CORROSION OF ALLOY 22 INDUCED BY 

DUST DELIQUESCENCE BRINES 

Prepared for 

U.S. Nuclear Regulatory Commission 

Contract NRC–02–07–006 

Prepared by 

L. Yang1 

R. Pabalan1 

P. Shukla1 

M. Juckett1 

X. He1

K.T. Chiang1

H. Gonzalez2

 

T. Ahn2

1

Center for Nuclear Waste Regulatory Analyses 

San Antonio, Texas 

2

U.S. Nuclear Regulatory Commission 

Washington, DC 

March 2011 
ii

ABSTRACT 

In the potential repository at Yucca Mountain, Nevada, nuclear waste materials would be 

housed in waste packages. The U.S. Department of Energy is considering Alloy 22 as the 

material for the outer container of the waste packages. During the preclosure period, the 

potential repository drift would be ventilated and atmospheric aerosols and tunnel dusts could 

be deposited onto the waste package surfaces by the ventilation, leading to the accumulation of 

small amounts of hygroscopic salts on the waste package surfaces. The hygroscopic salts 

could deliquesce and form a brine solution on the waste package surfaces when the relative 

humidity of the in-drift environment is near or above the salt’s deliquescence relative humidity. 

This report evaluates Alloy 22 corrosion by the deliquescence brines under potential 

repository conditions. 

The Yucca Mountain dusts and their salt contents were reviewed, and the brine properties that 

may be formed from the dust salts under potential drift conditions (high temperature and low 

relative humidity) are discussed. Because the brines formed under potential Yucca Mountain 

conditions are expected to be mixed with large amounts of nondeliquescent dust mixtures, an 

experimental assessment on the corrosion of metals in contact with the deliquescent and 

nondeliquescent dust mixtures was also conducted. 

General corrosion of Alloy 22 in bulk deliquescence brines containing Na+

, K+, Cl−, and NO3

−

ions at elevated temperatures in closed systems (autoclaves) has been reported in the 

literature. Because the potential drift would be open to the atmosphere, experiments on 

Alloy 22 corrosion in open systems were conducted in the present study. In these experiments, 

the pH effects of the brine solution on the corrosion of Alloy 22 and the verification test for the 

pH effect were evaluated. The general corrosion rates of Alloy 22 in NaCl–NaNO3–KNO3 brines 

were equal to or less than 1 μm/yr [0.04 mil/yr] in the temperature range from 120 to 220 C 

[248 to 428 F] and a pH range of the brine solution from 4.8 to 8.5 (as measured at room 

temperature after tenfold dilution by weight). 

Literature data on electrochemical tests and immersion tests for localized corrosion of Alloy 22 

in large quantities of deliquescence brine systems were reviewed. A detailed model analysis on 

the possibility of localized corrosion of Alloy 22 in a nondeliquescent dust and deliquescent salt 

mixture under the potential Yucca Mountain drift conditions was conducted. Electrochemical 

potential (repassivation and corrosion) data suggest that localized corrosion in the form of 

crevice corrosion would not initiate for Alloy 22 in large quantities of brines at temperatures 

between 120 and 150 C [248 and 302 F]. However, previous investigators have reported 

pitting and crevice corrosion in immersion tests for specimens exposed to both closed and open 

systems containing large amounts of brines that contrast the electrochemical test results. 

Additional long-term immersion tests may resolve this discrepancy. Calculated cathodic 

capacities of Alloy 22 in dust deliquescence brines suggest that the high penetration rate due 

to localized corrosion may not be sustained under limited brine amounts in potential 

Yucca Mountain drift conditions. 
iii

CONTENTS 

Section Page 

ABSTRACT ................................................................................................................................... ii 

FIGURES ...................................................................................................................................... v 

TABLES ....................................................................................................................................... vii 

ACKNOWLEDGMENTS ............................................................................................................. viii 

1 INTRODUCTION ................................................................................................................. 1-1 

 1.1 Background ............................................................................................................... 1-1 

 1.2 Scope and Organization of the Report ...................................................................... 1-2 

 1.3 References ................................................................................................................ 1-2 

2 YUCCA MOUNTAIN DUSTS AND BRINE FORMATION AT 

 ELEVATED TEMPERATURES ........................................................................................... 2-1 

 2.1 Overview of Salt Deliquescence ................................................................................ 2-1 

 2.2 Characterization Data on Dusts Sampled From Yucca Mountain 

 and Vicinity ................................................................................................................ 2-3 

 2.2.1 Yucca Mountain Crest Data ............................................................................. 2-3 

 2.2.2 Yucca Mountain Tunnel Data .......................................................................... 2-4 

 2.2.3 Atmospheric (Cyclone) Dust Data ................................................................... 2-5 

 2.2.4 Remote Dust Data ........................................................................................... 2-7 

 2.3 Potential Composition of Brines Formed by Deliquescence of Salts 

 Present in Yucca Mountain Dusts ............................................................................. 2-7 

 2.4 Effect of Capillary Retention by Dusts on the Corrosivity of 

 Deliquescence Brines .............................................................................................. 2-18 

 2.4.1 Test Methods ................................................................................................. 2-18 

 2.4.2 Test Results ................................................................................................... 2-19 

 2.4.2.1 Test 1—NaCl–Quartz Mixture at 70 C [158 F] ............................... 2-19 

 2.4.2.2 Test 2—NaCl–Dust Mixture at 70 C [158 F] .................................. 2-21 

 2.4.2.3 Test 3—NaCl–Quartz and NaCl–Dust Mixtures at 70 C [158 F] .... 2-23 

 2.4.2.4 Test 4—Effect of Dust Particle Sizes ................................................ 2-24 

 2.4.2.5 Test 5—Alloy 22 and Carbon Steel Corrosion in 

 NaCl–NaNO3–KNO3 Dust Mixture at 125 C [257 F] ....................... 2-24 

 2.5 Summary and Conclusions ...................................................................................... 2-26 

 2.6 References .............................................................................................................. 2-28 

3 CORROSIVITY OF DELIQUESCENT BRINES TO ALLOY 22 AT 

 ELEVATED TEMPERATURES ........................................................................................... 3-1 

 3.1 Corrosion Rate of Alloy 22 in Nitrate–Chloride Salt Systems Measured in 

 Closed Systems ......................................................................................................... 3-1 

 3.2 Corrosion Rate of Alloy 22 in Nitrate–Chloride Salt Systems Measured in 

 Open Systems ........................................................................................................... 3-2 

 3.2.1 Test Specimens ............................................................................................... 3-2 

 3.2.2 Test Vessels for Weight Loss Corrosion Rate Measurements ........................ 3-4 

 3.2.3 Test Vessels for Characterization of Vapor Phase Acidity ............................... 3-5 

 3.2.4 Weight Loss Results ........................................................................................ 3-6 

 3.2.5 Effect of pH on Corrosion Rate ...................................................................... 3-10 
iv

CONTENTS (continued) 

Section Page 

 3.3 Verification of the pH Effect on Alloy 22 Corrosion and Understanding of the 

 pH Variations in the NaCl–NaNO3–KNO3 Systems ................................................. 3-15 

 3.3.1 Corrosion Rates Measured With Coupled Multielectrode 

 Array Sensors ................................................................................................ 3-15 

 3.3.2 Causes for the Variations in the pH of the NaCl–NaNO3–KNO3 Systems ..... 3-16 

 3.4 Summary ................................................................................................................. 3-21 

 3.5 References .............................................................................................................. 3-23 

4 LOCALIZED CORROSION OF ALLOY 22 IN DUST DELIQUESCENCE BRINES ............ 4-1 

 4.1 Electrochemical Test Data on Localized Corrosion in Large Quantity 

 of Brines .................................................................................................................... 4-1 

 4.2 Immersion Test Data on Localized Corrosion in Large Quantity 

 of Brines .................................................................................................................... 4-5 

 4.3 Localized Corrosion Model for Alloy 22 in Dust Deliquescence Brines 

 Under Potential Repository Conditions ...................................................................... 4-6 

 4.3.1 Model Development ......................................................................................... 4-7 

 4.3.2 Model Parameters ......................................................................................... 4-12 

 4.3.3 Calculated Cathodic Capacities ..................................................................... 4-16 

 4.3.4 Discussion on Cathodic Capacity Results ..................................................... 4-19 

 4.4 Summary ................................................................................................................. 4-21 

 4.5 References .............................................................................................................. 4-21 

5 SUMMARY AND CONCLUSIONS ...................................................................................... 5-1 

 5.1 Formation of Deliquescence Brines at Elevated Temperatures and Low 

 Relative Humidities .................................................................................................... 5-1 

 5.2 Corrosivity of Potential Deliquescence Brines to Alloy 22 Waste 

 Package Materials ..................................................................................................... 5-2 

 5.3 Localized Corrosion of Alloy 22 in Deliquescence Brines Under Potential 

 Yucca Mountain Conditions ....................................................................................... 5-2 
v

FIGURES 

Figure Page 

2-1 Plot of Calculated Waste Package Temperature and Relative Humidity at the 

 Potential Yucca Mountain Repository ............................................................................. 2-2 

2-2 Calculated Relative Humidity at 25 °C [77 °F] Over an Aqueous Solution Saturated 

 With Either NaCl, NaNO3, or Both as a Function of Salt Mixture Composition ............... 2-2 

2-3 Calculated Deliquescence Relative Humidity Shown as Contour Lines of 

 Different Salts and Salt Mixtures in the System NaCl–KCl–MgCl2 at 90 C [194 F] ..... 2-4 

2-4 Calculated Deliquescence Relative Humidity of Different Salts and Salt Mixtures 

 in the System Na–K–Cl–NO3 as a Function of Temperature .......................................... 2-5 

2-5 Mineral Precipitates Calculated to Form Upon Evaporation of Initially Dilute 

 Dust Leachates ............................................................................................................. 2-12 

2-6 Evolution of the Nitrate-to-Chloride Molar Ratio as More Water Is Taken Up by the 

 Deliquescent Salt Assemblage at 110 C [230 F] ....................................................... 2-17 

2-7 Schematic of Test Setup Used in Studying the Effect of Capillary Retention by 

 Dusts on the Corrosivity of Deliquescence Brines ........................................................ 2-19 

2-8 Maximum Corrosion Current Measured From Test 1 Carbon Steel Probes ................. 2-20 

2-9 Posttest Photographs of the Test 1 Carbon Steel Probes ............................................ 2-21 

