containing corrosion products can occur. Cations and anions are 
transported through the very thin passive oxide barrier film adjacent to the metal surface, with 
the growth kinetics controlled by diffusion and electric field-driven electromigration (BSC 
2003b, Section 6.4.1.1.2). 
Localized Corrosion (Pitting and Crevice Corrosion)–Pitting corrosion is any type of 
distributed, nonuniform corrosive attack of the surface, and is due to the localized failure of the 
passive film. Such localized failure may be initiated due to surface inclusions of relatively 
soluble species, the precipitation of small soluble halide crystallites on the passive film, or 
destabilization of the passive film within occluded areas such as crevices. This destabilization is 
due to the lowering of pH, which results from the combined effects of differential aeration, the 
December 2003 1-8 No. 6: Waste Package and Drip Shield Corrosion
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hydrolysis of dissolved metal cations within the crevice, and the electric field-driven 
electromigration of aggressive halide anions into the crevice. Localized attack may also occur at 
sites where precipitation has occurred. Both the drip shield and waste package outer shell are 
resistant to pitting corrosion under a range of the exposure conditions in the repository. 
Crevices may be formed between the waste package and supports, beneath mineral precipitates, 
corrosion products, dust, rocks, cement, and bio-films, and between layers of the containers. In 
the absence of inhibitor and buffer ions, the hydrolysis of dissolved metal can lead to the 
accumulation of H+ and a corresponding decrease in pH. Electromigration of Cl- (and other 
anions) into the crevice must occur to balance cationic charge associated with H+ ions (Gartland 
1997; Walton et al. 1996; Farmer and McCright 1998). These exacerbated conditions can set the 
stage for subsequent attack of the corrosion resistant material by passive corrosion, pitting 
(initiation and propagation), SCC, or other mechanisms (Farmer and McCright 1998; Farmer, Lu 
et al. 2000; Farmer, McCright et al. 2000). Gamma radiolysis may produce hydrogen peroxide, 
thereby increasing open-circuit corrosion potential (Glass et al. 1986; Kim 1987, 1988, 1999a, 
1999b). 
3 
The drip shield is not subject to crevice corrosion under the exposure conditions in the repository 
(BSC 2003c, Section 8.4). The waste package outer shell is not subject to crevice corrosion if 
the solution contacting the waste package contains significant concentrations of inhibitive ions 
such as nitrate. The waste package outer shell is, however, potentially susceptible to crevice 
corrosion if an acidic chloride-containing solution with a low NO -/Cl- ratio contacts the waste 
package while it is at elevated temperatures. Additional details are provided in Sections 6 and 7. 
However, no crevice corrosion was observed after 5-year exposure to a range of brines in the 
Long Term Corrosion Test Facility. 
Stress Corrosion Cracking–SCC is the initiation and propagation of cracks in structural 
components due to the simultaneous interaction of three factors: metallurgical susceptibility, 
critical environment, and static (or sustained) tensile stresses. Both Alloy 22 and Titanium 
Grade 7 are susceptible to SCC under repository exposure conditions. The drip shields 
(including their fabrication welds) will be fully stress-relief annealed before placement in the 
drifts (Plinski 2001, Section 8.3.17). Therefore, the drip shield is not subject to SCC upon 
emplacement in the repository. However, the drip shields are potentially subject to SCC under 
the action of seismic-induced loading and rockfall. An analysis of the consequence of SCC of 
the drip shield (BSC 2003a, Section 6.3.7) indicates that stress corrosion cracks in passive alloys 
such as Titanium Grade 7 tend to be very tight and could be plugged by corrosion products or 
mineral deposits and, thereby, preventing water transport. Since the primary role of the drip 
shield is to keep water from contacting the waste package, SCC of the drip shield does not 
compromise its intended design purpose and is not of significant consequence to repository 
performance. 
Similar to the drip shield, all regions of the waste package (including fabrication welds) except 
the waste package closure lid welds are stress-relief annealed before the waste packages are 
loaded with waste (Plinski 2001, Section 8.1.7) and, thus, do not develop residual stress or 
stress-intensity factors high enough for SCC to occur (BSC 2003a, Section 6.4.2). Residual 
stress mitigation techniques are applied to the waste package outer closure lid weld region to 
induce compressive stresses in the outer layers and delay the initiation of crack growth until 
December 2003 1-9 No. 6: Waste Package and Drip Shield Corrosion
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these layers are removed by general corrosion processes. The SCC model makes use of a 
threshold stress for the initiation of stress corrosion cracks on smooth surfaces and a stressintensity 
factor threshold for the initiation of crack growth. Manufacturing flaws (e.g., weld 
defects) are also considered to propagate by SCC. It is conservatively assumed that 
manufacturing flaws behave like stress corrosion cracks and, thus, only the stress intensity factor 
threshold for crack growth is applied to the manufacturing flaws. 
1.3 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 Project’s 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 
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. 
December 2003 1-10 No. 6: Waste Package and Drip Shield Corrosion
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2. THERMAL OPERATING REGIMES FOR WASTE PACKAGE 
AND DRIP SHIELD 
2.1 EXPECTED IN-DRIFT ENVIRONMENT 
The current repository design enables the repository to operate in three temperature regimes: 
dryout, transition, and low-temperature. The relevant attributes of each regime are summarized 
below. 
Figures 2-1 and 2-2 show a representative time-temperature and time-relative humidity profile 
for a variety of waste packages (located near the center of the repository) in the three operational 
regimes. 
Source: BSC 2003e. 
NOTE: These histories are plotted for the P3R7C12 location (in the mountain scale thermal-hydrologic model), 
which is near the center of the repository in the Tptpll (tsw35) unit. PWR = pressurized water reactor; WP = 
waste package; DHLW = defense high-level radioactive waste; BWR = boiling water reactor. 
Figure 2-1. Typical Waste Package Temperature Histories 
December 2003 2-1 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
Source: BSC 2003e. 
NOTE: These histories are plotted for the P3R7C12 location (in the mountain scale thermal-hydrologic model), 
which is near the center of the repository in the Tptpll (tsw35) unit). PWR = pressurized water reactor; 
WP = waste package; DHLW = defense high-level radioactive waste; BWR = boiling water reactor. 
December 2003 
Figure 2-2. Typical In-Drift Relative Humidity Histories 
Dryout–Drift walls will first be dried by ventilation air during the preclosure period. Eventually 
during postclosure, heat generated by radioactive decay will increase the temperature of waste 
packages and drift walls above the boiling point of water. Since no significant seepage is 
expected for drift wall temperatures above the boiling point of water, no aqueous phase corrosion 
due to seepage is expected (calcium chloride type brines are possible and predicted to occur in 
this regime, but they occur in the host rock when temperatures are above boiling and seepage 
into the drift is prevented by the vaporization barrier effect as described in the technical basis 
document on the in-drift chemical environment). However, depending on the surface 
temperature and relative humidity conditions, the existence of liquid-phase water on the waste 
package or drip shield is possible due to the presence of a dust or salt deposit. In the presence of 
such a deposit, a thin-film liquid phase can be established at a higher temperature and lower 
relative humidity than otherwise possible. Thus, formation of deliquescent brines in the absence 
of seepage may occur, and corrosion of the waste package and drip shield is considered in the 
context of these solutions. 
Transition–Seepage into the drifts will become possible as the waste package cools, as the 
temperature of the drift wall drops below the boiling point of water, and while the waste package 
2-2 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
surface temperature is at or above the boiling point of the water. Seepage waters will undergo 
evaporative concentration on the drip shield surface or the waste package surface at the time 
when the drip shield seepage diversion function is lost, thereby evolving into either carbonate- or 
sulfate-type brines. The drip shield will mitigate seepage effects on the waste package and is 
expected to last through this period, unless there are low-probability seismic events that would 
shorten its performance lifetime. However, as in the dryout regime, formation of deliquescent 
brines could occur in this regime. 
Low Temperature–As the waste package cools to a temperature below the boiling point of 
water, the in-drift relative humidity will increase, so evaporated solutions cannot be as 
concentrated. With further cooling, the temperature will drop to below the threshold for 
localized corrosion for the repository-relevant environments. This threshold temperature is a 
function of the presence of beneficial ions, such as nitrates and sulfates. 
2.2 RELATION OF IN-DRIFT CHEMICAL MODEL RESULTS TO CORROSION 
TESTING ENVIRONMENTS 
The project has developed an understanding of the in-drift chemical environment for the three 
regimes described in Section 2.1. The understanding is based on geochemical models and 
supporting data and analysis appropriate for the repository conditions. A detailed description of 
the evolution of the chemical environment is provided in the in-drift chemical environment 
technical basis document, which includes a discussion of the relationship between the 
geochemical process model results and the chemical environments used in corrosion related 
testing. A high-level summary of the chemical environment applicable to corrosion related 
testing extracted from the technical basis document on in-drift chemical environment follows. 
The geochemical model developed by the project to characterize the range of expected in-drift 
environments is presented in Engineered Barrier System: Physical and Chemical Environment 
Model (BSC 2003f). The model output is in the form of lookup tables, listing ion concentrations 
and pH as a function of relative humidity, temperature, and carbon-dioxide partial pressure. 
Brines that develop on the waste packages and drip shields will be the result of either evaporative 
concentration of seepage water or deliquescence of deposited salts. Deposited salts can be due to 
entrained matter in the ventilation air, dust and debris deposited within the drifts, or seepage 
waters that have evaporated to dryness. Seepage waters are not expected to enter the drifts until 
host rock temperatures fall below 100°C. Dust salts will be able to deliquesce water from the 
atmosphere to form thin films on waste packages and drip shields above the normal boiling point 
of water (up to about 140°C) (BSC 2003f). 
The water types or bins that have been predicted with the in-drift environment geochemical 
model for drift crown seepage waters are listed in Table 2-1. (This water binning methodology 
is discussed in detail in the technical basis document on the in-drift chemical environment.) The 
crown seepage waters are those anticipated to contact the drip shields and the waste package 
outer shell after the drip shield failure once the temperature of the drift wall falls below 100°C. 
The seepage waters have been categorized into 11 bins, or water types, based on their chemical 
composition (BSC 2003f, Section 6.6). A characteristic water has been determined for each bin. 
These seepage waters are characterized by their dominant constituents shown in the table near 
December 2003 2-3 No. 6: Waste Package and Drip Shield Corrosion
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the beginning of the evaporation process (identified in the table as at 98 percent relative 
humidity) and under conditions just before dryout. The time-integrated occurrence fraction that 
water in a particular bin will form at the drift crown is also listed. The associated brine type 
described and the test solution in which corrosion testing was conducted for each bin are also 
listed in the table and discussed in the following paragraphs. 
Less than 1 percent of the time-integrated drift crown waters fall into Bins 1 through 3, which 
would be expected to be CaCl2-type brines. These brine types are projected to occur in the rocks 
above the drift during the time when the drift wall temperature is above 100°C. Most of the 
time-integrated waters (99 percent or more) fall into Bins 4 through 11, which evolve to sodiumand 
potassium-dominated solutions. The standard solutions used for corrosion testing by the 
project have also been categorized in terms of the representative bins. 
Corrosion testing to determine the response of waste package, drip shield, and other in-drift 
materials is carried out in environmental conditions consistent with those predicted by in-drift 
chemical modeling. Corrosion testing environments were chosen based on the three types of 
natural brines: calcium chloride, carbonate, and sulfate. Initial studies focused on the 
carbonate-type brine, based on reasoning that carbonate-type waters, typified by J-13 well water 
from the saturated zone near Yucca Mountain, were the expected types of waters at the 
repository (Harrar et al. 1990). Details of the aqueous test solutions are given after the 
discussion of the brine types. 