2-10 (a) Measured Maximum Corrosion Current and (b) Calculated Corrosion Rate 

 for Test 2 Carbon Steel Probes .................................................................................... 2-22 

2-11 Posttest Photographs of the Carbon Steel Probes From Test 2 .................................. 2-23 

2-12 Maximum Corrosion Current Measured From the Test 3 Carbon Steel Probes ........... 2-24 

2-13 Maximum Corrosion Current Measured From the Test 4 Carbon Steel Probes ........... 2-25 

2-14 (a) Maximum Corrosion Current Measured From Test 5 Probes ................................. 2-27 

3-1 Schematic Diagrams for the Creviced Specimens Used in the Tests ............................ 3-3 

3-2 Schematic of Typical Glass Vessels Used for the Weight Loss Experiment .................. 3-5 

3-3 Schematic of Typical Glass Flask Used for the Characterization of Vapor 

 Phase pH ........................................................................................................................ 3-6 

3-4 Optical Appearance of the Specimens After the Tests in Liquid Phase ......................... 3-9 

3-5 pH of the Vent-Bubbler Water and the Deionized Makeup Water That Was Added 

 to the Test Vessels and Vent-Bubblers During the Tests ............................................. 3-11 

3-6 pH of the Vent-Bubbler Water During the Tests Conducted in Vessels 

 A', B', and C' ................................................................................................................. 3-12 

3-7 Average Corrosion Rates for Alloy 22 as a Function of pH of the 

 Vent-Bubbler Water ...................................................................................................... 3-13

3-8 Relationship Between the pH of the Diluted Sample From the Liquid Phase and the 

 pH of the Bubbler Water ............................................................................................... 3-14 

3-9 Relationship Between Corrosion Rates of the Alloy 22 Specimens Exposed 

 to the Liquid Phase ....................................................................................................... 3-14 

3-10 (a) Schematic Diagram of the Coated Multielectrode Array Sensor and 

 (b) Photograph of a Coated Alloy 22 Multielectrode Array Sensor ............................... 3-15 

3-11 Maximum Corrosion Rates for Alloy 22 Measured With a Coupled 

 Multielectrode Array Sensor at Different pH in the NaCl–NaNO3–KNO3 System ......... 3-17 

3-12 Relationship Between the Maximum Corrosion Rates Measured With the 

 Alloy 22 Coupled Multielectrode Array Sensor and the pH of the Diluted Samples ..... 3-17 

3-13 Effect of Atmospheric Components on the pH of the Condenser Water Above the 

 Test Flasks Containing Wet NaCl, NaNO3, and KNO3 Salt Mixtures ............................ 3-19 
vi

FIGURES (continued) 

Figure Page 

3-14 Calculated Effect of Addition of SiO2 and Dissolved SiO2 Concentration on the 

 pH of the NaCl–NaNO3–KNO3 Solution at 160 C [320 F] .......................................... 3-19 

3-15 Calculated Effect of the Addition of B(OH)3 and Dissolved B(OH)3 Concentration 

 on the pH of NaCl–NaNO3 Solution at 160 C [320 F] ................................................ 3-20 

3-16 Calculated Effect of the Addition of P2O5 and the Dissolved P2O5 Concentration 

 on the pH of NaCl–NaNO3–KNO3 Solution at 160 C [320 F] ..................................... 3-20 

3-17 (a) Corrosion Rates and (b) Open Circuit Potential Versus pH for Alloy 22 in 

 Hydrochloric, Sulfuric, and Nitric With 4 m NaCl at 90 C [194 F] ............................... 3-22 

4-1 Corrosion Potential Versus Time of Alloy 22 in 25 m NO3

−

 +1 m Cl− Solution With 

 the pH Adjusted to Different Values at 110 C [230 F]. ................................................. 4-2 

4-2 Graphical Representation of the Erp Data Tabulated in Table 4-1 .................................. 4-4 

4-3 Schematic Diagram of the Crevice Corrosion Process. ................................................. 4-7 

4-4 Schematic Representation of the Cathodic Region With Corresponding Cathodic 

 and Anodic Polarization Curves ................................................................................... 4-10 

4-5 Schematic Representation of Dust Deliquescence Brine Layer and Dust 

 Layer on Alloy 22 Surface ............................................................................................ 4-15 

4-6 (a) Electrode Potential and (b) Current Density Distribution of Alloy 22 in the 

 Cathodic Region ........................................................................................................... 4-17 

4-7 Cathodic Capacity of Alloy 22 in Dust Deliquescence Brine as a Function of 

 the Difference Between the Corrosion Potential and the Repassivation Potential ....... 4-18 

4-8 Cathodic Capacity of Alloy 22 in Dust Deliquescence Brines as a Function of the 

 Difference Between the Corrosion Potential and the Repassivation Potential ............. 4-19 

4-9 Cathodic Capacity of Alloy 22 in Dust Deliquescence Brines as a Function of the 

 Difference Between the Corrosion Potential and the Repassivation Potential ............. 4-20 
vii

TABLES 

Table Page 

2-1 Center for Nuclear Waste Regulatory Analyses Data on the Composition of 

 Leachates of Dusts Sampled From the Surface of Yucca Crest (mg/L) ......................... 2-5 

2-2 Mineralogical Composition of Surface Samples Based on Center for Nuclear 

 Waste Regulatory Analyses Powder X-Ray Diffraction Analysis .................................... 2-6 

2-3 Composition of Leachates of Dusts Sampled From Inside the Tunnel at 

 Yucca Mountain (mg/L) .................................................................................................. 2-6 

2-4 Mineralogical Composition of Tunnel Samples Based on Center for Nuclear 

 Waste Regulatory Analyses Powder X-Ray Diffraction Analysis .................................... 2-6 

2-5 Center for Nuclear Waste Regulatory Analyses Data on the Composition of 

 Leachates of Atmospheric Dusts Collected With the Torit Cyclone (mg/L) .................... 2-8 

2-6 Mineralogical Composition of Atmospheric Samples Based on Center for Nuclear 

 Waste Regulatory Analyses Powder X-Ray Diffraction Analysis .................................... 2-8 

2-7 Leachate Composition From the National Atmospheric Deposition Program 

 Regional Precipitation (Rain Out) ................................................................................... 2-9 

2-8 Mineralogical Composition of Two Remote Sample Based on Center for Nuclear 

 Waste Regulatory Analyses Powder X-Ray Diffraction Analysis .................................... 2-9 

2-9 Composition of Leachable Portion of Dust Samples Taken From 

 Yucca Mountain Tunnels .............................................................................................. 2-10 

2-10 Salt Assemblages Derived From Evaporation of Dust Leachates With Composition 

 Listed in Table 2-9. ....................................................................................................... 2-11 

2-11 Calculated Mutual Deliquescence Relative Humidity Percent As a Function of 

 Temperature of Six Samples Representing Six Key Salt Assemblages ....................... 2-16 

2-12 Calculated Molar Concentration Ratio of Chloride-to-Nitrate Ions As a Function of 

 Temperature of Six Samples Representing Six Key Salt Assemblages ....................... 2-17 

2-13 Chemical Compositions of Dust Leachate (ppm) ......................................................... 2-19 

2-14 Chemical Composition of Brine Formed by Deliquescence of a 

 NaCl–NaNO3–KNO3 Salt Mixture at a Temperature of 125 C [257 F] ....................... 2-26 

3-1 Specimens Used in the Experiments .............................................................................. 3-3 

3-2 Chemical Compositions of Specimen and Weld Fillers (in Weight Percent) .................. 3-4 

3-3 Corrosion Rates of Alloy 22 Specimens in the Liquid Vapor Phases and the pH 

 of the Water in the Vent Bubbler Conducted in 2006 ..................................................... 3-7 

3-4 Corrosion Rates of Alloy 22 Specimens in the Liquid and Vapor Phases and the pH 

 of the Water in the Vent Bubbler Conducted in 2005 ................................................... 3-10 

4-1 Reported Value of Repassivation Potential Erp............................................................... 4-3 

4-2 Dust Particle Size, Dust Layer Thickness, Brine Film Thickness, and Ionic 

 Conductivity and Effective Ionic Conductivity of Dust Deliquescence Brine ................. 4-13 

4-3 Calculated Deliquescence Relative Humidity, Dissolved Oxygen Concentration, 

 and Ionic Conductivity of Deliquescence Salts at Various Temperatures .................... 4-14 

4-4 Effective Ionic Conductivity of Dust Deliquescence Brine Formed Between 

 120 and 170 C [248 and 338 F] ................................................................................. 4-15 
viii

ACKNOWLEDGMENTS 

This report describes work performed by the Center for Nuclear Waste Regulatory Analyses 

(CNWRA®) and its contractors for the U.S. Nuclear Regulatory Commission (NRC) under 

Contract No. NRC–02–07–006. The activities reported here were performed on behalf of the 

NRC Office of Nuclear Material Safety and Safeguards, Division of High-Level Waste 

Repository Safety. This report is an independent product of CNWRA and does not necessarily 

reflect the view or regulatory position of the NRC. The NRC staff views expressed herein are 

preliminary and do not constitute a final judgment or determination of the matters addressed or 

of the acceptability of a license application for a geologic repository at Yucca Mountain. 

The authors would like to thank J. Auguste, S. Young, and B. Derby for laboratory assistance. 

The authors acknowledge the technical review of G. Cragnolino, editorial review of L. Mulverhill, 

programmatic review of B. Sagar and D. Pickett, the technical discussions with Y.-M. Pan and 

K. Axler, and the administrative support of A. Ramos, J. Gonzalez, and B. Street in preparing 

this report. 

The fabrication of the coupled multielectrode array sensor used in this study was financially 

supported by a CNWRA-sponsored Focused Internal Research and Development program. 

QUALITY OF DATA, ANALYSES, AND CODE DEVELOPMENT 

DATA: All CNWRA-generated original data contained in this report meet the quality assurance 

requirements described in the Geosciences and Engineering Division Quality Assurance 

Manual. Sources for other data should be consulted for determining the level of quality for those 

data. Experimental data analyses have been recorded in CNWRA Scientific Notebooks 679E, 

771, 737, and 851E. 

ANALYSES AND CODES: The computer software Corrosion AnalyzerTM (OLI Systems, Inc., 

2006) was used in the analyses contained in this report. Corrosion Analyzer is commercial 

software controlled under the CNWRA quality assurance procedure “Development and Control 

of Scientific and Engineering Software” (TOP–18). Documentation for thermodynamic 

calculation in Section 2 can be found in Scientific Notebook 679E. 