The brine type name reflects a characteristic that distinguishes it from the other brines. In terms 
of sulfate brines and in other cases where modeling has gone beyond the classic chemical divides 
(see the in-drift chemical environment technical basis document), it does not necessarily reflect 
the dominant species in the brine. This characterization of surface brine types has, in part, 
guided the expected range of brine water chemistry in the repository. However, some 
differences are expected between brines formed at the earth’s surface and brines formed in the 
repository. These differences are mainly due to differences in the chemistry of seepage waters 
and surface waters giving rise to brines, and differences between the salt chemistry of dust and 
the dissolved salt content of such surface waters. Two important general factors specific to the 
repository brines are the presence of nitrate and more effective mechanisms for the removal of 
magnesium. It is expected that nitrate will be present in the deliquescent brines owing to 
multiple potential sources (BSC 2003g, Section 6.7.2.8) and the generally high solubility of 
nitrate minerals (BSC 2003g, Section 4.1.1.7). It is expected that magnesium will not be 
significant owing to a combination of low source (for the dust, as well as for at least some 
groundwaters) and multiple removal mechanisms, most of which are enhanced by elevated 
temperature (BSC 2003g, Sections 6.7.2.10 and 6.7.2.11). 
The calcium-chloride brines have near neutral pH and no significant bicarbonate/carbonate, 
fluoride, or sulfate content. These brines may contain other cations such as sodium, potassium, 
and magnesium and other anions such as nitrate, but not carbonate, sulfate, or fluoride. The 
endpoint of the evaporative concentration of this type of brine would contain Ca-Cl/NO3 or a 
mixture of Ca/Mg-Cl/NO3. The quantity of magnesium and calcium in this type of brine would 
be limited due to the precipitation of calcium carbonates and sulfates, and magnesium silicates. 
This is consistent with information on saline lakes where sodium is the dominant cation with the 
percentage of calcium varying from insignificant to about 20 percent (Drever 1997). In the 
December 2003 2-4 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
repository, the concentration of magnesium in any type of brine is expected to be insignificant as 
noted earlier. Thus, a magnesium chloride brine is not expected. Nitrate is expected to be 
present, and an end-point brine of this type is likely to be dominated by calcium chloride and 
calcium nitrate. A brine of the calcium chloride type is expected to have a very limited 
occurrence in the repository, as indicated in Table 2-1. For brine generated by dust 
deliquescence, the brine is actually expected to be more of a potassium nitrate–sodium chloride 
brine with only a small probability of calcium present due to the compositional nature of the dust 
leachate. 
Relative humidity dependence of the calcium-chloride brine composition is as follows. At low 
relative humidity, the aqueous solutions will be dominated by calcium cations (very low sodium 
and potassium) and chloride and nitrate anions, since both calcium nitrate and calcium chloride 
are very soluble. At higher relative humidity, chloride and nitrate salts of sodium and potassium 
become soluble and could dominate the aqueous solution compositions. This would occur at or 
above the deliquescence relative humidity for salts composed of these ions. 
Corrosion test solutions corresponding to this calcium chloride type of brine include calcium 
chloride, calcium chloride plus calcium nitrate, the simulated saturated water (SSW), and sodium 
chloride aqueous solutions. The SSW and sodium chloride test solutions simulate the moderate 
relative humidity scenario where calcium is a minor component in the aqueous solution. 
Numerous electrochemical studies were performed in these test solutions. Thin film studies were 
also performed using these types of solutions on coupons of Alloy 22 using an environmental 
thermogravimetric analyzer. See Appendix A for the corrosion tests performed in these types of 
solutions. 
The carbonate brines are alkaline and do not contain significant calcium or magnesium content. 
In the early stages of the evaporative concentration, calcium precipitates predominately as 
carbonate mineral (calcite or aragonite) under equilibrium conditions. Magnesium precipitates 
as a minor component in the calcium carbonate species and as magnesium silicate. In the 
repository, it is expected that magnesium will be removed even more efficiently. Potassium may 
be significant in some of these brines. Nitrate is expected to be an important component, and a 
brine of this type may evolve through a high extent of evaporation into one in which nitrate is 
actually the dominant anion. The carbonate brine is likely to be represented as alkali metal 
(sodium, potassium) carbonate brine. 
Relative humidity dependence of carbonate brine composition is as follows. At low relative 
humidity, the aqueous solutions will be dominated by nitrate and chloride anions with nitrate 
ions dominating at the lowest relative humidity. At moderate relative humidity (greater than 
70 percent relative humidity), chloride ions could dominate the solution composition. The 
nitrate-chloride solutions are expected to have slightly elevated pH due to residual carbonate in 
solution. Significant amounts of carbonate and sulfate ion are not expected until the relative 
humidity is greater than 85 percent. 
Corrosion test solutions corresponding to the carbonate type of brine include the simulated dilute 
water (SDW), simulated concentrated water (SCW), basic saturated water (BSW), and under 
certain circumstances, SSW and simulated acidic water (SAW) aqueous test solutions (see 
Table 2-2). The BSW test solution is a highly concentrated alkaline solution and could be 
December 2003 2-5 No. 6: Waste Package and Drip Shield Corrosion
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expected under repository conditions where temperatures could be at its measured boiling point 
of nominally 112°C to 113°C or where the relative humidity is nominally 70 to 75 percent. The 
SCW test solution is a moderately concentrated alkaline solution and solutions in this 
concentration range could be expected to form for relative humidity in the range of 90 to 
95 percent. The SDW test solution is a dilute alkaline solution and solutions in this 
concentration range could be expected to form for high relative humidity (greater than 
99 percent). These may have characteristics of solutions at the drift wall, that is, typical of 
in-drift seepage waters. 
Under conditions of extreme evaporative concentration (i.e., low relative humidity) this type of 
brine containing high nitrate and chloride content would evolve into a nitrate-chloride brine with 
low carbonate content. The SSW test solution has characteristics of this type of brine. Likewise, 
the SAW test solution has characteristics of low carbonate brine and would have characteristics 
of solutions in equilibrium with relative humidity of nominally 90 percent. The calcium and 
magnesium addition to this test solution tends to make it more able to sustain lower pHs due to 
the hydrolysis properties of these cations. 
The sulfate brines have near-neutral pH and no significant bicarbonate/carbonate and calcium 
content. Calcium precipitates as carbonates and possibly sulfates. In addition, they typically 
have only a small amount of magnesium, though some surface brines have been observed to have 
high magnesium (Drever 1997, Table 15-1, p. 333, brines 1-3). The dominant cation is typically 
sodium. In the repository brines, potassium may be comparable to sodium, and magnesium is 
expected to be insignificant. A brine of this type may also evolve through a high extent of 
evaporation into one in which nitrate is the dominant anion. 
Relative humidity dependence of the sulfate brine composition is as follows. At low relative 
humidity, the aqueous solutions will be dominated by nitrate and chloride anions with nitrate 
ions dominating at the lowest relative humidity. At moderate relative humidity (greater than 
70 percent relative humidity), chloride ions could dominate the solution composition. However, 
unlike the carbonate brines, these brines are expected to have near-neutral to slightly acidic pH 
because of the lack of a carbonate component. Significant amounts of carbonate and sulfate ion 
are not expected until the relative humidity is greater than 85 percent because of the increase in 
solubility of expected sulfate minerals (sodium and potassium sulfates). (Magnesium sulfate is 
expected to be present in insignificant quantities in these brines.) 
The corrosion test solutions corresponding to the sulfate type of brine include the SAW and 
SSW. This type of brine has near-neutral to slightly acidic pH and, as noted, magnesium is not 
expected to be present in seepage waters to any significant extent. The SAW test solution has 
characteristics of solutions in equilibrium with nominally 90 percent relative humidity. The 
SSW has characteristics of water that have undergone evaporative concentration to the extent 
that sulfate precipitates out of solution (this is for the magnesium-free situation). 
Two important general factors specific to the repository brines are the presence of nitrate and 
more effective mechanisms for the removal of magnesium. It is expected that nitrate will always 
be present in the deliquescent brines because of multiple potential sources (BSC 2003g, 
Section 6.7.2.8) and the generally high solubility of nitrate minerals (BSC 2003g, 
Section 4.1.1.7). The presence of nitrate is evident in the endpoint brines, as shown in Table 2-1. 
December 2003 2-6 No. 6: Waste Package and Drip Shield Corrosion
It is expected that magnesium ions will never be significant because of a combination of low 
concentration (for the dust, as well as for at least some groundwaters) and multiple removal 
mechanisms, most of which are enhanced by elevated temperature (BSC 2003g, 
Sections 6.7.2.10 and 6.7.2.11). 
Table 2-1. Drift Crown Seepage Water Limiting Compositions and Probabilities of Their Formation, the 
Associated Brine Type, and the Corresponding Corrosion Test Solutions 
Dominant 
Constituents in 
Bin Water at 
98% Relative 
Humidity* 
Probability of 
Crown 
Seepage* Bin Water* 
Ca-Cl 0.00 1 
Na-Cl 0.00 2 
Na-Cl 0.22 3 
Na-Cl 1.42 4 
Na-Cl 0.79 5 
Na-Cl 6 5.46 
Na-Cl 27.15 7 
16.2 8 Na-CO3 
15.55 9 Na-CO3 
11.7 10 Na-CO3 
Na-Cl 21.5 11 
Source: BSC 2003f; Table 6.14-8 for columns identified with *. 
NOTE: The probability of crown seepage represents the 20,000-year time-integrated occurrence fraction (in 
percent) of the representative water for each bin. However, the frequency of occurrence can be 
significantly higher for short durations. 
Dominant 
Constituents in 
Endpoint 
Brines* 
Ca-Cl 
Ca-Cl 
Ca-Cl; K-Cl 
K-NO3; Na-NO3 
Na-Cl; K-Cl 
Na-Cl; Na-NO3; 
K-Cl 
Na-Cl; Na-NO3; 
K-Cl 
Na-Cl; Na-NO3; 
K-Cl 
K-NO3; K-Cl 
Na-Cl; Na-NO3; 
K-Cl 
Na-Cl; Na-NO3; 
K-Cl 
Brine Type 
Calcium chloride 
Calcium chloride 
Calcium chloride 
Sulfate 
Sulfate 
Carbonate 
Carbonate 
Carbonate 
Carbonate 
Carbonate 
Carbonate 
Corrosion Test 
Solution 
CaCl2; CaCl2 + 
Ca(NO3)2 
CaCl2; CaCl2 + 
Ca(NO3)2 
CaCl2; CaCl2 + 
Ca(NO3)2 
SSW, SAW, 
NaCl 
SSW, SAW, 
NaCl 
SDW, SCW, 
BSW, SSW, 
NaCl 
SDW, SCW, 
BSW, SSW, 
NaCl 
SDW, SCW, 
BSW, SSW, 
NaCl 
SDW, SCW, 
BSW, SSW, 
NaCl 
SDW, SCW, 
BSW, SSW, 
NaCl 
SDW, SCW, 
BSW, SSW, 
NaCl 
Aqueous corrosion test solutions include several multiionic solutions (Table 2-2) based on a 
carbonate-base, J-13 well water and test solutions containing the major species expected to effect 
corrosion. The standardized solutions developed by the project as relevant test environments are 
presented in Table 2-2. These solutions include SDW, SCW, and SAW at 30°C, 60°C, and 
90°C, as well as SSW at 100°C and 120°C. The SSW formulation is based upon the assumption 
No. 6: Waste Package and Drip Shield Corrosion 2-7 
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December 2003
Revision 1 
that evaporation of J-13 eventually leads to a sodium-potassium-chloride-nitrate solution. The 
absence of sulfate and carbonate in the target composition for this test medium is conservative. 
For example, carbonate would help buffer pH in any occluded geometry such as a crevice, and 
sulfate can act as a corrosion inhibitor. The compositions of these environments, as well as the 
solution known as BSW, are given in Table 2-2. Small amounts of carbonate will form in the 
SSW, SAW, and BSW solutions by interaction with gas phase carbon dioxide. The amount of 
carbonate formed was not determined experimentally, since the small amounts were not expected 
to affect the corrosion processes significantly. 
Ion 
K+ 3.400 × 103 3.400 × 101 
Na+ 4.090 × 104 4.090 × 102 
Mg2+ 
<1 Ca2+ 5.000 × 10-1 
F- 1.400 × 103 1.400 × 101 
Cl- 6.700 × 103 6.700 × 101 
NO3
- 6.400 × 103 6.400 × 101 
SO4
2- 1.670 × 104 1.670 × 102
1 <1 1.000 × 103 
1.000 × 103 
0 0 1.470 × 103 
2.425 × 104 
2.30 × 104 
3.86 × 104
0 0 0 HCO3
- 7.000 × 104 9.470 × 102 
Si 
Table 2-2. Target Composition of Standard Test Media Based on J-13 Well Water 
SSW 
(mg/L) 
SAW 
(mg/L) 
SCW 
(mg/L) 
SDW 
(mg/L) 
1.420 × 105 3.400 × 103 
4.870 × 104 3.769 × 104 
1.280 × 105 
1.313 × 106 
0 
5.5 to 7 
27 (60°C), 49 (90°C) 27 (60°C), 49 (90°C) 27 (60°C), 49 (90°C) 
2.7 9.8 to 10.2 9.8 to 10.2 pH 
Source: DTN: LL000320405924.146. 
BSW-12 
(mg/L) 
6.762 × 104 
1.0584 × 105 
0 0 
0 0 
1.3083 × 105 
1.3965 × 106 
1.470 × 104 
0 0
12 
December 2003 
NOTE: pH measured for actual solutions at room temperature. 
BSW can have a pH between 11 and 13, and has a boiling point near 110°C (BSW-12 with a pH 
of 12 shown in Table 2-2). This test medium was established based on results from a distillation 
experiment. The total concentration of dissolved salts in the starting liquid was more 
concentrated than that in the standard SCW solution. After evaporation of approximately 
90 percent of the water from the starting solution, the residual solution reaches a maximum 
chloride concentration and has a boiling point of approximately 110°C, with a pH of about 11. 
The synthetic BSW solution composition can be slightly modified (mainly by adding sodium 
hydroxide) to cover a range of pH values, yielding BSW-13, BSW-12, and BSW-11. 
Deliquescence of dust deposited on the waste packages and drip shield is another means by 
which brines can form on these engineered barrier system components. In the absence of salts, 
condensed water can be present on smooth surfaces only if the relative humidity is 100 percent. 
At lower relative humidity values, most of the water evaporates, with residual water existing on 
the surface as a very thin adsorbate layer. Dissolved salts lower the relative humidity at which 
such dryout occurs. Salt minerals in a dry system lower the relative humidity required for an 
aqueous solution to form. If the dissolved salt composition of a solution is known, the relative 
humidity at which dryout occurs at a given temperature can be determined. Conversely, the 
relative humidity for a given salt or set of salt minerals at which deliquescence occurs at a 
specified temperature can also be found. 
2-8 No. 6: Waste Package and Drip Shield Corrosion
Table 2-3 lists the brines that would develop on the waste packages using the analysis (BSC 
2003f) for dust deliquescence in drift thermal environments up to 140°C. Included in the table 
are the associated brine type and the corresponding aqueous corrosion test solutions. 
17.31 3 Na-SO4 
23.08 4 Na-NO3 
44.23 5 Na-NO3 
Table 2-3. Brines from Dust Deposited on the Waste Packages, Including Bromide, the Probabilities of 
Their Formation, the Associated Brine Type, and the Corresponding Corrosion Test Solutions 
Probability of 
Deliquescence 
Dominant 
Constituents in 
Bin Water at 
98 Percent 
Relative 
Humidity 
Bin 
Water Bin 
Dominant 
Constituents in 
Endpoint Brines 
Ca-NO3 1 5.77 Na-NO3 
Brine Type 
Calcium chloride 
Sulfate 
Carbonate 
Sulfate 
Sulfate 
Carbonate 
Corrosion Test 
Solution 
CaCl2; CaCl2 + 
Ca(NO3)2 
SSW, SAW, NaCl 
SDW, SCW, BSW, 
SSW, NaCl 
SSW, SAW, NaCl 
SSW, SAW, NaCl 
SDW, SCW, BSW, 
SSW, NaCl 
2 7.69 Na-NO3 
K-NO3; Na-NO3 
K-NO3; Na-NO3 
Na-NO3 
K-NO3; Na-NO3 
K-NO3 Na-Cl 1.92 6 
Source: BSC 2003f, Tables 6.10-6 and 6.14-9. 
NOTE: The probabilities represent the percentage of waste packages subject to dust that may deliquesce into the 
2-9 December 2003 
identified brine type. 
In all cases, the nitrate component is the most soluble species and would dominate the solution 
composition at the deliquescent relative humidity or eutectic point of a mineral assemblage at 
elevated temperatures. At higher relative humidity, chloride minerals become soluble and could 
become a dominant ion. It is not until the relative humidity is much higher that the sulfate and 
carbonate compositions could become appreciable. In essence, solutions will be dominated by 
chloride and nitrate at low to moderate (less than 70 percent) relative humidity and, at higher 
relative humidity, sulfate and carbonate could be appreciable. This is discussed in more detail in 
Environment on the Surfaces of the Drip Shield and Waste Package Outer Barrier (BSC 2003g). 
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2-10 No. 6: Waste Package and Drip Shield Corrosion 
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3. LONG-TERM THERMAL AGING OF ALLOY 22 IN THE REPOSITORY 
As described in Section 1.2, Alloy 22 phase stability issues resulting from exposure at higher 
temperatures during welding can potentially degrade corrosion resistance and mechanical 
properties. To minimize potential phase instabilities resulting from waste package fabrication, a 
solution heat treatment and rapid water quench of the as-fabricated waste package is 
implemented. Following solution heat treatment in the fabrication shop and the final thermal 
cycle for the lid closure weld, the temperature of the Alloy 22 waste package outer shell will 
remain significantly below 200°C as shown in Figure 1-2. With these constraints, the impact of 
thermal aging and phase instability on the corrosion of Alloy 22 are expected to be minimal 
consistent with early work (CRWMS M&O 2000b). This previous work indicated that 
degradation resulting from phase instability for Alloy 22 base metal will not affect waste 
package performance at less than about 300°C. 
3.1 MODEL DEVELOPMENT 
More recently, the precipitation kinetics for two of the higher temperature tetrahedrally close 
packed phases known as P and ó and the lower temperature ordered phase known as oP6-ordered 
phase have been modeled based on fundamental thermodynamics and kinetic concepts and 
principles (BSC 2003d). This modeling, benchmarked with thermal aging kinetic measurements, 
is being used to provide predictive insight into the long-term metallurgical stability of Alloy 22 
under repository relevant conditions. 
3.2 KINETICS OF TRANSFORMATION IN A NICKEL.CHROMIUM. 
MOLYBDENUM ALLOY SURROGATE FOR ALLOY 22 
To simulate the kinetic transformations of Alloy 22 over long time periods, a more simplified 
ternary nickel-chromium-molybdenum alloy that can be considered as a surrogate for Alloy 22 
was used. The nominal composition used was 21.1 weight percent chromium, 
13.5 weight percent molybdenum, and the remainder nickel. Three transformations were 
considered: the face-centered cubic matrix to the oP6-ordered phase, the face-centered cubic 
matrix to the P phase, and the face-centered cubic matrix to the ó phase. The corresponding 
thermodynamic-model-generated phase-fraction versus temperature diagrams are shown in 
Figure 3-1. In this study of phase-formation kinetics for the oP6-ordered phase, 
7.5 weight percent molybdenum is used instead of 13.5 weight percent in the extreme left panel. 
This reduction in molybdenum accounts for the role of tungsten in destabilizing the ordered 
phase, and the relative contribution of tungsten is approximately twice that of molybdenum 
(1 weight percent tungsten is approximately 2 weight percent molybdenum). 
In Figure 3-2, the time-temperature-transformation diagram associated with the face-centered 
cubic-to-oP6 ordered phase transformation of Ni.21.1Cr.7.5Mo is displayed with 2, 10, and 
15 percent transformation rates. 
Similarly, the transformation of Ni.21.1Cr.13.5Mo from the face-centered cubic-based matrix 
to the P phase with 2, 10, and 15 percent transformation rates was predicted (Figure 3-3) as a 
function of time and temperature. The results are comparable with the qualitative observations 
cited by Turchi (2001). Finally, the transformation of Ni.21.1Cr.13.5Mo from the 
December 2003 3-1 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
face-centered cubic-based matrix to the ó phase with 2, 5, and 10 percent transformation rates 
was predicted as a function of time and temperature (Figure 3-4). 
It can be seen from Figures 3-2, 3-3, and 3-4 that extrapolation of each curve to lower 
temperatures is consistent with the expected low probability of forming less than 2 percent (an 
insignificant amount) of these phases out of the face-centered cubic-solid solution at less than 
200°C for as long as 10,000 years. Thus, phase and time temperature-transformation diagrams 
predicted for Alloy 22, and validated with experimental data (BSC 2003d), indicate no 
significant phase instabilities (long-range ordering and tetrahedrally close packed phase 
precipitation) at temperatures below 200°C for 10,000 years. Further, examination of 
Figures 3-2, 3-3 and 3-4 indicate that no significant phase formation is expected even for the 
highest potential waste package temperatures of about 300°C resulting from early drift collapse. 
The estimated temperature increase due to drift collapse is based on the previously analyzed back 
fill case (with Overton sand) (BSC 2001a, Section 6.11.1.4 and Figures 6-7 and 6-14). This 
analysis suggests that the waste package temperature could increase to peak temperatures of 
about 300°C, but for only decades to centuries. 
Source: DTN: LL030106312251.013. 
NOTE: The face-centered cubic matrix and the oP6-ordered phase (left panel); in this case, 7.5 weight percent 
molybdenum, was considered instead of 13.5 weight percent (see text) P phase (middle panel) and ó phase 
(right panel). All other phases are suspended during the calculations. 
Figure 3-1. Time-Temperature-Transformation Diagram for Ternary Alloy Nickel.21.1 Chromium.13.5 
December 2003 3-2 
Molybdenum 
No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
Source: DTN: LL030106312251.013. 
NOTE: At 596°C (see Figure 3-1, left panel), the phase fraction of oP6-ordered phase drops to 0. 
Figure 3-2. Calculated Isothermal Time-Temperature-Transformation for a Face-Centered 
Cubic-Based Matrix of a Ternary Ni-21.1Cr-7.5Mo (in Weight Percent) Alloy (Surrogate for 
Alloy 22) Transforming into the oP6-Ordered Phase for 2, 10, and 15 Percent 
Transformation Rates 
December 2003 3-3 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
Source: DTN: LL030106312251.013. 
NOTE: At 836°C (see Figure 3-1, middle panel), the phase fraction of P phase drops to 0. 
Figure 3-3. Calculated Isothermal Time-Temperature-Transformation for a Face-Centered 
Cubic-Based Matrix of a Ternary Ni-21.1Cr-13.5Mo (in Weight Percent) Alloy (Surrogate 
for Alloy 22) Transforming into the P Phase for 2, 10, and 15 Percent Transformation Rates 
December 2003 3-4 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
Source: DTN: LL030106312251.013. 
NOTE: At 727°C (see Figure 3-1, right panel), the phase fraction of ó phase drops to zero. 
Figure 3-4. Calculated Isothermal Time-Temperature-Transformation for a Face-Centered 
Cubic-Based Matrix of a Ternary Ni.21.1Cr.13.5Mo (in Weight Percent) Alloy (Surrogate 
for Alloy 22) Transforming into the ó Phase for 2, 5, and 10 Percent Transformation Rates 
3.3 EXPERIMENTAL RESULTS 
Volume-Fraction Measurements in Alloy 22–In order to measure the amount of tetrahedrally 
close packed phase precipitation in Alloy 22 base metal and welds, area-fraction measurements 
have been made, using scanning electron microscope image analyses as a function of aging time 
and temperature. Using standard methods of quantitative stereology, it has been shown that 
area-fraction measurements (and correspondingly, linear-fraction measurements on grain 
boundaries) are mathematically equivalent to volume-fraction measurements (Vander Voort 
2000; Hilliard and Cahn 1961). Thus, the area-fraction measurements or bulk precipitation 
presented here are equivalent to the volume-fraction values in Alloy 22 as a function of time and 
temperature which in turn can be related to any potential effects on corrosion and (or) 
mechanical property degradation. The scanning electron microscope image analyses that were 
performed measured gross tetrahedrally close packed phase precipitation in the samples as a 
function of time and temperature. No tetrahedrally close packed phase extraction was conducted 
to differentiate the types of tetrahedrally close packed phases that may be present in these 
measurements. 