Reference 

OLI Systems, Inc. “Corrosion Analyzer Software.” Version 2.0. Morris Plains, New Jersey: 

OLI Systems, Inc. 2006. 

Pabalan, R.T. “Verification and Validation of Mixed-Solvent Electrolyte Model for the 

Environmental Simulation Program and Analyzer Codes Developed by OLI Systems, Inc. 

(Morris Plains, NJ).” SN 679E. San Antonio, Texas: CNWRA. 2004.
1-1 

1 INTRODUCTION 

1.1 Background 

The U.S. Department of Energy (DOE) is considering Alloy 22 as the outer container material 

of the waste package (Anderson, et al., 2003), which is an important component of the 

engineered barrier system designed to prevent or delay radionuclide release to the accessible 

environment. The U.S. Nuclear Regulatory Commission (NRC) and the Center for Nuclear 

Waste Regulatory Analyses staff are evaluating the corrosion behavior of Alloy 22 in the 

potential Yucca Mountain repository environment. Under the nominal-case scenario, corrosion 

of Alloy 22 is an important degradation process limiting the lifetime of the waste package 

(Bechtel SAIC Company, LLC, 2004). 

During the preclosure period, which may last for several decades after waste package 

emplacement, the repository emplacement drifts would be actively ventilated (Sandia National 

Laboratory, 2007a). Atmospheric aerosols and tunnel dusts could be deposited onto the waste 

package surfaces by forced ventilation, leading to the accumulation of small amounts of 

hygroscopic salts on such surfaces (Bechtel SAIC Company, LLC, 2005a,b; Sandia National 

Laboratories, 2007a; Craig, et al., 2004). Hygroscopic salts could deliquesce (sorb moisture 

from the atmosphere) and form a brine solution on the waste package surfaces when the 

relative humidity of the in-drift environment is near or above the salt’s deliquescence 

relative humidity. 

According to Sandia National Laboratories (2007b), the temperature of the waste package 

surface is expected to be relatively low and constant during the ventilation period. However, the 

temperature is expected to increase after ventilation ceases or the drift is closed, reaching a 

maximum value shortly after the closure, followed by a gradual decrease thereafter. The 

maximum temperature at the waste package surface could be 203 C [397 F]. In parallel, the 

relative humidity in the drift is expected to be low during the ventilation period because of the 

relatively dry air blown from the outside and is expected to increase as soon as the drift is 

closed because water vapor pressure would tend to reach the ambient pressure. After drift 

closure, the relative humidity of the waste package is estimated to decrease slightly with the 

increase in temperature and then rise after the waste package temperature has peaked after 

approximately 200 years (Sandia National Laboratory, 2007b). 

According to recent studies (Rard, 2004; Rard, et al., 2006), some salt assemblages (such as 

NaCl–NaNO3–KNO3) that may develop under the potential repository conditions can deliquesce 

at ambient pressure and temperatures over 190 C [374 F] with a corresponding relative 

humidity below 4.4 percent. Hence, brine may be present during the entire high 

temperature period. 

The presence of brines on the waste package surfaces has important implications because 

many metals are more susceptible to localized corrosion or have higher general corrosion 

rates at higher temperatures when immersed in a brine system (Orme, et at., 2004; Yang, et al., 

2003). 

This report assesses Alloy 22 corrosion induced by dust deliquescence brines under potential 

drift conditions. 
1-2 

1.2 Scope and Organization of the Report 

This report is organized into five chapters, including this introduction as Chapter 1. 

The potential Yucca Mountain dusts and their salt contents were assessed, and the brine 

properties that may be formed from the dust salts under the potential emplacement drift 

conditions (high temperature and low relative humidity) are detailed in Chapter 2. Chapter 2 

also includes an experimental assessment on the effect of nondeliquescent dusts mixed with 

soluble deliquescent salts on the corrosion of the metals contacting dust mixtures. 

The general corrosion of Alloy 22 in bulk deliquescence brines containing Na+

, K+, Cl−, and NO3

−

ions at elevated temperatures is the subject of Chapter 3. A review on the corrosion of Alloy 22 

in closed systems (autoclaves) and experiments on the corrosion of Alloy 22 in open systems 

(no deaeration, no exclusion of air) is presented in the chapter. The pH effects of the brine 

solution on Alloy 22 corrosion and the verification test for the pH effect are detailed in that 

chapter as well. 

Localized corrosion of Alloy 22 in deliquescence brines at elevated temperatures is discussed in 

Chapter 4. This chapter includes a review of electrochemical test and immersion test results 

dealing with localized corrosion of Alloy 22 exposed to large quantities of deliquescence brines. 

A detailed model analysis on the possibility of localized corrosion of Alloy 22 in the mixture of 

nondeliquescent dusts and deliquescent salts under the potential Yucca Mountain drift 

conditions is also presented in this chapter. 

Finally, a summary of conclusions is included in Chapter 5. 

1.3 References 

Anderson, M.J., N.R. Brown, J.D. Cloud, P.R.Z. Russell, and J. Trautner. “Waste Package 

Design for License Application.” Proceedings of the 10th International High-Level Radioactive 

Waste Management Conference, Las Vegas, Nevada, March 30–April 3, 2003. Section E–7. 

Published on CD-ROM. La Grange Park, Illinois: American Nuclear Society. pp. 714–720. 

2003. 

Bechtel SAIC Company, LLC. “Analysis of Dust Deliquescence for FEP Screening.” Report 

ANL–EBS–MD–000074. Rev. 01. Las Vegas, Nevada: Bechtel SAIC Company, LLC. 2005a. 

–––––. “FEPs Screening of Processes and Issues in Drip Shield and Waste Package 

Degradation.” Report ANL–EBS–PA–000002. Rev. 05. Las Vegas, Nevada: Bechtel SAIC 

Company, LLC. 2005b. 

_____. “WAPDEG Analysis of Waste Package and Drip Shield Degradation.” 

Report ANL–EBS–PA–000001. Rev. 02. Las Vegas, Nevada: Bechtel SAIC Company, LLC. 

2004a. 

Craig, L., S. Carrol, and T. Wolery. “Deliquescence of NaCl–NaNO3 and KNO3–NaNO3 Salt 

Mixtures at 90 °C in Water-Rock Interaction.” R.B. Wanty and R.R. Seal II, eds. Vol. 2. 

London, England: Taylor Francis Group. pp. 1,275–1,278. 2004. 
1-3 

Orme, C.A., J. Gray, J. Hayes, L. Wong, R. Rebak, S. Carroll, J. Harper, and G. Gdowski. 

“Fiscal Year 2004 Summary Report: General Corrosion and Passive Film Stability.” 

UCRL–TR–208588. Livermore, California: Lawrence Livermore National Laboratory. 2004. 

Rard, J.A. “Results From Boiling Temperature Measurements for Saturated Solutions in the 

Systems NaCl+KNO3+H2O, NaNO3+KNO3+H2O and NaCl+NaNO3+KNO3+H2O.” Report 

UCRL–TR–207054. Livermore, California: Lawrence Livermore National Laboratory. 2004. 

Rard, J.A., K.J. Staggs, S.D. Day, and S.A. Carroll. “Boiling Temperature and Reversed 

Deliquescence Relative Humidity Measurements for Mineral Assemblages in the 

NaCl+NaNO3+KNO3+Ca(NO3)2+H2O System.” Journal of Solution Chemistry. Vol. 35. 

pp. 1,187–1,215. 2006. 

Sandia National Laboratories. “Analysis of Dust Deliquescence for FEP Screening.” 

Report ANL–EBS–MD–000074. Rev. 01, AD 01. Las Vegas, Nevada: Sandia National 

Laboratories. 2007a. 
1-4 

–––––. “Multiscale Thermohydrologic Model.” Report ANL–EBS–MD–00004. Rev. 03. 

Las Vegas, Nevada: Sandia National Laboratories. 2007b. 

Yang, L., R.T. Pabalan, L. Browning, and D.S. Dunn. “Corrosion Behavior of Carbon Steel and 

Stainless Steel Materials Under Salt Deposits in Simulated Dry Repository Environments.” 

Scientific Basis for Nuclear Waste Management XXVII. Symposium Proceedings 757. 

R.J. Finch and D.B. Bullen, eds. Pittsburgh, Pennsylvania: Materials Research Society. 

pp. 791–797. 2003.
2-1 

2 YUCCA MOUNTAIN DUSTS AND BRINE FORMATION 

AT ELEVATED TEMPERATURES 

Deliquescence is a process in which a soluble substance sorbs water vapor from the air to form 

a saturated aqueous solution. At the potential high-level waste repository at Yucca Mountain, 

Nevada, deposition of aerosols and dusts entrained in ventilation air in the drifts of the potential 

repository could lead to the accumulation of hygroscopic salts on drip shield and waste package 

surfaces. The deliquescence of these salts, which would occur when the relative humidity in the 

environment is at or above the deliquescence relative humidity of the salt or the mutual 

deliquescence relative humidity of the salt mixture, could form brines potentially corrosive to the 

metallic engineered barriers. Formation of these brines could affect the performance of the 

waste package during the thermal period when the drift wall temperature is above the boiling 

point of water and seepage water is unlikely to enter the repository drift. During this time period 

(see Figure 2-1 for a typical illustration), localized corrosion of the Alloy 22 waste package outer 

container due to dust deliquescence could be facilitated by the elevated system temperature if 

corrosive brine contacts the waste package surface. 

This chapter assesses the potential for deliquescence brines to form in the potential repository 

drift environment, to contact the waste package surface, and to initiate localized corrosion of the 

Alloy 22 waste package material. The assessment is based on geochemical data on dust 

samples taken from Yucca Mountain and its vicinity (e.g., Bechtel SAIC Company, 2004). The 

geochemical information is used in thermodynamic calculations to determine whether 

deliquescence brines can potentially form at elevated temperatures—corresponding to the dust 

deliquescence period indicated in Figure 2-1—and whether the brines that do form would have 

chemical compositions that could initiate localized corrosion of the Alloy 22 waste package outer 

container. In addition, experiments were conducted to determine whether capillary retention by 

nondeliquescent dust particles could prevent brines from contacting the waste package surface 

and mitigate localized corrosion degradation of the material. 