Volume-Fraction Measurements in Base Metal–The measurements of area-fraction of 
tetrahedrally close packed precipitation for base metal, as a function of time and temperature, are 
December 2003 3-5 No. 6: Waste Package and Drip Shield Corrosion
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presented in Figure 3-5. A trend line is included for the results at 760°C, where more than 
two measurements were made. Tetrahedrally close packed phases are seen to readily form at 
higher temperatures (760°C to 800°C) in less than 1,000 hours. In general, as the temperature is 
decreased, the onset of tetrahedrally close packed phase precipitation is delayed, and it also 
appears that the slopes of trend lines at lower temperatures may become shallower, indicating a 
slower rate of phase precipitation. The measurements presented here are reasonably consistent 
with model predictions shown in Figure 3-3 which in turn indicate that forming tetrahedrally 
close packed phases from the face-centered cubic solid solution within 10,000 years at less than 
200°C is unlikely. 
Source: DTN: LL030606912251.020. 
Figure 3-5. Precipitation of Tetrahedrally Close Packed Phases in Alloy 22 Base Metal as a Function of 
December 2003 3-6 
Time and Temperature 
Area-Coverage Measurements on Grain Boundaries–Area-coverage measurements 
(linear-fraction measurements) on grain boundaries have also been performed using scanning 
electron microscope image analyses, and are shown in Figure 3-6. These measurements are 
similar to the area-fraction measurements on base metal samples discussed in the previous 
section and give a quantitative measure of grain boundary precipitation kinetics. Extrapolation 
of Arrhenius plots or comparison to time-temperature-transformation diagrams in Figure 3-3 
indicate that significant precipitation will not occur under repository conditions (BSC 2003d). 
No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
Source: DTN: LL030606912251.020. 
Figure 3-6. Precipitation of Tetrahedrally Close Packed Phases at Alloy 22 Grain Boundaries as a 
December 2003 3-7 
Function of Time and Temperature 
Long-Range Ordering–The kinetics of long-range ordering are treated in a manner similar to 
that discussed for tetrahedrally close packed phase precipitation. However, very little kinetic 
data exist for long-range ordering in Alloy 22. A very fine dispersion of ordered domains was 
seen after aging for 30,000 and 40,000 hours at 427°C in Alloy 22 base metal, and in a weld 
similarly aged (BSC 2003d). The ordering in these cases is so fine that it would be difficult to 
measure the volume fraction of the ordered domains. Long-range ordering was also observed in 
Alloy 22 base metal aged at 593°C for 16,000 hours and at 538°C and 593°C for 1,000 hours. 
Alloy 22 base metal samples aged for 40,000 hours at 260°C and 343°C and for 1,000 hours at 
482°C were also examined with transmission electron microscopy, but no long-range ordering 
was observed. 
Unlike tetrahedrally close packed phase precipitates, long-range ordering results in very small 
and finely dispersed precipitates. As a result, scanning electron microscope image analysis is not 
well suited to determine the extent of long-range ordering kinetics. However, due to the 
uniformly and finely dispersed nature of long-range ordering, microhardness measurements are 
indicative of long-range ordering, because analogous to precipitation in age-hardened alloys, 
hardness increases with the amount of long-range ordering precipitation (Reed-Hill 1973). 
Figure 3-7 (BSC 2003d, Figure 96), shows such microhardness (Hv) measurements made on 
Alloy 22 as a function of time and temperature. Trend lines have been included for the results at 
500°C and 550°C. The microhardness of “as-received” material was 217 Hv. The 
microhardness measurements indicate that long-range ordering has occurred at temperatures in 
the range of 500°C to 550°C, for up to 40,000 hours. In addition, for results up to 40,000 hours, 
no long-range ordering is evident for temperatures below 400°C, and little long-range ordering is 
seen at temperatures around 600°C. The observation of very little long-range ordering near 
600°C conforms to the critical order-disorder temperature of the computational model, which is 
about 596°C (BSC 2003d). 
No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
Source: DTN: LL030607112251.021. 
NOTE: Microhardness of “as-received” base metal is 217 Hv. 
Figure 3-7. Microhardness (Hv) Measurements on Aged Alloy 22 Base Metal Shown as a Function of 
December 2003 
Time and Temperature and Indicative of Long-Range Ordering 
Volume Fraction Measurements in Alloy 22 Welds–As observed by Cieslak et al. (1986), 
tetrahedrally close packed phases are present in the interdendritic regions of the as-welded 
structure. After aging, the amount and size of tetrahedrally close packed precipitates increases 
with both time and temperature up to 760°C. Nucleation of precipitation was also observed to 
form possibly along grain boundaries in some areas of these samples. The area fraction of 
precipitates is shown as a function of time in Figure 3-8. Each of the data points in Figure 3-8 
represents the average of 20 to 40 measurements. In the as-welded condition, there is 
approximately 0.02 volume-percent tetrahedrally close packed phase. The average activation 
energy calculated from the slopes of Arrhenius plots of these data is 241 kJ/mol, which is 
comparable to but lower than the values of 250 and 260 kJ/mol reported previously for base 
metal (Rebak et al. 2000). Extrapolations of these data do not indicate that precipitate nucleation 
and growth in the welds will occur at temperatures below approximately 200°C. 
3-8 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
Source: DTN: LL030606912251.020. 
Figure 3-8. Tetrahedrally Close Packed Phase Precipitation Kinetics for Alloy 22 Gas Tungsten Arc 
December 2003 3-9 
Weld as a Function of Temperature 
3.4 MODEL CONFIDENCE 
In Aging and Phase Stability of Waste Package Outer Barrier (BSC 2003d, Section 7.5), the 
computational phase kinetics results for the phase (from DICTRA, a code that predicts 
time-temperature-transformation diagram) for the ternary nickel-chromium-molybdenum alloy (a 
surrogate for Alloy 22) were compared with volume-fraction measurements on Alloy 22 base 
metal at temperatures of approximately 700°C and 760°C. The computation phase kinetics 
results are used to construct the computational time-temperature-transformation diagrams for a 
particular phase. The comparison is somewhat limited because (1) the volume-fraction 
measurements do not distinguish among the possible tetrahedrally close packed phases (P, µ, and 
ó) that have formed, and (2) the DICTRA results only display one type of phase fraction at a 
time. As a result, the P phase formation results from DICTRA were compared with the 
tetrahedrally close packed volume-fraction measurements, as the P phase is the most likely to 
form at these temperatures and times. For this comparison, only a few volume-fraction 
measurements were deemed most likely to contain primarily P phase at the times and 
temperature shown and thus the available data set was very limited. 
However, in both comparisons (i.e., at 700°C and 760°C), the agreement between the 
computational kinetics results and the volume-fraction measurements is reasonable. The 
comparison (although based on a limited number of data points) shows that the computational 
results are conservative compared to the measured data. 
No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
3.5 SUMMARY 
Analyses documented in Aging and Phase Stability of Waste Package Outer Barrier (BSC 
2003d) assumed that the precipitation mechanisms that operate at higher temperatures also 
operated at much lower temperatures, and that the phases seen at the higher temperatures were 
also stable at the lower temperatures. Information from josephinite, a material that has been 
proposed as a natural analog for Alloy 22, shows stability of metallic phases after exposure over 
millions of years. This is fully consistent with model predictions and experimental observations 
of no low-temperature mechanism with rates significantly greater than those predicted at lower 
temperatures. This observation provides confidence in the implicit assumption that the 
high-temperature mechanisms used to extrapolate kinetics are the same as those that occur at 
lower temperatures (approaching expected repository conditions). 
Long-range ordering formation occurs at relatively lower temperatures than for the tetrahedrally 
close packed phases. However, the associated kinetics based on model predictions and 
transmission electron microscopy and microhardness measurements support the expected lack of 
formation of the ordered phase of Ni2Cr-type at repository temperatures, as the phase formation 
kinetics are primarily driven by thermally activated diffusion. Thus, alloys homogenized (or 
annealed) at high temperatures and rapidly quenched should not display any deleterious phases. 
Extrapolation of computationally derived time-temperature-transformation curves to lower 
temperatures, which are expected in a repository, indicate that formation of the P phase or 
oP6-ordered phase from the face-centered cubic solid solution will not occur. 
In summary, model predictions and extrapolation of higher temperature results to lower 
temperatures show that formation of tetrahedrally close packed or ordered phases in Alloy 22 
base metal and annealed welds will not occur during the repository period of 10,000 years. 
Further, as discussed, analysis shows that tetrahedrally close packed and long-range ordering 
phases would not form even for the unlikely case of early steady state drift collapse (acting like 
backfill), where the waste package temperature could potentially increase to peak temperatures 
of about 300°C, but for only relatively short times, decades to centuries. On this basis, neither 
the waste package outer shell base metal nor weld metal are subject to enhanced degradation due 
to the effects of thermal aging (BSC 2003b, Section 6.4.6), and this process is not included in 
total system performance assessment (TSPA) modeling. 
December 2003 3-10 No. 6: Waste Package and Drip Shield Corrosion
4. WASTE PACKAGE OUTER SHELL OXIDATION AND CORROSION 
IN DRYOUT REGIME 
While in the higher temperature dryout regime, Alloy 22 may undergo dry oxidation at relative 
humidities below the critical value for humid air or aqueous corrosion (CRWMS M&O 2000a, 
Section 6.4.2). However, the extent of dry oxidation is believed to be insignificant. To confirm 
the expected low dry oxidation rates at repository temperatures, a bounding calculation using 
typical literature based oxidation rate laws has been performed. 
4.1 DRY OXIDATION ABOVE DELIQUESCENCE POINT 
Dry oxidation of Alloy 22 at repository temperatures involves the formation of a relatively thin 
protective oxide film. To quantify this effect, a simple bounding analysis has been performed 
that assumes the uniform protective surface oxide film is primarily Cr2O3 (other oxides may also 
be present). The rate of dry oxidation is assumed to be limited by mass transport through this 
growing metal oxide film. Fick¡¯s First Law is applied, assuming a linear concentration gradient 
across the oxide film of thickness x. Integration shows that the oxide thickness should obey the 
following parabolic growth law, also known as Wagner¡¯s Law (Welsch et al. 1996), where the 
film thickness is proportional to the square root of time. This is represented by the following 
equation. 
k ¡¿ t 2
0 
where 
x + 
0 x2 is the initial oxide thickness, x is the oxide thickness at time t, and k is a 
temperature-dependent parabolic rate constant that follows an Arrhenius type relationship. The 
rate log term constant in the above equation is determined as follows: 
x = 
1 103 
2 log ( ) K T . .. 
. 
where T is defined as the absolute temperature. The highest waste package surface temperature 
is expected to be less than 200¡ÆC, which is well below 350¡ÆC (623 K), the boundary temperature 
used in calculations below (BSC 2003b, Section 6.4.2). The value of k corresponding to this 
very conservative upper limit of 350¡ÆC is 2.73 ¡¿ 10.24 m2/s (8.61 ¡¿ 10.5 ¥ìm2/yr). After 1 year, 
this corresponds to an oxide thickness of 0.0093 ¥ìm (about 9.3 nm/yr). As will be seen in 
subsequent discussion, this estimated rate is comparable to that expected for aqueous phase 
corrosion at lower temperatures. A logarithmic growth law may be more appropriate for use at 
low temperatures than the parabolic growth law used above. However, such a logarithmic 
expression predicts that the oxide thickness (penetration) will asymptotically approach a small 
maximum level. The parabolic growth law predicts continuous growth of the oxide, which is 
much more conservative than the logarithmic growth law. To be conservative, the parabolic 
growth law is used to model the dry oxidation of Alloy 22. Recent measurements (BSC 2003b, 
Section 6.4.2) of the thickness of the Alloy 22 oxide film exposed to air at 550¡ÆC showed the 
oxide film approaches a limiting thickness of about 0.025 to 0.050 ¥ìm after about 333 days 
exposure, which corresponds to a penetration rate of 0.027 to 0.054 ¥ìm/yr calculated at 350¡ÆC 
( ) 12.5 [ ] k m s. = .