2.1 Overview of Salt Deliquescence 

The concept of deliquescence relative humidity for salt mixtures is illustrated in Figure 2-2, 

which shows the calculated relative humidity at 25 °C [77 °F] over an aqueous solution 

saturated with NaCl, NaNO3, or both. Consider, for instance, a dry mixture of NaCl and NaNO3

salts with bulk composition xA. If the ambient relative humidity, initially below the mutual 

deliquescence relative humidity, is increased, no significant uptake of water by the salt mixture 

occurs until the mutual deliquescence relative humidity is reached. At the mutual deliquescence 

relative humidity, deliquescence occurs, and the salt mixture dissolves—NaCl completely and 

NaNO3 partially. As the ambient relative humidity is raised further, the amount of solid 

decreases, and the amount of liquid increases, with the liquid becoming richer in dissolved 

NaNO3 relative to dissolved NaCl. At a relative humidity slightly higher than RHA, the NaNO3

solid is dissolved completely and the aqueous composition will have changed from xC at the 

mutual deliquescent relative humidity to xA. Raising the ambient relative humidity further will 

cause the aqueous phase to be diluted. Likewise, for a mixture of NaCl and NaNO3 having a 

bulk composition of xB, the salt mixture dissolves at the mutual deliquescence relative 

humidity—NaCl partially and NaNO3 completely. As the ambient relative humidity is raised 

above the mutual deliquescence relative humidity, the aqueous phase becomes richer in 
2-2 

Figure 2-1. Plot of Calculated Waste Package Temperature and Relative Humidity at 

the Potential Yucca Mountain Repository (Pensado, 2006). The Temperature and 

Relative Humidity Are Kept Low by Ventilation During the Preclosure Period 

(Assumed to be 50 Years). 

Figure 2-2. Calculated Relative Humidity at 25 °C [77 °F] Over an Aqueous Solution 

Saturated With Either NaCl, NaNO3, or Both as a Function of Salt Mixture Composition. 

Xc Is the Eutonic Composition of the NaCl–NaNO3 Mixture.
2-3 

dissolved NaCl relative to dissolved NaNO3 until NaCl is dissolved completely at RHB. A salt 

mixture with bulk composition xC will dissolve completely when the ambient relative humidity is 

raised to values at or above the mutual deliquescence relative humidity. 

The deliquescence relative humidity of salts and the mutual deliquescence relative humidity of 

salt mixtures depend on composition. Figure 2-3 shows the calculated deliquescence relative 

humidity and mutual deliquescence relative humidity for salts and salt mixtures in the system 

NaCl–KCl–MgCl2 at 90 °C [194 °C]. The deliquescence relative humidity for the single salts 

NaCl, KCl, and MgCl2 are 74.3, 76.2, and 24.7 percent, respectively. For the mixture of NaCl 

and KCl, which both have relatively high deliquescence relative humidity, the depression of 

the deliquescence point due to the other salt is not large (less than 10 percent relative humidity). 

In contrast, the deliquescence point of NaCl or KCl is substantially reduced when mixed 

with MgCl2, and the mutual deliquescence relative humidity is lowest (24.2 percent) when 

the brine is saturated with the three salts NaCl (halite), MgCl26H2O (bishofite), and 

KMgCl36H2O (carnallite). 

The deliquescence relative humidity of salts and the mutual deliquescence relative humidity of 

salt mixtures also depend on temperature. Figure 2-4 shows the calculated deliquescence 

relative humidity as a function of temperature of different salts and salt mixtures in the system 

Na–K–Cl–NO3. The NaCl–NaNO3–KNO3 salt mixture is important because it has low 

deliquescence relative humidity and the U.S. Department of Energy (DOE) considers it to be a 

salt assemblage that could control the deliquescence of dusts deposited inside the drift of the 

potential Yucca Mountain repository (Bechtel SAIC Company, LLC, 2005). DOE experiments 

suggest that NaCl–NaNO3–KNO3 brines can form at temperatures exceeding 190 °C [374 °F] 

(Bechtel SAIC Company, LLC, 2005). 

2.2 Characterization Data on Dusts Sampled From Yucca Mountain 

and Vicinity 

DOE sampled dusts on the surface of Yucca Mountain, in the Yucca Mountain tunnels, and at 

remote locations near the Yucca Mountain site {within 20 mi [32 km]}. These dusts provided 

mineralogical, geochemical, and leachate data that contributed to an analysis of the potential 

effect of dust deliquescence brines on the localized corrosion of the Alloy 22 waste package 

outer container (Bechtel SAIC Company, LLC, 2005). 

2.2.1 Yucca Mountain Crest Data 

The U.S. Geological Survey staff collected samples from the ground surface of the 

Yucca Mountain crest {approximately 1,494 m [4,900 ft] elevation} in June 2004 on the lee of 

large rocks and bushes and from bedrock depressions (Peterman, 2006). These samples were 

collected using a brush and a collection tray and then were transferred to a sample bottle. 

Approximately 200 g [7 oz] of these samples were collected, and 10-g [0.4-oz] splits of these 

samples were provided to the Center for Nuclear Waste Regulatory Analyses (CNWRA®) for 

replicate analysis. The dust leachate compositions analyzed by the CNWRA are listed in 

Table 2-1. DOE data on the split samples are reported in Bechtel SAIC Company, LLC 

(2007, 2004). The CNWRA data are for the leachates of the whole dust samples, whereas DOE 

analyses were conducted on leachates of sized fractions of the dust samples. 
2-4 

Figure 2-3. Calculated Deliquescence Relative Humidity Shown as Contour Lines of 

Different Salts and Salt Mixtures in the System NaCl–KCl–MgCl2 at 90 °C [194 °F]. The 

Solutions Are Saturated With the Indicated Salts. Also Shown for Some Compositions Is 

the Stoichiometric Chloride Concentration [Cl−

] (moles/kg H2O) of the Saturated 

Solutions. Figure Taken From Pabalan, et al. (2002). 

As shown in Table 2-1, calcium is present in all samples in significant amounts as compared to 

sodium, and all samples show a nitrate-to-chloride molar ratio of less than 1. All samples have 

a very low relative amount (<0.1 percent) of soluble salts. The U.S. Geological Survey stated 

that samples taken from the Yucca Mountain crest were leached by rainfall and, therefore, had 

low amounts of soluble salts. CNWRA powder x-ray diffraction data on four surface samples 

(Table 2-2) indicated the presence of feldspars and quartz; one sample had calcite as a 

major phase. 

2.2.2 Yucca Mountain Tunnel Data 

The U.S. Geological Survey staff collected samples from inside the tunnel by vacuuming dust 

from tunnel walls and by trapping dust using a cyclone. Vacuuming several square meters of 

surface yielded 250 to 400 g [8.8 to 14 oz] of dust per sample (Peterman, 2006). Splits of 

samples collected in the Exploratory Studies Facility portion of the tunnel in February 2002 were 

provided to CNWRA for analysis. CNWRA data on dust leachate compositions are listed in 

Table 2-3. DOE data on the split samples are reported in Bechtel SAIC Company, LLC 

(2007, 2004). The CNWRA data are for leachates of whole dust samples, whereas DOE 

analyses were conducted on leachates of sized fractions of the dust samples. All tunnel 

samples showed close to 1 percent soluble material. 

 
2-5 

Figure 2-4. Calculated Deliquescence Relative Humidity of Different Salts and Salt 

Mixtures in the System Na–K–Cl–NO3 as a Function of Temperature. The Red Curve Is 

Calculated for a Total Pressure of 89 kPa [12.9 psi], Which Is the Approximate 

Atmospheric Pressure at the Elevation of Yucca Mountain. The Region Above the Red 

Curve Is Inaccessible (i.e., There Would Be No Film of Water on the Surface of the 

Engineered Barrier). The Symbols Represent Experimental Data From Rard (2005, 2004). 

Table 2-1. Center for Nuclear Waste Regulatory Analyses Data on the Composition of 

Leachates of Dusts Sampled From the Surface of Yucca Crest (mg/L)*

Sample 

Number Ca2+ Mg2+ K+

 Na+ SiO2 Cl− Br− F− NO3

 SO4

2

NO

Cl

3





Molar 

Ratio 

030800 61.5 5.7 58.1 157.0 119.0 8.0 <1.49 2.6 10.2 33.5 0.73 

030803 28.1 5.2 19.4 <14.9 64.7 4.2 <1.49 3.1 4.1 10.2 0.56 

030805 76.0 12.3 55.7 15.1 63.1 7.9 <1.49 3.6 4.1 20.2 0.30 

030807 83.5 12.7 161.0 38.4 49.5 52.4 <1.49 3.2 21.6 67.5 0.24 

*DOE data on split samples are reported in Bechtel SAIC Company, LLC. “Environment on the Surfaces of the Drip 

Shield and Waste Package Outer Barrier.” ANL–EBS–MD–000001. Rev. 01. Las Vegas, Nevada: Bechtel SAIC 

Company, LLC. 2004. 

Table 2-4 gives CNWRA x-ray diffraction data for the tunnel samples, each of which has 

feldspar and silica minerals as major components. Calcite is also a major phase in one sample. 

2.2.3 Atmospheric (Cyclone) Dust Data 

DOE collected atmospheric dust samples using a Torit® cyclone, which was located near the 

South Portal until its relocation to a different site in March 2007. The cyclone was set up to 

Temperature ( oC)

50 100 150 200

0

10

20

30

40

50

60

70

80

90

100

T

Inaccessible

RH Region

KNO3 NaCl

KCl KNO3 + KCl NaCl

+ KCl NaNO3

NaCl + NaNO3

NaCl + KCl + KNO3

NaNO3 + KNO3

NaCl + NaNO3+ KNO3

Relative Humidity (%)

Curves — calculated

Symbols — data from Rard et al. (2004, 2005)

Temperature ( oC)

50 100 150 200

0

10

20

30

40

50

60

70

80

90

100

T

Inaccessible

RH Region

KNO3 NaCl

KCl KNO3 + KCl NaCl

+ KCl NaNO3

NaCl + NaNO3

NaCl + KCl + KNO3

NaNO3 + KNO3

NaCl + NaNO3+ KNO3

Relative Humidity (%)

Temperature ( oC)

50 100 150 200

0

10

20

30

40

50

60

70

80

90

100

T

Inaccessible

RH Region

KNO3 NaCl

KCl KNO3 + KCl NaCl

+ KCl NaNO3

NaCl + NaNO3

NaCl + KCl + KNO3

NaNO3 + KNO3

NaCl + NaNO3+ KNO3

Relative Humidity (%)

Curves — calculated

Symbols — data from Rard et al. (2004, 2005)
2-6 

Table 2-2. Mineralogical Composition of Surface Samples Based on Center for Nuclear 

Waste Regulatory Analyses Powder X-Ray Diffraction Analysis 

Mineral 

Sample Number 

030800 030803 030805 030807 

Albite, calcian, ordered 

(Na,Ca)Al(Si,Al)3Si3O8

Major Major 

Albite, ordered NaAlSi3O8 Major Minor Major Major 

Anorthite, ordered CaAl2Si2O8 Major — — — 

Anorthoclase, disordered (Na,K)(Si3Al)O8 — — Major Minor 

Anorthite, sodian, disordered 

(Ca,Na)(Si,Al)4O8

— — — Major 

Calcite CaCO3 Major — — — 

Microcline, intermediate KAlSi3O8 — Major Major — 

Augite, aluminian Ca(Mg,Fe3+, Al)(Si,Al)2O6 — Minor — — 

Diopside CaMg(SiO3) 2 — Minor — — 

Quartz SiO2 Major Major — Major 

Table 2-3. Composition of Leachates of Dusts Sampled From Inside the Tunnel at 

Yucca Mountain (mg/L)* 

Sample 

Number Ca2+ Mg2+ K+

 Na+ SiO2 Cl− Br− F− NO3

 SO4

2 PO43

NO

Cl

3





Molar 

Ratio 

574980 389 44.1 89.8 107 134 73.6 4.66 2.7 59.9 310 — 0.47 

574982 737 104 196 518 95.2 178 12.5 13.5 176 926 — 0.57 

574983 1010 98.5 253 710 60.6 364 15.7 20.2 194 1170 — 0.31 

574987 1390 58.3 247 622 180 366 44.5 21 120 1510 — 0.19 

*DOE data on split samples are reported in the following references: (i) Bechtel SAIC Company, LLC. “Analysis of 

Dust Deliquescence for FEP Screening.” ANL–EBS–MD–000074. Rev 01, Addendum 01. Las Vegas, Nevada: 

Bechtel SAIC Company, LLC. 2007; and (ii) Bechtel SAIC Company, LLC. “Environment on the Surfaces of the 

Drip Shield and Waste Package Outer Barrier.” ANL–EBS–MD–000001. Rev 01. Las Vegas, Nevada: Bechtel 

SAIC Company, LLC. 2004. 

Table 2-4. Mineralogical Composition of Tunnel Samples Based on Center for Nuclear 

Waste Regulatory Analyses Powder X-Ray Diffraction Analysis 

Mineral 

Sample Number 

574980 574982 574983 574987

Albite, calcian, ordered (Na,Ca)Al(Si,Al)3Si3O8 Major Major — —

Albite, ordered NaAlSi3O8 Major — Major —

Anorthite, sodian, ordered (Ca,Na)(Al,Si)2Si2O8 — — Major —

Anorthoclase, disordered (Na,K)(Si3Al)O8 — Major — Major 

Calcite CaCO3 — — Major 

Cristobalite SiO2 Major Major Major Major 

Microcline, intermediate KAlSi3O8 — — Major Minor 

Orthoclase KAlSi3O8 — Major Major Major 

Stellerite Ca2Al4Si14O36–14H2O — — Major —

Quartz SiO2 Major Major Major Major 
2-7 

exclude particles greater than 0.5 μm [2 H 10−5

 in] in size. Sampling was conducted from 

June 2005 until March 2007 approximately monthly and yielded approximately 1.5 g [0.053 oz] 

per sample (Peterman, 2006), although the exact amounts collected are not available in 

published literature. Splits of samples collected between March 2006 and March 2007 were 

provided to CNWRA for analysis. CNWRA leachate data are given in Table 2-5. DOE data are 

reported in Bryan (2006) and Bechtel SAIC Company, LLC (2007). Soluble fractions in the 

samples provided to CNWRA ranged from 8 to 13 percent. 

The CNWRA x-ray diffraction data listed in Table 2-6 indicate that all atmospheric samples 

contained calcite, quartz, and feldspars as major mineral components. 

2.2.4 Remote Dust Data 

The U.S. Geological Survey and the National Atmospheric Deposition Program have several 

dust sampling locations near the Yucca Mountain site. The dusts from these sampling locations 

are called remote dusts in this report. The remote dusts might be informative to the project, 

although they are not directly from the Yucca Mountain site. Remote dust sampling captured 

airborne dusts using specialized dust traps approximately 2 m [6.6 ft] above the ground surface, 

as Reheis (2003) described. The samples were collected semiannually, and the particular 

samples that the U.S. Geological Survey provided to CNWRA for analysis were collected in 

April 2004. Sample T3 was collected at a location a few miles north of Yucca Mountain, and 

T14 was collected in Crater Flat, west of the southern end of Yucca Mountain. These samples 

primarily contain dust from playas and alluvial fans in the area. Remote dusts contained 

approximately 10 percent soluble fraction based on CNWRA analyses. CNWRA analyses of the 

leachate of the combined T3 and T14 samples showed calcium as the primary cation (55 mole 

percent) and sulfate as the primary anion (69 mole percent), with sodium and chloride as the 

secondary cation and anion, respectively. The nitrate-to-chloride molar ratio of the combined T3 

and T14 leachate sample is 0.81. For comparison, data for regional precipitation (rain out) in 

the same area taken from the National Atmospheric Deposition Program are given in Table 2-7. 

The CNWRA x-ray diffraction analysis results for the remote samples T3 and T14 are listed in 

Table 2-8. The two remote samples have calcite, quartz, and feldspars as the major mineral 

components, similar to the other dust samples. 

2.3 Potential Composition of Brines Formed by Deliquescence of 

Salts Present in Yucca Mountain Dusts 

A useful indicator of the composition of salts that may deliquesce and form brines on the surface 

of drip shields and waste packages at the potential repository is the composition of salts present 

in dusts sampled from Yucca Mountain. However, there is no direct information regarding the 

identity of salt minerals present in dust samples taken from Yucca Mountain and the vicinity. 

The dust samples generally have salt amounts too low for phase identification using powder 

x-ray diffraction analysis. 

 
2-8 

Table 2-5. Center for Nuclear Waste Regulatory Analyses Data on the Composition of 

Leachates of Atmospheric Dusts Collected With the Torit Cyclone (mg/L)* 

Sample 

Number Ca2+ Mg2+ K+

 Na+ Si4+ Cl− F− NO3

 SO4

2 PO43

NO

Cl

3





Molar 

Ratio 

1037535 4960 1110 1510 4750 160 1850 30.3 3790 6480 31.3 1.17 

2043004 3940 835 1520 3280 203 1470 50.9 2610 4590 <8.31 1.02 

2043008 6920 1440 3270 4770 408 1550 47.6 4830 7170 53.9 1.78 

2043025 5750 1290 2210 5100 235 1510 38.7 5260 5960 11 1.99 

2043027 4630 986 1700 3910 245 1620 39.4 3690 5040 11.3 1.30 

2043031 4340 1150 2100 3490 260 1500 40.7 2970 4490 173 1.13 

2043033 5330 1190 1720 5270 275 1290 41.9 4260 8210 49.4 1.89 

2043037 3690 980 1410 4910 365 2860 21.3 2800 4770 17.6 0.56 

*DOE data on split samples are reported in the following references: (i) Bechtel SAIC Company, LLC. “Analysis of 

Dust Deliquescence for FEP Screening.” ANL–EBS–MD–000074. Rev. 01, Addendum 01. Las Vegas, Nevada: 

Bechtel SAIC Company, LLC. 2007; and (ii) Bryan, C. “Evolution of Waste Package Environments in a Repository 

at Yucca Mountain.” Nuclear Waste Technical Review Board Meeting, September 25–26, 2006, Las Vegas, 

Nevada. 2006. <http://www.nwtrb.gov/meetings/2006/sept/bryan.pdf>. (7 April 2008). 

Table 2-6. Mineralogical Composition of Atmospheric Samples Based on Center for 

Nuclear Waste Regulatory Analyses Powder X-Ray Diffraction Analysis 

Mineral 

Sample Number

1037535 2043004 2043008 2043025 2043027 2043031 2043033 2043037

Albite, calcian, ordered 

(Na,Ca)Al(Si,Al)3Si3O8 

Major Major — Major Major Major Major Major 

Albite, ordered 

NaAlSi3O8

Major — Major Major Major Major Major Major 

Anorthite, sodian, 

ordered 

(Ca,Na)(Al,Si)2Si2O8

— — Major Major — — Major —

Anorthite, sodian, 

intermediate 

(Ca,Na)(Si,Al)4O8

Minor — — — — Minor — —

Anorthoclase, 

disordered 

(Na,K)(Si3Al)O8

— — — — — — — Major 

Calcite 

CaCO3

Major Major Major Major Major Major Major Major 

Cristobalite 

SiO2

Minor — — — — — — —

Anorthite ordered 

 CaAl2Si2O8

— Major — — — — — —

Gismondine 

CaAl2Si2O8–4H2O 

— Major Major — — — — —

Quartz 

SiO2

Major Major Major Major Major Major Major Major 

 
2-9 

Table 2-7. Leachate Composition From the National Atmospheric Deposition Program 

Regional Precipitation (Rain Out). Data in mg/L 

Sample Number Ca2+ Mg2+ K+ Na+ NO3



Cl− SO4

2

NO

Cl

3





Molar 

Ratio 

NV00–2002 0.48 0.044 0.013 0.059 1.14 0.09 0.46 7.3 

NV00–2001 0.66 0.068 0.042 0.133 2.15 0.16 1.01 7.7 

NV00–2000 1.21 0.137 0.055 0.263 3.24 0.36 1.35 5.2 

Table 2-8. Mineralogical Composition of Two Remote Samples Based on Center for 

Nuclear Waste Regulatory Analyses Powder X-Ray Diffraction Analysis 

Mineral 

Sample Number 

T3 T14 

Albite, calcian, ordered (Na,Ca)Al(Si,Al)3Si3O8 Major Major 

Albite, ordered NaAlSi3O8 Major Major 

Anorthite ordered CaAl2Si2O8 Major Major 

Anorthite, sodian, intermediate (Ca,Na)(Si,Al)4O8 Minor 

Calcite CaCO3 Major Major 

Halite NaCl Minor 

Quartz SiO2 Major Major 

Thermodynamic simulations of the evaporation of dust leachates were conducted using 

StreamAnalyzer Version 2.0 (Gerbino, 2006). The software can be used to simulate aqueous 

chemical systems for temperatures up to 300 °C [572 °F], pressures up to 150 bar [2,205 psi], 

and ionic strengths as high as 30 molal. Higher ionic strengths—up to pure molten salts or pure 

acids—can be simulated with the software using a mixed-solvent electrolyte model (Wang, et 

al., 2002). Salt solubilities and aqueous solution water activities calculated using the 

mixed-solvent electrolyte model are consistent with experimental data (Gruszkiewicz, et al., 

2007), including those that Rard (2005, 2004) recently reported for the systems that contain 

NaCl, KNO3, and NaNO3. This latter model was applied in the calculations reported here to 

enable the evaporation simulations to extend to dry out conditions. The dust leachate 

compositions used as input to the StreamAnalyzer calculations were taken from Bechtel SAIC 

Company, LLC. (2004) and are tabulated in Table 2-9. 