4-1 No. 6: Waste Package and Drip Shield Corrosion 
Revision 1 
(Eq. 4-1) 
. 5 . 3 (Eq. 4-2) ÿ ÿ. 
. 
December 2003
Revision 1 
(BSC 2003b). Therefore, the upper bound conservative rate of 0.0093 µm/yr is consistent with 
the measurement data considering the higher test temperature. This precludes the need for 
additional uncertainty bounds. Thus, this analysis indicates dry oxidation will not be a life 
limiting waste package outer shell degradation mechanism. On this basis, dry oxidation will not 
have a significant effect on the performance of the waste package outer shell, and this process is 
not included in TSPA modeling (BSC 2003b, Section 6.4.2). 
4.2 CORROSION UNDERNEATH DELIQUESCENT BRINES 
At a given surface temperature, the existence of liquid water on the waste package surface 
depends upon the hygroscopic nature of any salt and mineral deposits on the surface. In the 
presence of such deposits, a thin liquid-phase surface brine film can be established at a higher 
temperature and lower relative humidity than otherwise possible. The chemistry–temperature– 
relative humidity stability range of this liquid phase brine film is modeled. A summary of the 
results and expected ranges of deliquescent brine compositions and their relative probabilities of 
occurrence are summarized in the in-drift chemical environment technical basis document. 
Environmental thermogravimetric analysis has also been used to study the corrosion of waste 
package materials underneath deliquescent brines, and the evolution of acid gas due to the 
thermal decomposition of those brines (BSC 2003g, Section 6.7.2.1.4.2). The subsequent weight 
loss is due to chemical transformations that are occurring in the aqueous solutions due to 
volatilization of HCl. The accompanying pH increase causes precipitation of calcium containing 
species with the loss of the aqueous phase. Because of its lower corrosion resistance, Alloy 825 
was tested in parallel to provide insight into localized modes of attack that might occur under 
deliquescent brine films. Weight change data for the two materials (Alloy 825 and Alloy 22) at 
150°C and 22.5 percent relative humidity are shown in Figure 4-1. 
Initial weight gains shown in Figure 4-1 are due to the formation of highly concentrated aqueous 
films due to deliquescence of the deposited CaCl2. No sustained oxidation of Alloy 22, usually 
indicated by an increase in weight due to the addition of oxygen to the surface, is evident from 
the thermogravimetric analysis data. 
December 2003 4-2 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
Source: Farmer 2003, Slide 14. Data taken from Hailey and Gdowski 2003. 
Figure 4-1. Thermogravimetric Analysis Data Comparing the Weight Gains of Alloy 825 and Alloy 22 at 
2 
December 2003 
150°C and 22.5 Percent Relative Humidity 
Figure 4-2 shows photographs of the samples after such testing in the thermogravimetric 
analysis. It is evident that Alloy 22 is highly resistant to localized attack in a deliquescent CaCl 
brine at 150°C and 22.5 percent relative humidity. There is no evidence of localized corrosion of 
Alloy 22 (UNS N06022, 55.5 Ni 22 Cr 13 Mo 3 W 4 Fe 2.5 Co) due to deliquescence. The 
white spots visible are deposits. However, in contrast, substantial attack of a less corrosion 
resistant material, Alloy 825 (UNS N08825, 42Ni 22Cr 3Mo 0.9 Titanium 2.2Cu 1Mn 28.9Fe) is 
evident. Thus, Alloy 22 has been shown to be resistant to localized attack under aggressive 
CaCl2-type deliquescent brines at 150°C and 22.5 percent relative humidity. 
Figure 4-3 (BSC 2003g, Figure 38) shows a scanning electron micrograph of the white 
precipitates on the Alloy 22 surface, which are formed from the deliquescent brine and are 
evident in the photographs of Figure 4-2. Elemental analysis of the deposit was done with 
energy dispersive spectroscopy and indicated that these precipitates contain calcium, chlorine 
and oxygen. Raman spectroscopy indicates that precipitates are not Ca(OH)2 or CaCO3, but may 
be CaOHCl (calcium hydroxy chloride). Furthermore, energy dispersive spectroscopy and 
wet-chemical analyses indicate a loss of chlorine relative to calcium, which is believed to be due 
to the formation of volatile hydrochloric acid gas (BSC 2003g, Section 6.7.14-2). 
4-3 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
Source: Farmer 2003, Slide 15. Data taken from Hailey and Gdowski 2003. 
NOTE: No such localized corrosion of Alloy 22 was observed under identical conditions. 
Figure 4-2. Localized Corrosion of Alloy 825 Underneath a Deliquescent Brine 
Source: Farmer 2003, Slide 16. Data taken from Hailey and Gdowski 2003. 
NOTE: Elemental analysis with energy dispersive spectroscopy indicates that the precipitates are probably CaOHCl. 
Figure 4-3. Scanning Electron Microscope View of the White Precipitates on the Alloy 22 Surface 
(Shown in Figure 4-2) Formed from the Deliquescent Brine 
December 2003 4-4 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
5. GENERAL CORROSION OF THE WASTE PACKAGE OUTER SHELL 
General corrosion is the uniform thinning of the waste package outer shell at its open-circuit 
corrosion potential (Ecorr). The general corrosion rate is temperature dependent, and for a given 
temperature, is assumed to be constant (i.e., time-independent). Therefore, for a given 
temperature, the depth of penetration or thinning of the waste package outer shell by general 
corrosion is equal to the general corrosion rate at that temperature multiplied by the time that the 
waste package is exposed to the environment under which general corrosion occurs. This 
assumption is considered conservative because the general corrosion rate of metals and alloys 
tends to decrease with time. As discussed in the following sections, general corrosion rates of 
the waste package outer shell have been estimated based on weight-loss and dimensional-change 
measurements of descaled Alloy 22 samples after a 5-year exposure in the LTCTF. The LTCTF 
provides a comprehensive source of corrosion data for Alloy 22 in environments relevant to the 
repository (BSC 2003b, Section 6.4.3). In addition to LTCTF results, general corrosion rates 
were also measured electrochemically to help model the temperature dependence of this 
corrosion mode in Alloy 22. 
5.1 LONG-TERM WEIGHT LOSS MEASUREMENTS 
The LTCTF is equipped with an array of fiberglass tanks. Each tank has a total volume of 
approximately 2,000 L and is filled with approximately 1,000 L of aqueous test solution. The 
temperature of the solution in a particular tank is controlled at either 60°C or 90°C, covered with 
a blanket of air flowing at approximately 150 cm3/min, and agitated. Four generic types of 
samples, U-bends, crevices samples, weight loss samples, and galvanic couples, are mounted on 
insulating racks and placed in the tanks. Approximately half of the samples are submersed, half 
are in the saturated vapor above the aqueous phase, and a limited number are at the water line. It 
is important to note that condensed water is present on specimens located in the saturated vapor 
(BSC 2003b, Section 6.4.3.). 
The testing includes a wide range of plausible test media, including SDW, SCW, simulated 
cement-modified water, and SAW. The compositions of three of these solutions are summarized 
in Table 2-2. In addition, these data along with a detailed discussion of corrosion rate 
measurement and analysis results are presented in General Corrosion and Localized Corrosion 
of Waste Package Outer Barrier (BSC 2003b, Section 6.4.3). The corrosion rate of Alloy 22 
was determined according to applicable ASTM G 1-90 1999. Results from the two types of 
coupons were used. These were labeled weight loss coupons and crevice coupons (BSC 2003b, 
Section 6.4.3). Figures 5-1 and 5-2 show the corrosion rate determined on these two types of 
samples after a 5-year exposure period. 
The SCW test medium is about three orders of magnitude (1,000×) more concentrated than J-13 
well water and is alkaline (pH approximately 10). The SAW test medium is about three orders 
of magnitude (1,000×) more concentrated than J-13 well water and is acidic (pH approximately 
2.7). These concentrated solutions are intended to mimic the evaporative concentration of the 
various potential electrolytes on the hot waste package surface. Two temperature levels (60°C 
and 90°C) are included (BSC 2003b, Section 6.4.3). 
December 2003 5-1 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
After a 5-year exposure to each solution/environmental condition, specimens were removed from 
their respective test vessel to determine the corrosion rate by weight-loss and 
dimensional-change measurements. Some samples had been previously removed after 6 month, 
1-year and 2-year exposures. In all of the tested conditions, the coupons were covered with 
deposits. Therefore, the coupons were cleaned prior to final weighing. Cleaning was carried out 
using ASTM G 1. 
A detailed analysis of these results based on weight loss coupons exposed to these environments 
at 60°C and 90°C for over 5 years is reported in the model report on the general and localized 
corrosion of the Alloy 22 waste package outer shell (BSC 2003b, Section 6.4.3), and a summary 
of these Alloy 22 corrosion rate results is presented here. The average corrosion rates and 2ó 
ranges are presented for the uncreviced (designated as weight loss specimens) and the creviced 
specimens in Figures 5-1 and 5-2, respectively. 
The 2ó range represents a 95 percent confidence level. The individual corrosion rates for the 
weight loss coupons ranged from 0 to 12 nm/yr, with the lowest rates observed for the coupons 
in the SDW solution. The individual corrosion rates for the crevice coupons, shown in 
Figure 5-2, ranged from 0 to 23 nm/yr with the highest rates observed in the SAW solution and, 
again, the lowest rates observed in the SDW solution. In most cases, the crevice coupons 
exhibited corrosion rates two to five times higher than the weight loss coupons in the same 
solutions. Stereomicroscopic, scanning electron microscope and atomic force microscopy 
observations of both weight loss and crevice specimens indicated little or no corrosion for 
Alloy 22. The machining grooves remained uniform and sharp throughout each coupon. It is not 
yet clear why the corrosion rates of the crevice coupons were higher than those of the weight loss 
coupons because crevice corrosion was not observed in any of the tested coupons. However, it is 
noteworthy that among all test specimens, a maximum measured corrosion rate of only 23 nm/yr 
was observed (BSC 2003b, Section 6.4.3). 
For both the weight loss and crevice coupons, the corrosion rates were generally lower for those 
specimens exposed to vapor than immersed in liquid, regardless of the test temperature or 
electrolyte solution. For the weight loss coupons exposed to liquid, the corrosion rates were 
generally lower at 90°C than at 60°C. For the weight loss coupons exposed to vapor, the 
corrosion rates were generally higher at 90°C than at 60°C. Overall, coupons in the SAW 
solution exhibited slightly lower corrosion rates at the higher temperature. Similar to the weight 
loss coupons, the corrosion rates for the crevice coupons exposed to liquid were lower at 90°C 
than at 60°C, while the corrosion rates were generally higher at 90°C than at 60°C for the crevice 
coupons exposed to vapor. In general, for corrosion processes, the corrosion rate increases with 
temperature. However, since in this study the corrosion rates were so low and the temperature 
range studied (60°C to 90°C) is small, a clear dependence with the temperature cannot be 
established for any set of coupons. Finally, for the weight loss coupons, there appeared to be no 
effect of the presence of welds on the corrosion rate, however, the nonwelded crevice coupons 
exhibited slightly higher rates than their welded counterparts (BSC 2003b, Section 6.4.3). 
December 2003 5-2 No. 6: Waste Package and Drip Shield Corrosion
Source: Output DTN: SN0308T0506303.004. 
Figure 5-1. Corrosion Rates for Alloy 22 Weight-Loss Coupons in Simulated Acidified Water, Simulated 
Concentrated Water, and Simulated Dilute Water 
Source: Output DTN: SN0308T0506303.004. 