The StreamAnalyzer evaporation simulations yielded six potential deliquescent salt 

assemblages, which are listed in Table 2-10. More than 80 percent of the dust leachate 

compositions resulted in either NaCl–KNO3 or NaCl–NaNO3–KNO3 salt mixtures. The dominant 

salt assemblages are NaCl–KNO3 and NaCl–NaNO3–KNO3 (neglecting KBr), and an infrequent 

salt assemblage resulting from the evaporation simulations includes a calcium nitrate salt. Note 

that nondeliquescent solid phases also are present in the mineral assemblage. For example, 

Figure 2-5 illustrates the mineral phases—deliquescent and nondeliquescent—that form upon 

evaporation of six dust leachate samples. 
2-10 

Table 2-9. Composition of Leachable Portion of Dust Samples Taken From Yucca 

Mountain Tunnels* 

Laboratory 

Sample 

Number 

Ca2+ Mg2+ Na+

 K+ Si4+ Cl− F− NO3

 SO4

2

Br−

ppm ppm ppm ppm ppm ppm ppm ppm ppm ppm 

C–186098 268 26.7 346 126 34.5 184 10 240 360 38 

C–186091 980 83.6 358 183 43.5 182 6 800 1,120 10 

C–186099 392 42 344 149 27.9 196 8 280 480 28 

C–186092 638 79.7 431 193 46.6 180 6 580 980 22 

C–186093 332 35.7 386 128 14.9 240 10 280 340 54 

C–186100 630 31.4 389 150 82.5 220 40 260 640 64 

C–186094 974 79.1 617 260 58.7 300 8 680 1,340 26 

C–186101 229 18.9 262 101 42.3 220 8 114 300 28 

C–186102 543 9.76 158 80 97.5 114 18 62 320 22 

C–186095 575 34.3 415 206 94.9 200 14 400 660 26 

C–186096 394 53.4 511 237 65 260 32 420 1,020 34 

C–186097 480 34.8 287 142 61.2 160 8 380 620 18 

C–186090 226 10.6 157 101 79.1 56 4 82 162 6 

C–186089 635 43.6 397 214 100 200 24 420 720 22 

C–186088 439 13.7 345 479 120 114 12 300 400 16 

C–186087 919 32.8 332 221 96.4 140 18 440 1,180 20 

C–186086 893 35.7 374 248 107 130 12 400 1,480 26 

C–186085 863 25.9 369 220 134 162 12 380 740 22 

C–186084 939 27.1 343 219 121 130 12 380 1,160 24 

C–186082 630 28.3 388 231 95.1 154 10 340 1,060 50 

C–186081 941 6.32 378 242 213 200 30 340 840 56 

C–186080 430 19.4 304 221 89.2 162 24 220 640 34 

C–186077 2,490 12.8 455 350 173 260 8 1,820 2,200 14 

C–203112 147 16.3 55.4 159 161 <24 8 <7 <32 <1.6 

C–203113 91.7 9.47 46.4 127 96.5 <24 4 10 <32 <1.6 

C–203114 119 12.6 66 174 144 <24 8 24 <32 <1.6 

C–203115† 270 34.3 115 193 169 76 18 220 220 <1.6 

C–203116† 466 49.3 124 181 76.6 98 4 400 360 4 

C–203117† 1,080 80.1 195 206 176 154 12 640 840 8 

C–203118 772 44.1 188 280 33.4 74 8 116 3,800 4 

C–203119 1,060 39.5 471 389 287 280 22 360 1,000 44 

C–203120 2,340 130 392 389 42.8 320 8 1,760 4,600 6 

C–203121† 246 39.6 262 196 164 86 20 240 320 4 

C–203122† 458 64.1 339 244 51.6 118 4 520 520 6 

C–203123† 1,010 99.6 556 345 99.2 170 6 1,000 1,060 12 

C–203124† 262 23.7 296 281 94.9 114 12 198 380 8 

C–203125† 770 50.8 425 303 57.3 260 4 540 720 10 

C–203126† 1,240 85.9 666 369 62.7 360 10 980 1,200 10 

 
2-11 

Table 2-9. Composition of Leachable Portion of Dust Samples Taken From Yucca 

Mountain Tunnels* (continued)

Laboratory 

Sample 

Number 

Ca2+ Mg2+ Na+

 K+ Si4+ Cl− F− NO3

 SO4

2

Br−

ppm ppm ppm ppm ppm ppm ppm ppm ppm ppm 

C–203127 335 49.4 265 339 87.3 128 10 220 440 10 

C–203130 994 65.4 349 303 248 170 18 340 880 20 

C–203131 1,260 52.4 461 333 181 168 22 480 1,180 14 

C–203132 1,030 54.7 480 275 130 166 22 500 1,220 16 

C–203133 248 23.4 222 234 75.4 96 6 146 320 6 

C–203134 806 38 292 220 131 170 12 300 740 10 

C–203135 1,190 51.4 408 260 159 220 18 440 1,140 14 

C–203136† 1,290 84.6 201 257 21.8 188 8 600 5,800 2 

C–203137† 1,180 64.1 564 313 135 360 22 540 1,400 42 

C–203138† 1,280 70.5 570 288 143 320 24 520 1,480 32 

C–203139 274 19.8 149 186 44.7 88 8 170 500 10 

C–203140 434 25.9 186 151 38.9 102 6 220 760 16 

C–203141 689 34.7 370 251 82.2 136 14 300 940 22 

C–203142 281 55.4 128 139 161 88 22 122 440 12 

C–203143 319 56.8 121 114 79.4 76 4 156 520 10 

C–203144 622 53.6 196 177 215 82 10 164 700 12 

C–203145 122 10.3 171 155 57.4 98 8 70 116 6 

C–203146 110 9.48 126 121 12.3 82 2 68 124 <1.6 

C–203147 305 19.8 187 173 60.2 166 6 150 280 2 

*Values for phosphate, arsenic, and lead are reported also in Bechtel SAIC Company, LLC. “Environment on the 

Surfaces of the Drip Shield and Waste Package Outer Barrier.” ANL–EBS–MD–000001. Rev. 01. Las Vegas, 

Nevada: Bechtel SAIC Company, LLC. 2004. However, these are not included in the evaporation simulations. 

Where “<x” values are listed, x is used as input in the StreamAnalyzer calculations. 

†The laboratory sample numbers identified in this table correspond to the following field sample numbers tabulated in 

Table 2-3: C–203115 (574980A), C–203116 (574980B), C–203117 (574980C), C–203121(574982A), C–203122 

(574982B), C–203123 (574982C), C–203124 (574983A), C–203125 (574983B), C–203126 (574983C), C–203136 

(574987A), C–203137 (574987B), C–203138 (574987C). 

Table 2-10. Salt Assemblages Derived From Evaporation of Dust Leachates With 

Compositions Listed in Table 2-9. The Evaporation Was Simulated Using 

StreamAnalyzer Version 2.0. 

Salt Assemblage 

Total 

Percentage of 

Occurrence 

Sample Number of 

Representative Leachate 

Composition 

NaCl–NaNO3–KNO3 43.9 C–186098 

NaCl–KNO3 40.3 C–186101 

KCl–KNO3 5.3 C–203112 

NaCl–KCl–KNO3 5.3 C–186088 

NaCl–NaNO3–Ca(NO3)2–Ca(NO3)24KNO3 3.5 C–203120 

NaCl–NaNO3–KNO3–Ca(NO3)24KNO3 1.7 C–203117 
2-12 

C–186098 

C–186101 

Figure 2-5. Mineral Precipitates Calculated to Form Upon Evaporation of Initially 

Dilute Dust Leachates. Decreasing Amount of H2O (Leftward on the X-Axis) 

Corresponds to Increasing Degree of Evaporation.
2-13 

C–203112 

C–186088 

Figure 2-5. Mineral Precipitates Calculated to Form Upon Evaporation of Initially 

Dilute Dust Leachates. Decreasing Amount of H2O (Leftward on the X-Axis) 

Corresponds to Increasing Degree of Evaporation. (continued)
2-14 

C–203120 

C–203117 

Figure 2-5. Mineral Precipitates Calculated to Form Upon Evaporation of Initially 

Dilute Dust Leachates. Decreasing Amount of H2O (Leftward on the X-Axis) 

Corresponds to Increasing Degree of Evaporation. (continued)
2-15 

To determine whether the resulting mineral assemblages can deliquesce and form brines at 

elevated temperatures, their mutual deliquescence relative humidity was calculated as a 

function of temperature using StreamAnalyzer. Not all 57 mineral assemblages derived from 

the evaporation simulations were used in this analysis. Instead, each key salt assemblage was 

represented by one leachate sample, as identified in Table 2-10. The mutual deliquescence 

relative humidity of each sample was derived by running isothermal, bubble point 

StreamAnalyzer calculations, incrementally adding small amounts of water to the system until 

an aqueous phase is first formed. The resulting activity of water when an aqueous phase first 

forms is taken as the mutual deliquescence relative humidity of the salt assemblage. 

The calculated mutual deliquescence relative humidity as a function of temperature for the six 

representative samples is listed in Table 2-11. The tabulated values indicate that all the 

resulting salt assemblages have very low deliquescence points, particularly at temperatures 

above 100 °C [212 °F]. To place the calculated mutual deliquescence relative humidity into 

perspective, those values can be compared to the in-drift relative humidity predicted for the 

potential Yucca Mountain repository, which are shown in Figure 2-1. A comparison of the 

mutual deliquescence relative humidity listed in Table 2-11 and the relative humidity in 

Figure 2-1 indicates that brines can potentially form on the waste package surface by 

deliquescence of salts when the waste package is relatively hot {>100 °C [>212 °F]}. 