Figure 5-2. Corrosion Rates for Alloy 22 Crevice Coupons in Simulated Acidified Water, Simulated 
Concentrated Water, and Simulated Dilute Water 
No. 6: Waste Package and Drip Shield Corrosion 
Revision 1 
December 2003 5-3
Revision 1 
The empirical cumulative distribution functions for the general corrosion rates of Alloy 22 
weight-loss and crevice samples are presented in the model report on the general and localized 
corrosion of the waste package outer shell (BSC 2003b, Section 6.4.3.2). These provide a 
comparison of the effect of various experimental factors on the general corrosion rate, such as for 
the solution chemistry, temperature, and metallurgical condition. For the weight-loss samples, 
the mean corrosion rate is 2.75 nm/yr and the standard deviation is 2.74 nm/yr. For the crevice 
samples, the mean corrosion rate is 7.24 nm/yr, and the standard deviation is 4.95 nm/yr 
(Figure 5-3). 
Source: Output DTN: SN0308T0506303.004. 
Figure 5-3. Cumulative Distribution Function of Ro of Base-Case General Corrosion Rate for Alloy 22 
1 (Eq. 5-1) o 
C
T 
December 2003 5-4 
Waste Package Outer Shell at 60°C 
The crevice coupon rates were used as the base case general corrosion rate of the waste package 
outer shell. A Weibull distribution, with scale factor of 8.88, shape factor of 1.62, and location 
factor of 0, best fits the corrosion rate distribution (BSC 2003b, Section 6.4.3). 
5.2 GENERAL CORROSION MODEL 
The temperature dependence of general corrosion rate can be represented by the logarithm of the 
Arrhenius relationship. 
ln( T R ) = C + 
RT is the temperature-dependent general corrosion rate in nanometers per year, T is the absolute 
temperature in Kelvin, and C and C1 are constants. The temperature-dependence term (C1) is 
determined from short-term polarization resistance data for Alloy 22 specimens tested for a 
range of sample configurations, metallurgical conditions, and exposure conditions (temperature 
No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
and water chemistry). Figure 5-4 shows the temperature dependence of Alloy 22 corrosion rates 
measured by the polarization resistance technique over the temperature range from 45°C to 
170°C. From fitting the data to the Arrhenius relationship, the temperature-dependence term 
(C1) was found to obey a normal distribution with a mean of approximately 3,000 and a standard 
deviation of approximately 300. This temperature dependence is characterized by the activation 
energy of 26 kJ/mol. Sample configuration (crevice, disk, or rod), metallurgical conditions 
(mill-annealed or weld), and water chemistry within the range expected in the repository appear 
to have no significant effect on the temperature dependence of general corrosion rate. 
The Arrhenius relationship is used to predict the temperature-dependent general corrosion rate 
(RT) from the general corrosion rate at 60°C (R0) and the temperature-dependence term (C1). 
+ = ln( ( ln (Eq. 5-2) R ) R ) ) 1 o T 
1 
333.15 
C 
T
1 ( - 
Predictions of the distribution of the temperature-dependent general corrosion rate (RT) at 25°C, 
50°C, 75°C, 100°C, 125°C, and 150°C are shown in Figure 5-5 (BSC 2003b, Figure 6-23). 
Source: Output DTN: SN0308T0506303.004. 
NOTE: Corrosion rates were determined from 24-hour polarization resistance measurements. 
Figure 5-4. Temperature Dependence of Corrosion Rates for Alloy 22 for Various Metallurgical 
Conditions and Sample Configurations and in a Wide Range of Test Solutions
December 2003 5-5 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
Source: Output DTN: SN0308T0506303.004. 
NOTE: The calculation was performed using the mean value (.3116.47) of the temperature-dependency term (C1). 
The calculated general corrosion rate range represents the variability of the rate. 
Figure 5-5. Calculated Model Outputs of the Base-Case Temperature-Dependent General Corrosion 
Model, Based on the Crevice Sample Data at 25°C, 50°C, 75°C, 100°C, 125°C, and 150°C 
5.3 UNCERTAINTY ANALYSIS OF GENERAL CORROSION RATE DATA 
Most of the uncertainties in general corrosion rate of Alloy 22 have resulted from insufficient 
resolution of the weight-loss and dimensional-change measurements of the samples due to the 
extremely low corrosion rates of Alloy 22 in the test media. It was concluded that measurement 
uncertainty was the main source of uncertainty. The combined standard uncertainty is estimated 
to be approximately 0.185 nm/yr in the case of crevice samples and 0.314 nm/yr in the case of 
weight-loss samples (BSC 2003b, Section 6.4.3). These estimates correspond to 1 standard 
deviation (1 ó). Therefore, for the crevice samples, about 3 percent of the variation in the 
measured general corrosion rate is due to the measurement uncertainty, and 97 percent of it is 
from the variations of the corrosion rate among the specimens. For the weight loss samples, 
most of the variation (about 89 percent) in the measured corrosion rate is due to variations 
among the specimens, and the rest is from measurement uncertainty. 
5.4 TIME-DEPENDENT GENERAL CORROSION BEHAVIOR OF THE WASTE 
PACKAGE OUTER SHELL 
The general corrosion model implemented in TSPA assumes that general corrosion of the 
Alloy 22 waste package outer shell progresses uniformly over a large surface. The general 
December 2003 5-6 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
corrosion rate is temperature dependent; for a given temperature, it is assumed to be constant 
(i.e., time-independent). Therefore, for a given temperature, the depth of penetration or thinning 
of the waste package outer shell by general corrosion is equal to the general corrosion rate at that 
temperature multiplied by the time duration that the waste package is at that temperature. In 
general, however, the corrosion rates of metals and alloys decrease with time. This is shown in 
Figure 5-6 (BSC 2003b, Figure 6-24) for the mean general corrosion rates of Alloy 22 after 
0.5-, 1-, 2-, and 5-year exposures in the LTCTF and for shorter exposure time results based on 
other measurement techniques such as weight loss, polarization resistance and potentiostatic 
polarization tests. The mean general corrosion rate after a 5-year exposure in the LTCTF is 
0.007 µm/yr (7 nm/yr). The trend of decreasing general corrosion rate with time is consistent 
with the expected corrosion behavior of passive alloys such as Alloy 22 under repository-type 
aqueous conditions. The time-dependent general corrosion behavior of the waste package outer 
shell was not included in TSPA because the constant (time-independent) rate model is more 
conservative and should bound the general corrosion behavior of the waste package outer shell 
over the repository time period. Since the long-term rates will be lower than the 5-year rates, the 
5-year corrosion rates were conservatively selected for extrapolation over the repository time 
scale. 
Figure 5-6. Time-Dependent General Corrosion Rate of Alloy 22 
NOTE: Trend line was obtained by a linear regression fit for the data shown in the figure. 
5.5 OTHER FACTORS INFLUENCING GENERAL CORROSION OF WASTE 
PACKAGE OUTER SHELL MATERIAL 
Effect of Microbial Activity on General Corrosion–MIC is the contribution to the corrosion 
rate of a metal or alloy by the presence or activity, or both, of microorganisms (BSC 2003b, 
Section 6.4.3). MIC most often occurs due to the increase in anodic or cathodic reactions due to 
the direct impact of microorganisms on the alloy, or by indirect chemical effects on the 
surrounding solution. Microorganisms can affect the corrosion behavior of an alloy either by 
acting directly on the metal or through their metabolic products. For example, some types of 
5-7 No. 6: Waste Package and Drip Shield Corrosion December 2003
Revision 1 
aerobic bacteria may produce sulfuric acid by oxidizing reduced forms of sulfur (elemental, 
sulfide, sulfite), and certain fungi transform organic matter into organic acids (Fontana 1986, 
Section 8-10). 
It has been observed that nickel-based alloys such as Alloy 22 are relatively resistant to MIC 
(Lian et al. 1999). The impact of microbially influenced corrosion is accounted for by adjusting 
the rate of general corrosion. 
There are no standard tests designed specifically to investigate the susceptibility of an 
engineering alloy to MIC (Stoecker 1987). One commonly used type of evaluation to determine 
the MIC factor is to test the alloy of interest in situ (field) using the same variables as for the 
intended application. However, testing in the laboratory with live organisms can provide more 
controlled conditions of various environmental variables, and sterile controls can be incorporated 
to better assess MIC-specific effects (Horn and Jones 2002). This approach was used to evaluate 
the microbiological processes on general corrosion of the waste package outer shell. For general 
corrosion of the waste package outer shell, the effect of microbially influenced corrosion can be 
described as follows: 
(Eq. 5-3) CR · fMIC st CRMIC = 
where CRMIC is the general corrosion rate in the presence of microorganisms, CRst is the general 
corrosion rate of the alloy in the absence of MIC, and fMIC is the microbially influenced corrosion 
factor. If fMIC is greater than 1, there is an enhancement of the corrosion rate of the alloy as a 
consequence of the presence or activity of microorganisms. 
Lian et al. (1999) showed with polarization resistance measurements that MIC can enhance 
corrosion rates of Alloy 22 by a factor of at most two. Measurements for Alloy 22 and other 
similar materials are shown in Table 5-1 (BSC 2003b, Table 6-10). The microbially influenced 
corrosion factor fMIC is calculated as the ratio of corrosion rates (microbes to sterile) from the 
table. The value of fMIC for Alloy 22 in sterile media is set to 1 (fMIC = 1), whereas the value of 
fMIC for Alloy 22 in inoculated media is larger (fMIC = 2). It is assumed that the microbially 
influenced corrosion factor fMIC is uniformly distributed between 1 and 2, and that this 
distribution is all due to uncertainty. The MIC factor is applied to the waste package outer shell 
general corrosion rate when the relative humidity at the waste package outer shell surface is 
above 90 percent. This MIC initiation threshold relative humidity is based on the analysis 
documented in In-Drift Microbial Communities (CRWMS M&O 2000c, Sections 6.3.1.6 and 
6.5.2, Table 23), which determined that microbial activity would not occur at relative humidity 
less than 90 percent. 
Other environmental factors that could affect bacterial growth include temperature and radiation. 
These factors, however, are closely coupled to relative humidity; as temperature and radiation 
decrease in the repository, relative humidity is predicted to increase. At the same time, while 
there are some types of microorganisms that can survive elevated temperatures (less than or 
equal to 120°C) and high radiation doses, if there is no available water, then bacterial activity is 
completely prevented. Thus, because water availability is the primary limiting factor and this 
factor is coupled to other less critical limiting factors, water availability (as expressed by relative 
humidity) was used as the primary gauge of microbial activity. 
December 2003 5-8 No. 6: Waste Package and Drip Shield Corrosion
Microbes 
Sterile Stainless Steel Type 304 0.003 
Table 5-1. Alterations in Corrosion Rates and Potentials Associated with Microbial Degradation 
Tested Sample Initial Condition 
CS1020 + YM Microbes 
Sterile CS 1020 
M400 + YM Microbes 
Sterile M400 
C-22 + YM Microbes 
Sterile C-22 
I625 + YM Microbes 
Sterile I625 
Stainless Steel Type 304 + YM 0.035 
Corrosion Potential Ecorr (V versus SCE) 
Initial 
-0.660 
-0.500 
-0.415 
-0.135 
-0.440 
-0.260 
-0.440 
-0.160 
-0.540 
-0.145 
Average Corrosion 
Rate (µm/yr) 
8.80 
1.40 
1.02 
0.005 
0.022 
0.011 
0.013 
0.003 
5-9 
Endpoint 
-0.685 
-0.550 
-0.315 
-0.070 
-0.252 
-0.200 
-0.285 
-0.130 
-0.280 
-0.065 
December 2003 
Source: DTN: LL991203505924.094. 
Effect of Aging and Phase Stability on General Corrosion–The waste package outer shell 
base metal and all fabrication welds (not including the welds for the closure lids) are fully 
annealed before the waste packages are loaded with waste (BSC 2003a). The analysis 
documented in the model report titled Aging and Phase Stability of Waste Package Outer Barrier 
(BSC 2003d, Sections 6.6.5.3 and 8.0) has shown that phase instabilities are not expected in 
Alloy 22 base metal and welded material as long as the temperature remains below about 200°C. 
Further, as discussed in Section 3, analysis suggests that even for the case of early steady state 
drift collapse (acting like backfill), the waste package temperature could potentially increase by 
about 130°C, leading to peak temperatures of about 300°C but for only relatively short times 
(decades to centuries). 