The potential for these deliquescence brines to initiate localized corrosion will depend on the 

ratio of the molar concentration of chloride to corrosion-inhibiting oxyanions (nitrate, sulfate, 

bicarbonate, and carbonate), relative humidity, and temperature (Dunn, et al., 2005). The 

compositions of the deliquescence brines at the mutual deliquescence relative humidity of the 

salt assemblages were derived from the output file of the StreamAnalyzer runs described in the 

preceding paragraphs. Table 2-12 lists the calculated chloride-to-nitrate molar ratios for the six 

key salt assemblages that may deliquesce at temperatures above 100 °C [212 °F] at ambient 

pressure. When the temperature is at or above 120 °C [248 °F], the calculated chloride-tonitrate ratios of these deliquescence brines are low—less than 0.2 (or higher than 5 for the 

nitrate-to-chloride ratios) under environmental conditions in which deliquescence brine-induced 

corrosion could occur (Figure 2-1). When the temperature is below 120 °C [248 °F] and above 

the normal boiling temperature, the calculated chloride-to-nitrate ratio for the KCl–KNO3 salt 

assemblage (represented by sample C–203112) is high—13 to 18 [0.077 to 0.056] for the 

nitrate-to-chloride ratio. But the brine formed by deliquescence of this salt assemblage would 

have a ratio of chloride to localized corrosion inhibitors (nitrate + sulfate + bicarbonate + 

carbonate) less than 0.3 (or 3.6 for the inhibiting anions-to-chloride molar ratio). 

Figure 2-6 shows the calculated evolution of the brine chloride-to-nitrate molar ratio as 

more water is taken up by the salt assemblages at a temperature of 110 °C [230 °F]. The 

chloride-to-nitrate molar ratio initially remains constant as water is added and the soluble nitrate 

salts dissolve, but the ratio subsequently increases as the less soluble chloride salts start to 

dissolve. With further increases in the amount of water, the nitrate and chloride salts completely 

dissolve and the chloride-to-nitrate molar ratio becomes constant. Note that this report is 

concerned only with the chloride-to-nitrate molar ratio near the mutual deliquescence relative 

humidity because dilute solutions would occur only after the temperature has dropped below 

boiling points and the relative humidity has increased above the corresponding deliquescence 

relative humidity. 
2-16 

Table 2-11. Calculated Mutual Deliquescence Relative Humidity Percent As a Function of 

Temperature of Six Samples Representing Six Key Salt Assemblages* 

Temp 

(°C) 

Sample Number 

C–186098 C–186101 C–203112 C–186088 C–203120 C–203117 

25 64.4 65.1 70.6 65.0 4.5 37.5 

30 63.2 63.8 67.4 64.7 5.1 38.2 

40 60.0 61.0 57.2 63.5 6.4 39.1 

50 55.9 57.8 56.7 61.7 7.7 39.3 

60 51.3 54.2 56.1 59.2 9.0 38.8 

70 47.6 50.2 55.3 56.3 10.2 37.6 

80 43.7 45.9 54.4 53.1 11.1 35.8 

90 39.6 41.5 52.9 49.7 11.7 33.4 

100 35.4 36.8 51.2 46.0 12.1 30.6 

110 31.1 32.0 49.2 42.2 12.2 27.4 

120 26.7 27.0 47.1 38.1 12.4 23.9 

130 22.3 22.3 44.8 33.9 12.5 20.0 

140 18.3 18.3 42.4 29.5 12.5 16.4 

150 17.3 17.2 39.8 25.0 12.2 15.7 

*Each sample number represents a key salt assemblage: 

C–186098: NaCl–NaNO3–KNO3

C–186101: NaCl–KNO3

C–203112: KCl–KNO3

C–186088: NaCl–KCl–KNO3

C–203120: NaCl–NaNO3–Ca(NO3)2– Ca(NO3)24KNO3

C–203117: NaCl–NaNO3–KNO3– Ca(NO3)24KNO3

The results of these analyses suggest that deliquescence brines could form on the waste 

package surface at elevated temperatures. However, at temperatures at or above 120 °C 

[248 °F], the nitrate-to-chloride molar ratio of the brines are expected to be high (>5) under the 

potential drift conditions during the deliquescence period. The high nitrate-to-chloride molar 

ratios are consistent with the values Felker, et al. (2006) reported. At temperatures between 

100 and 120 °C [212 and 248 °F], the nitrate-to-chloride ratio may be low (less than 0.1), but the 

ratio of total inhibitors (nitrate + sulfate + bicarbonate + carbonate) to chloride is still higher than 

3.3 for the KCl–KNO3 salt assemblage. 

Section 2.2 noted that the U.S. Geological Survey provided split samples of dusts sampled from 

Yucca Mountain to CNWRA for replicate chemical analysis and that DOE data on leachate 

compositions are reported by Bryan (2006) and Bechtel SAIC Company, LLC (2007, 2004). To 

determine whether the CNWRA leachate data would yield different salt assemblages compared 

to DOE leachate compositions, additional StreamAnalyzer evaporation simulations were 

conducted. The CNWRA leachate data for Yucca Mountain tunnel samples 574980, 574982, 

574983, and 574987 listed in Table 2-3 were used as input to the calculations, which all yielded 

the salt assemblage NaCl–KNO3. This assemblage is consistent with the simulation results for 

DOE leachate samples C–203115, C–203121, and C–203136, but not for samples C–203116, 

C–203122, C–203123, C–203124, C–203125, C–203126, C–203137, and C–203138, which 

resulted in an NaCl–NaNO3–KNO3 assemblage, and sample C–203117, which yielded a salt 
2-17 

Table 2-12. Calculated Molar Concentration Ratio of Chloride-to-Nitrate Ions As a 

Function of Temperature of Six Samples Representing Six Key Salt Assemblages* 

Temp (°C) 

Sample Number 

C–186098 C–186101 C–203112 C–186088 C–203120 C–203117

25 0.56 0.60 0.15 (0.48)† 2.1 2.59 × 10–04 0.042 

30 0.45 0.50 3.4 (0.37) 1.9 3.24 × 10–04 0.040 

40 0.28 0.34 4.2 (0.15) 1.3 4.82 × 10–04 0.036 

50 0.17 0.24 4.8 (0.19) 0.85 6.77 × 10–04 0.032 

60 0.11 0.16 5.5 (0.22) 0.59 9.13 × 10–04 0.028 

70 0.081 0.12 6.4 (0.25) 0.44 1.19 × 10–03 0.025 

80 0.063 0.084 7.7 (0.28) 0.33 1.53 × 10–03 0.022 

90 0.051 0.063 9.7 (0.28) 0.26 1.94 × 10–03 0.021 

100 0.042 0.049 13 (0.28) 0.20 2.44 × 10–03 0.020 

110 0.036 0.039 18 (0.27) 0.16 3.07 × 10–03 0.020 

120 0.031 0.032 0.18 (0.18) 0.13 3.87 × 10–03 0.021 

130 0.028 0.028 0.16 (0.15) 0.11 4.89 × 10–03 0.022 

140 0.027 0.027 0.14 (0.14) 0.097 6.17 × 10–03 0.023 

150 0.032 0.033 0.13 (0.13) 0.088 7.75 × 10–03 0.028 

*Each sample number represents a key salt assemblage: 

C–186098: NaCl–NaNO3–KNO3

C–186101: NaCl–KNO3

C–203112: KCl–KNO3

C–186088: NaCl–KCl–KNO3

C–203120: NaCl–NaNO3–Ca(NO3)2– Ca(NO3)24KNO3

C–203117: NaCl–NaNO3–KNO3– Ca(NO3)24KNO3

†Values in parentheses are the concentration ratios of Cl!

/(NO3

!

–SO4

2!

–HCO3

!

–CO3

2!

) 

H2O added (moles)

10-5 10-4 10-3 10-2 10-1 100

Chloride-to-nitrate molar ratio

10-3

10-2

10-1

100

101

102

C-186098

C-186101

C-203112

C-186088

C-203112

C-203117

Figure 2-6. Evolution of the Nitrate-to-Chloride Molar Ratio as More Water Is Taken up by 

the Deliquescent Salt Assemblage at 110 °C [230 °F] 
2-18 

mixture of NaCl–NaNO3–KNO3– Ca(NO3)2KNO3. Although the resulting salt assemblages 

based on the CNWRA leachate chemistry are generally different than those based on the DOE 

leachate data, the conclusion stated in the previous paragraph still holds true. This is because 

brines formed by deliquescence of a NaCl–KNO3 salt mixture would have chloride-to-nitrate 

ratios much lower than the critical value that could lead to initiation of localized corrosion of 

Alloy 22. 

2.4 Effect of Capillary Retention by Dusts on the Corrosivity of 

Deliquescence Brines 

The deliquescent salts that may deposit on the waste package surface likely would be mixed 

with nondeliquescent rock dusts produced during potential repository construction. A DOE 

numerical analysis indicates that rock dusts would retain by capillarity any brine formed by 

deliquescence, thereby limiting brine contact with and potential corrosion of the waste package 

(Bechtel SAIC Company, LLC, 2005). Experiments discussed in this section (i) evaluated the 

effect of capillary retention by dusts on the corrosivity of deliquescence brines and (ii) assessed 

Alloy 22 corrosion in brines formed by deliquescence of salts mixed with rock dusts. 

2.4.1 Test Methods 

The experiments were conducted at 70 °C [158 °F] inside a chamber with controlled 

temperature and relative humidity using mixtures with varying mass ratios of NaCl to quartz or 

rock dust in contact with the metals of interest. Quartz powder, which was sieved to <250 μm 

[<0.01 in] and acid washed to remove soluble impurities, was used for comparison with the rock 

dust. Simulated rock dust was made by crushing specimens of Topopah Spring tuff, which is 

the potential repository host rock at Yucca Mountain, and by sieving the crushed material to 

<74 μm [<0.003 in]. To determine the composition of the soluble components of the rock dust, 

5-g [0.18-oz] samples of the rock dust were leached in 50 mL [1.7 fluid oz] of deionized water at 

25 and 70 °C [77 and 158 °F] for various durations. The leachates were analyzed for soluble 

composition using inductively coupled plasma emission spectrometry for cations and ion 

chromatography for anions. The leachate compositions, which are shown in Table 2-13, 

indicate that the soluble constituents of the rock dust are primarily sodium salts of chloride, 

fluoride, nitrate, and sulfate. The molar ratio of chloride and sulfate to chloride is less than or 

equal to 0.30. The charge imbalance between cations and anions shown in Table 2-13 could be 

due to analytical error and to the lack of data on carbonate species, which was not analyzed in 

this study. 