Mechanical stress mitigation processes, currently planned for the closure weld, however, may 
introduce cold work into the material. This cold work might accelerate phase transformation 
kinetics, however the kinetics would have to be at least two orders of magnitude higher before it 
would be expected to be observed at low temperatures in 10,000 years. For a range of thermal 
loading designs of the repository, the waste package surface temperature will be always kept 
below 200°C (BSC 2001b, Section 5.4.1, Figures 5.4.1-2 and 5.4.1-6) in the absence of drift 
collapse. Although peak temperatures of about 300°C are possible under drift collapse, they 
would only exist for decades to centuries. These times are too short for significant aging effects 
at 300°C (Section 3). With this constraint, the effect of aging and phase instability on the 
corrosion of the waste package outer shell will therefore be insignificant in the repository. 
In order to analyze the effects of thermal aging on corrosion of Alloy 22, three metallurgical 
conditions of Alloy 22 were studied using the multiple crevice assembly samples: mill-annealed, 
as-welded, and as-welded plus thermally aged (at 700°C for 173 hours). The samples were 
tested, using electrochemical methods similar to those presented in Sections 6 and 7, in 5 M 
CaCl2 and 5 M CaCl2 + 0.5 M Ca(NO3)2 solutions at temperatures up to 120°C. After immersion 
in the test solution at open-circuit potential for 24 hours, the polarization resistance of the 
samples was measured. The corrosion rates from the polarization resistance measurements were 
No. 6: Waste Package and Drip Shield Corrosion 
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Revision 1 
only for comparative analysis of the effects of thermal aging on corrosion of Alloy 22; the tests 
were not intended to obtain the absolute values of the corrosion rate. It is to be noted that the 
corrosion rates derived from the short-term electrochemical testing are significantly higher that 
the rates obtained from long-term testing. This is characteristic of short-term tests and the 
results, therefore, are being used for comparative purposes only. 
Comparison of the calculated corrosion rates of the mill-annealed, as-welded, and as-welded plus 
thermally aged samples are shown in Figure 5-7 for 5 M CaCl2 solutions and Figure 5-8 for 5 M 
CaCl2 + 0.5 M Ca(NO3)2 solutions. The mill-annealed multiple crevice assembly samples in 5 M 
CaCl2 solutions at differing temperatures were considered as the baseline condition for the 
analysis. The baseline condition rates were compared with those of the as-welded and as-welded 
plus thermally aged multiple crevice assembly samples tested in the same electrolyte solution 
condition. A data trend-line for the baseline condition data was obtained by linear regression fit 
for an easier comparison with the as-welded and as-welded plus thermally aged sample data. 
The comparison shown in Figure 5-7 clearly shows that there is no apparent enhancement of the 
corrosion rate due to welding or thermal aging of the welded samples for the tested conditions. 
As shown in Figure 5-8, a similar comparison was made for the corrosion rates measured in 5 M 
CaCl2 + 0.5 M Ca(NO3)2 solutions. As for the 5 M CaCl2 solution case, the mill-annealed 
multiple crevice assembly samples at differing temperatures were considered as the baseline 
condition, and a data trend-line was drawn for the baseline condition data for an easier 
comparison. The comparison in Figure 5-8 again clearly shows no apparent enhancement of the 
corrosion rate due to welding or thermal aging of the welded samples. It is also noted that the 
corrosion rates of all three type samples (mill-annealed, as-welded, and as-welded plus thermally 
aged) were reduced by a factor of 3 to 4 in the nitrate containing solutions, compared to those in 
5 M CaCl2 solutions. 
The above analyses are consistent with the results by Brossia et al. (2001, Section 3.2.1.3, 
Figure 3-13) and Rebak et al. (2002). Comparison of the anodic passive current densities of the 
as-welded Alloy 22 samples to those of the base metal samples showed no significant effect of 
the welds on the passive corrosion behavior of the alloy (Brossia et al. 2001, Section 3.2.1.3, 
Figure 3-13). 
Based on the above analysis and insignificant aging and phase stability processes under the 
thermal conditions expected in the repository (BSC 2003d, Sections 6.6.5.3 and 8.0), the 
corrosion performance of the waste package outer shell is not expected to be affected by the 
aging and phase stability in the repository. Hence thermal aging and phase instability effects of 
the waste package outer shell on corrosion are not included in TSPA. 
Similarly, analyses described in Section 8 indicate the effects of radiolysis on corrosion 
performance of Alloy 22 will not be significant enough to lead to corrosion-induced failure of 
the waste package outer shell under repository relevant conditions. For this reason, the effects of 
radiolysis on corrosion performance of Alloy 22 are not included in TSPA modeling.
December 2003 5-10 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
Source: Output DTN: SN0308T0506303.003. 
NOTE: The trend line for the mill-annealed samples was obtained by linear regression fit of the data set. 
Figure 5-7. Comparison of Corrosion Rates from 24-Hour Polarization Resistance Measurements of 
Samples in 5 M CaCl 
Mill-Annealed, As-Welded, and As-Welded Plus Aged Alloy 22 Multiple Crevice Assembly 
2 Brines at Varying Temperatures 
December 2003 5-11 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
Source: Output DTN: SN0308T0506303.003. 
Figure 5-8. Comparison of Corrosion Rates from 24-Hour Polarization Resistance Measurements of 
Samples in 5 M CaCl 
Mill-Annealed, As-Welded, and As-Welded Plus Aged Alloy 22 Multiple Crevice Assembly 
2 + 0.5 M Ca(NO3)2 Brines at Varying Temperatures 
5.6 SUMMARY OF GENERAL CORROSION OF WASTE PACKAGE OUTER 
SHELL 
Because of extremely slow corrosion rates of Alloy 22, there is little data for Alloy 22 
in the scientific literature that could be used to evaluate the general corrosion model 
presented above. However, similar passive corrosion behavior has also been observed for 
nickel-chromium-molybdenum type corrosion-resistant alloys. For example, Alloy C is found to 
retain a very thin passive film, indicated by the retained mirror-like finish after 44 years of 
exposure at Kure Beach to a marine environment (i.e., salt air with alternate wetting and drying 
as well as the presence of surface deposits) (Baker 1988, p. 134, Table 6). More recent 
examination of specimens from this alloy after more than 50 years of exposure indicates that the 
samples continue to maintain a mirror-like finish and passive film behavior (McCright 1998, 
Figure ES-1). Under these exposure conditions, the less corrosion-resistant Alloy 600 exhibited 
a corrosion rate of 8 nm/yr after 36 years of exposure. This long-term corrosion rate is consistent 
with the model prediction. The 50th and 99.99th percentile rates at 25°C predicted by the model 
are 2.4 and 11.7 nm/yr respectively. These long-term results provide corroborative support for 
the expected excellent long-term passive corrosion behavior of Alloy 22 under 
chloride-containing aqueous environments that are relevant to repository exposure conditions. 
December 2003 5-12 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
In addition, the general corrosion model implemented in TSPA assumes that general corrosion of 
the waste package outer shell progresses uniformly over a large surface. The general corrosion 
rate is temperature dependent, and for a given temperature, the depth of penetration or thinning 
of the waste package outer shell by general corrosion is equal to the general corrosion rate at that 
temperature multiplied by the time duration that the waste package is at that temperature. 
However, general corrosion rates of metals and alloys tend to decrease with time. This 
dependence of the general corrosion rate of Alloy 22 on the exposure time was seen previously 
for the 6-month, 1-year, and 2-year data from the LTCTF. This is shown in Figure 5-6 for the 
mean general corrosion rates of Alloy 22 for those data. 
The mean general corrosion rate of the crevice samples after 5-year exposure at the LTCTF was 
0.0073 µm/yr, and the standard deviation was 0.005 µm/yr. Each data point for up to 2 years is 
the mean of the measurements on at least 144 samples, and the data point for 5-year is the mean 
of 59 samples. The trend of decreasing general corrosion rate with time is consistent with the 
expected corrosion behavior of passive alloys such as Alloy 22 under repository-type aqueous 
conditions. Therefore, the 5-year corrosion rates were conservatively selected for extrapolation 
over the repository time scale. 
Corroboration of the very low rates obtained from the temperature-dependent general corrosion 
model with the rates from alternative techniques from the scientific literature confirms that the 
life of the waste package is not limited by the rate of uniform, general corrosion. 
December 2003 5-13 No. 6: Waste Package and Drip Shield Corrosion
INTENTIONALLY LEFT BLANK 
5-14 No. 6: Waste Package and Drip Shield Corrosion 
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December 2003
Revision 1 
6. LOCALIZED CORROSION OF ALLOY 22 
IN AGGRESSIVE BRINES EVOLVED IN THE TRANSITION REGIME 
6.1 CRITERION FOR LOCALIZED CORROSION 
This section describes the localized corrosion behavior of the corrosion-resistant outer shell of 
the waste package. The threshold temperature is the temperature above which spontaneous 
localized corrosion can occur for any given environment, and no localized attack will occur 
below this temperature. The threshold temperature for localized corrosion can be determined 
from measurements of the open-circuit corrosion potential (Ecorr) and the critical potential 
(Ecritical) as functions of temperature and exposure chemistry. Spontaneous breakdown of the 
passive film and localized corrosion require that the open-circuit corrosion potential exceed or 
equal the critical potential (BSC 2003b, Sections 6.4.4 and 6.4.4.1; Farmer, McCright et al. 
2000): 
(Eq. 6-1) E ¡Ã Ecritical corr 
Localized corrosion (pitting or crevice corrosion) is a type of corrosion in which the attack 
progresses at discrete sites or in a nonuniform manner. The alloys under consideration form 
relatively stable oxide films (passive films) which impede the rate of electrochemical reactions. 
Under aggressive environmental exposure conditions, the passive films may breakdown locally 
(typically at defect sites in the film) leading to localized attack of the underlying alloy. The rate 
of localized corrosion is generally much higher than the rate of general corrosion. The current 
analysis conservatively considers crevice corrosion as the potential mode of localized corrosion 
of the waste package outer shell under repository exposure conditions. This is conservative 
because the initiation thresholds for crevice corrosion of Alloy 22 in terms of water chemistry 
and temperature are lower than for pitting corrosion (BSC 2003b, Section 6.3) 
The localized corrosion model for the waste package outer shell consists of an initiation 
component and a propagation component. In the initiation component, localized corrosion of the 
waste package outer shell occurs when the open-circuit corrosion potential (Ecorr) is equal to or 
greater than a critical potential (Ecritical), that is .E (equal Ecritical . Ecorr) less than or equal to 0. 
Both of the crevice corrosion initiation model components (i.e., Ecorr and Ecritical) are represented 
as functions of temperature, pH, chloride ion concentration, nitrate ion concentration. When 
localized corrosion occurs, the rate of localized corrosion propagation is assumed to occur at a 
(time-independent) constant rate (Section 6.7). This assumption may be conservative (BSC 
2003b, Section 6.3). 
Section 6.2 focuses on the determination of the long-term open-circuit or corrosion potential 
(Ecorr). Section 6.3 discusses cyclic potentiodynamic polarization techniques and their use to 
determine the critical potential (Ecritical) for breakdown of the passive film. The breakdown 
potential corresponds to the onset of passive film destabilization and is higher than the critical 
potentials designated for reformation of the passive film during the negative (cathodic) potential 
scan. The different types of cyclic potentiodynamic polarization curves relevant to the current 
analysis are reviewed. In Section 6.4 different methods of choosing the value of Ecritical are 
discussed. The localized corrosion initiation model is discussed in Section 6.5 with application 
December 2003 6-1 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
to the characteristic seepage waters presented in Section 6.6. Section 6.7 discusses the localized 
corrosion penetration rate model. 
6.2 LONG-TERM OPEN-CIRCUIT CORROSION POTENTIAL DATA ANALYSIS 
Because the corrosion potential of Alloy 22 may change over time, it is important to know the 
most probable value of long-term corrosion potential (Ecorr) for Alloy 22 under different 
environmental conditions and the uncertainty associated with that corrosion potential to evaluate 
the localized corrosion susceptibility of the waste package outer shell in the repository. The 
localized corrosion initiation model of the waste package outer shell assumes that localized 
corrosion will only occur when Ecorr is equal to or greater than a critical potential (crevice 
repassivation potential (Ercrev) in the current model analysis) (i.e., Ecritical equals Ercrev). That is, if 
Ecorr is less than Ercrev, general or passive corrosion will occur. General corrosion rates are 
expected to be exceptionally low. 
The specimens used to evaluate Ecorr of Alloy 22 as a function of immersion time were machined 
from sheet and bar stock. There were two main groups of specimens, (1) welded U-bend 
specimens and (2) untested rod specimens. Approximately half of the U-bend specimens tested 
for long-term corrosion potential were from the LTCTF, and other half of the U-bend specimens 
were not previously exposed to any electrochemical test condition. The U-bend specimens from 
the LTCTF already had the passive film and other surface alterations from the exposure in the 
LTCTF. The data and test details are reported in the source DTN: LL020711612251.017. 
Long-term corrosion potential behaviors of some selected Alloy 22 samples in the SDW, SAW 
and SCW solutions from the LTCTF tanks are shown in Figure 6-1. The figure shows that after 
initial changes, the corrosion potentials of the Alloy 22 samples become stable after about 
100 days of testing. It is shown that the results of the welded U-bend samples 
(Samples DUB052 and DUB159) and (nonwelded) rod samples (Samples DEA2850, DEA2851 
and DEA2852) in aged SAW at 90°C show no significant differences in their long-term 
open-circuit corrosion potential behaviors. 
December 2003 6-2 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
Source: DTN: SN0308T0506303.003. 
Figure 6-1. Open-Circuit Corrosion Potential Measurements for Samples of Alloy 22 in Three Types of 
Long Term Corrosion Test Facility Solutions (Simulated Dilute Water, Simulated Acidified 
December 2003 
Water, and Simulated Concentrated Water), as a Function of Time 
The values of Ecorr of Alloy 22 in SAW are higher than those in other aged LTCTF solutions. 
The apparent steady state Ecorr values of Alloy 22 in SAW (an acidic solution) at 60°C and 90°C 
are in the order of 300 to 400 mV versus SSC (SSC is Ag/AgCl). Figure 6-2 compares the 
long-term corrosion potential evolution of freshly polished Alloy 22 rods in freshly prepared 
SAW at 90°C with the corrosion potential evolution of the welded U-bend and rod samples in 
aged SAW at 90°C. Because of the long-term nature of the tests, the reference electrodes had to 
be replaced regularly during the tests due to their operations outside the specified accuracy 
range. The reference electrode replacements are indicated in the figure. Initially the Alloy 22 
rods in the fresh SAW solution had corrosion potentials on the order of about -150 mV versus 
SSC. However, over approximately 100 days of testing, the corrosion potential increased to a 
more oxidizing potential value of approximately 330 mV versus SSC. This apparent stable 
corrosion potential was maintained over the balance of the approximately 300-day test, slowly 
reaching a maximum oxidizing value near 400 mV versus SSC. This high value of Ecorr is 
probably due to the formation of a protective chromium rich oxide film on the surface of the 
Alloy 22 electrodes. The test results show that, regardless of the initial condition of the metal 
surface or the age of the electrolyte solution, Alloy 22 eventually undergoes ennoblement in 
SAW. This ennoblement is probably promoted by both the pH value and the presence of nitrate 
in the solution (Estill et al. 2003). Such an ennoblement of Alloy 22 with time has also been 
reported recently, and the ennoblement was significant in acidic solutions (Jayaweera et al. 2003, 
Figures 9.12 and 9.13; Dunn et al. 2003, Figures 8 and 9). 
In addition, the results in Figure 6-2 show that the Ecorr values of Alloy 22 rods in SAW at 25°C 
are lower than the values at 90°C by about 150 mV. A similar trend is also observed for the 
welded U-bend samples in the SDW solutions. This effect could be attributed to kinetic 
6-3 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
mechanisms either in the behavior of the oxide film or on the redox reactions in solution. The 
temperature effect in the SCW solution was not evaluated because the tests at 60°C were 
terminated early. 
Source: DTN: SN0308T0506303.003. 
Figure 6-2. Open-Circuit Corrosion Potential Measurements for Samples of Alloy 22 in Simulated 
Acidified Water, at Different Conditions, and as a Function of Time 
Figure 6-3 shows the test results for the Ecorr behavior of Alloy 22 in CaCl2 solutions with 
varying chloride concentrations and the effect of the addition of nitrate. The data show that Ecorr 
of Alloy 22 in the CaCl2 solutions is affected mostly by the chloride concentration, and addition 
of nitrate ions slows down the process of attaining steady state. Ecorr for Alloy 22 in 5 M CaCl2 
solution at 120°C reached steady state in less than 50 days. The average values of Ecorr after 
more than 300 days of testing were -129 mV versus SSC (see BSC 2003b, Attachment V for the 
numerical values of the data). 
After approximately 100 days of testing, the Ecorr values for Alloy 22 in 5 M CaCl2 solutions 
with nitrate added seemed to be approaching a steady state value. The average value of Ecorr for 
Alloy 22 in 5 M CaCl2 + 0.05 M Ca(NO3)2 solution at 90°C was -39 mV versus SSC after 
100 days of testing. This solution represents a nitrate to chloride ratio of 0.01. A similar 
behavior was observed for Alloy 22 tested in 5 M CaCl2 + 0.5 M Ca(NO3)2 solution (nitrate to 
chloride ratio of 0.1) at 90°C, and the average Ecorr value was –46 mV versus SSC. For Alloy 22 
electrodes immersed in 1 M CaCl2 + 1 M Ca(NO3)2 solution (nitrate to chloride ratio of 1) at 
90°C, it appears that, after 120 days of testing, the Ecorr values had approached a steady state 
value, and the average Ecorr value was 168 mV versus SSC. 
December 2003 6-4 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
Source: DTN: SN0308T0506303.003. 
NOTE: Higher resolution and longer duration data are shown in Figure 6-11 for DEA 105 and DEA 106 and in 
Figure 6-15 for DEA 2800 and DEA 2801. 
Figure 6-3. Open-Circuit Corrosion Potential Measurements for Samples of Alloy 22 in CaCl2 Solutions, 
December 2003 
with Various Concentrations and Inhibitor Levels, and as a Function of Time 
Figure 6-4 shows the steady state open-circuit potentials (or corrosion potentials) of all the 
Alloy 22 samples (those shown in the previous figures plus the rest of the data in the input 
DTN: LL020711612251.017), as a function of chloride concentration. The figure shows that the 
sample geometry or configuration and the metallurgical condition have negligible effect on the 
long-term steady-state corrosion potential of Alloy 22. It is shown that the steady state corrosion 
potential decreases with chloride concentration, which is consistent with the fact that higher 
chloride concentration makes the metal more active. Also, part of the potential decrease with 
chloride concentration may be due to the ‘salting out effect’ because the dissolved oxygen 
decreases with increasing salt concentration. 
As discussed later in the model analysis section, the steady-state corrosion potential is affected 
significantly by the solution pH. 
6-5 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
Figure 6-4. Open-Circuit Corrosion Potential for Alloy 22, with Various Sample Configurations, 
Metallurgical Conditions, and Chloride Concentration, after Exposure for Extended Periods 
Source: DTN: SN0308T0506303.003. 
December 2003 6-6 No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
6.3 CYCLIC POTENTIODYNAMIC POLARIZATION TECHNIQUES 
Initial Approach–Cyclic potentiodynamic polarization has been used by the project for several 
years as a means of measuring the critical potential of the corrosion-resistant outer layer of waste 
package (Ecritical), relative to the open-circuit corrosion potential (Ecorr). Hypothetical cyclic 
potentiodynamic polarization curves for Type 316L stainless steel, Alloy 22, and titanium are 
shown in Figure 6-5 to illustrate the differences in the corrosion resistance of these materials. 
Cyclic potentiodynamic polarization measurements have been based on a procedure similar to 
ASTM G 5-94, with slight modification. For example, ASTM G 5-94 calls for an electrolyte of 
1N H2SO4, whereas SDW, SCW, SAW, SSW, BSW (all aerated) and near-saturation CaCl2 
solutions with various levels of nitrate (deaerated) are used here. Use of aerated solutions is also 
in contrast to the procedure that calls for de-aerated solutions. These choices were made to 
better replace the expected repository exposure conditions (aerated concentrated solutions). To 
illustrate the differences in the cyclic potentiodynamic polarization results for Alloy 22 and 
Type 316L stainless steel, the curves are categorized as Type 1, 2, or 3 (CRWMS M&O 2000a, 
Section 6.4.1). 
NOTE: The critical potential can be defined as either the breakdown potential, which corresponds to the onset of 
passive film destabilization, or the repassivation potential evident during the reverse scan. 
Figure 6-5. Hypothetical Cyclic Potentiodynamic Polarization Curves for 316L, Alloy 22, and Titanium 
December 2003 6-7 
in High-Chloride Solutions 
The initial approach selected threshold potentials for localized corrosion based on the 
three generic types of cyclic potentiodynamic polarization curves. The characteristics of each 
type are discussed below. The samples were noncreviced disc samples. 
Type 1 Cyclic Potentiodynamic Polarization Curves–A generic Type 1 curve exhibits 
complete passivity (no passive film breakdown) between the open-circuit corrosion potential and 
the breakdown potential, or the potential where oxygen evolution begins if no breakdown is 
No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
observed (CRWMS M&O 2000a). Type 1 behavior, shown in Figure 6-6, is observed for 
Alloy 22 in the standardized SSW test medium near its ambient-pressure boiling point of 
approximately 120°C. This saturated sodium-potassium-chloride-nitrate electrolyte was 
formulated to represent the type of concentrated electrolyte that might evolve on a hot waste 
package surface. It is evident from Figure 6-6 that Alloy 22 maintains passivity at potentials up 
to the reversal potential (1,200 mV versus Ag/AgCl), even under these relatively aggressive 
conditions. 
Source: CRWMS M&O 2000a, Figure 3. 
NOTE: Data for Alloy 22 in SSW at 120°C show no loss of passivity, even with a voltage reversal of 1,200 mV 
versus Ag/AgCl, and exhibit Type 1 behavior (noncreviced disc sample). 
Figure 6-6. Cyclic Potentiodynamic Polarization Data for Alloy 22 in Simulated Saturated Water at 
December 2003 6-8 
120°C 
Type 2 Cyclic Potentiodynamic Polarization Curves–A generic Type 2 curve (e.g., Section 7, 
Figure 7-5) exhibits a well-defined oxidation peak between the open-circuit corrosion potential 
and the breakdown potential, or the potential where oxygen evolution begins if no breakdown is 
observed (CRWMS M&O 2000a). The anodic oxidation peak (process) is due to the conversion 
of metals in the passive film to higher, perhaps more soluble, oxidation states. This oxidation 
process is accompanied by changes in X-ray photoelectron spectroscopy and surface 
morphology, both of which have been well documented (BSC 2003b, Section 6.4.1). 
Type 3 Cyclic Potentiodynamic Polarization Curves–A generic Type 3 curve exhibits a 
complete breakdown of the passive film and active pitting at potentials relatively close to the 
open-circuit corrosion potential. The critical pitting potential is evident in Figure 6-7. Type 3 
behavior has only been observed with Type 316L stainless steel. 
No. 6: Waste Package and Drip Shield Corrosion
Revision 1 
Source: CRWMS M&O 2000a, Figure 8. 
NOTE: Data show pitting potential very close to the open-circuit corrosion potential and exhibit Type 3 behavior. 
Figure 6-7. Cyclic Potentiodynamic Polarization Data for Type 316L Stainless Steel in Simulated 
December 2003 
Saturated Water at 120°C (Noncreviced Disc Sample) 
The initial approach to evaluate localized corrosion behavior utilized primarily uncreviced 
specimens with most testing performed in the initially defined set of relevant environments. 
Subsequent testing indicated these initial test environments were relatively benign (not 
conducive to initiation of localized corrosion). More recent environmental modeling and 
experimental assessments have expanded the full range of potential brine environments that 
could potentially contact the waste package over time resulting in the need to expand the test 
program to cover the full range of credible environments