The corrosivity of the deliquescence brines was monitored using coupled multielectrode array 

sensor probes (Yang, et al., 2002a). Each probe has 16 electrodes of the same material with a 

diameter of 1 mm [39 mil]. The electrode material was either carbon steel or Alloy 22. The 

carbon steel electrodes were used for their high sensitivity to the corrosivity of the brine. The 

deliquescent salts used were NaCl and a mixture of NaCl, KNO3, and NaNO3. Figure 2-7 shows 

the schematic of the test setup. The tests were conducted by setting four probes vertically in 

the humidity chamber. The salt-dust mixtures, each with a total mass of 10 g [0.35 oz], were 

placed inside Tygon sleeves above the probes. The relative humidity inside the chamber was 

raised stepwise from 20 percent to above the deliquescence point of NaCl, which is 75 percent 

at 70 °C [158 °F] (Yang, et al., 2002b). Four tests were conducted at 70 °C [158 °F] using 

carbon steel probes and mixtures of NaCl–dust or NaCl–quartz. These four tests were 

designed to evaluate the effect of capillary retention by dusts on the corrosivity of deliquescence 
2-19 

Table 2-13. Chemical Compositions of Dust Leachate (ppm) 

Leaching 

Time 

(Days) Cl−

 F− NO3

−

 SO4

2−

 Na+

Molar Ratio 

[Na ]

[Cl ] [NO ] [F ] 2[SO ] 3 4

2



    

Molar Ratio 

[NO ] [SO ]

[Cl ]

3 4

 2



 

5 39.3 12 <2 18.4 376 8 0.19 

12 37.8 13.2 <2 18.0 348 7 0.19 

52 35.4 13.9 <2 17.5 352 7 0.20 

Leaching at 70 °C [158 °F] 

5 55.4 19.6 12.0 19.0 407 6 0.25 

14 39.3 20.5 9.08 18.4 413 7 0.30 

Figure 2-7. Schematic of Test Setup Used in Studying the Effect of Capillary Retention 

by Dusts on the Corrosivity of Deliquescence Brines

brines. An additional test was conducted at 125 °C [257 °F] using both Alloy 22 and carbon 

steel probes to assess the corrosion of Alloy 22 material relative to that of carbon steel in brines 

formed by the deliquescence of NaCl–KNO3–NaNO3 salts mixed with rock dusts. 

2.4.2 Test Results 

2.4.2.1 Test 1—NaCl–Quartz Mixture at 70 °C [158 °F] 

Figure 2-8 shows the maximum corrosion current measured from the four carbon steel probes 

at 70 °C [158 °F] with the following solid phase compositions: (i) 100 percent NaCl on Probe 1, 

(ii) 30 percent NaCl–70 percent quartz on Probe 2, (iii) 10 percent NaCl–90 percent quartz on 

Tygon

sleeve 

Multielectrode Data 

Acquisition System 

Salt/Dust 

mixture 

Multielectrode 

array 

sensor 

probe 

Humidity Chamber 
2-20 

Figure 2-8. Maximum Corrosion Current Measured From Test 1 Carbon Steel Probes. 

Test 1 Was Conducted at 70 °C [158 °F] With the Following Solid Phase Compositions: 

(i) 100 Percent NaCl on Probe 1, (ii) 30 Percent NaCl–70 Percent Quartz on Probe 2, 

(iii) 10 Percent NaCl–90 Percent Quartz on Probe 3, and (iv) 1 Percent NaCl–99 Percent 

Quartz on Probe 4.

Probe 3, and (iv) 1 percent NaCl–99 percent quartz on Probe 4. The relative humidity was 

raised stepwise from 20 percent to 40 percent, 55 percent, 68 percent, and 80 percent. 

At relative humidities of 20 percent, 40 percent, and 55 percent, the carbon steel corrosion 

current was below 10!9

 A, which is the background noise level, indicating that the system was 

dry at those relative humidities. However, at a relative humidity of 68 percent, which is below 

the deliquescence relative humidity of NaCl (75 percent), the corrosion currents increased by 

1 to 2 orders of magnitude, which was probably due to the adsorption of water molecules on the 

salt crystals (Yang, et al., 2002b). At a relative humidity of 80 percent, brine was formed and 

overflowed from Probes 1, 2, and 3. At 80 percent relative humidity, the corrosion current from 

Probes 1, 2, and 3 decayed with time, whereas the corrosion current from Probe 4 remained 

high. The decrease in corrosion current observed for Probes 1, 2, and 3 was likely due to the 

brine solution slowing down oxygen transport and the salting out effect, which decreases the 

oxygen concentration in the solution. The brine solution was not present in Probe 4, so that 

oxygen transport was not a limiting factor for the carbon steel corrosion process. No significant 

difference was observed in the NaCl–quartz tests with 100, 30, and 10 percent NaCl 

(Probes 1, 2, and 3, respectively). 

Figure 2-9 shows posttest photographs of the Test 1 carbon steel probes. Corrosion was 

evident on each carbon steel probe indicating brine contacted the metal surface, which is 

consistent with the observed corrosion current in Figure 2-8. 

0 10 20 30

Time (days)

10-10

10-9

10-8

10-7

10-6

10-5

10-4

Carbon steel corrosion current (

A)

-40

0

40

80

Relative humidity (%)

100% NaCl, Probe 1 

30% NaCl and 70% Quartz, Probe 2

10% NaCl and 90% Quartz, Probe 3

1% NaCl and 99% Quartz, Probe 4

40%

55%

68%

80%

20%

1% NaCl

99% Quartz
2-21 

Figure 2-9. Posttest Photographs of the Test 1 Carbon Steel Probes

2.4.2.2 Test 2—NaCl–Dust Mixture at 70 °C [158 °F] 

Because no significant difference in corrosion current was observed in the NaCl–quartz tests 

with 100, 30, and 10 percent NaCl, the tests involving NaCl–dust mixtures used smaller relative 

amounts of NaCl. Figure 2-10(a) shows the maximum corrosion current measured from 

the four carbon steel probes with the following solid phase compositions: (i) 10 percent 

NaCl-90 percent dust on Probe 1, (ii) 1 percent NaCl–99 percent dust on Probe 2, 

(iii) 0.4 percent NaCl–99.6 percent dust on Probe 3, and (iv) 100 percent dust on Probe 4. The 

relative humidities used in the tests were the same as those in Test 1. 

The results in Figure 2-10(a) show that the corrosion current increased above background level 

at a relative humidity of 55 percent, even for the probe with no salt added (100 percent dust), 

indicating an aqueous solution formed at a relative humidity of 55 percent. This relative 

humidity is much lower than the NaCl deliquescence point, and the brine formation might have 

been caused by the lowering of the deliquescence relative humidity by the small amount of 

nitrate in the dust (Table 2-13). The corrosion rate of carbon steel, which was calculated from 

the measured corrosion current by assuming uniform corrosion on each electrode, is shown in 

Figure 2-10(b). For all mixtures, including the mixture with a dust-to-NaCl weight ratio of 249 

(Probe 3 with 0.4 percent NaCl–99.6 percent dust) and pure dust, the carbon steel corrosion 

rate increased above 1 μm/yr [0.04 mil/yr] once relative humidity reached 55 percent. This 

result suggests the brine that formed by deliquescence at 55 percent relative humidity would be 

sufficient to support ionic transport on the carbon steel surface and the leachable salt in pure 

rock dust would be sufficient to initiate corrosion of carbon steel. At a relative humidity of 

80 percent, the current decreased with time for Probes 2, 3, and 4. The small amount of nitrate 

present in the dust (Table 2-13) could have leached into the solution during the test and acted 

as a corrosion inhibitor, causing the observed decrease in corrosion current. 

Figure 2-11 shows posttest photographs of the probes. No visible liquid formed even at a 

relative humidity of 80 percent, but the solid mixtures were damp, indicating absorption of liquid 
2-22 

Figure 2-10. (a) Measured Maximum Corrosion Current and (b) Calculated Corrosion 

Rate for Test 2 Carbon Steel Probes. Test 2 Was Conducted at 70 °C [158 °F] Using the 

Following Solid Phase Compositions: (i) 10 Percent NaCl–90 Percent Dust On Probe 1, 

(ii) 1 Percent NaCl–99 Percent Dust on Probe 2, (iii) 0.4 Percent NaCl–99.6 Percent Dust 

on Probe 3, and (iv) 100 Percent Dust on Probe 4. 

0 10 50 60

Time (days)

1E-010

1E-009

1E-008

1E-007

1E-006

1E-005

0.0001

Carbon steel corrosion current (A)

0

20

40

60

80

Relative humidity (%)

10% NaCl and 90% Dust, Probe 1

1% NaCl and 99% Dust, Probe 2

0.4% NaCl and 99.6% Dust, Probe 3

0% NaCl and 100% Dust, Probe 4

40%

55%

68%

80%

20%

10% NaCl

90% Dust

100% Dust

(a)

0 10 50 60 70

Time (days)

10-1

100

101

102

103

104

105

Carbon steel corrosion rate (μm/yr)

0

20

40

60

80

Relative humidity (%)

10% NaCl and 90% Dust, Probe 1

1% NaCl and 99% Dust, Probe 2

0.4% NaCl and 99.6% Dust, Probe 3

0% NaCl and 100% Dust, Probe 4

40%

55%

68%

80%

20%

10% NaCl

90% Dust

100% Dust

(b)
2-23 

Figure 2-11. Posttest Photographs of the Carbon Steel Probes From Test 2 

by dust. Corrosion was observed on each carbon steel probe even for the 100 percent dust 

system, but the number of corrosion sites decreased with decreasing percentage of NaCl in 

the mixture. 

2.4.2.3 Test 3—NaCl–Quartz and NaCl–Dust Mixtures at 70 °C [158 °F] 

To compare directly the results for pure dust, pure quartz, and mixtures with a very small 

amount of NaCl, Test 3 was conducted with the following solid phase compositions: 

(i) 0.4 percent NaCl–99.6 percent quartz on Probe 1, (ii) 100 percent quartz on Probe 2, 

(iii) 0.4 percent NaCl–99.6 percent dust on Probe 3, and (iv) 100 percent dust on Probe 4. 

Figure 2-12 shows the measured corrosion current from all probes. The measured currents 

indicate that no deliquescence brine formed at all relative humidity in the test with 100 percent 

quartz, whereas brine formed in the test with 100 percent rock dust. Consistent with the 

observations in Test 2, deliquescence brine formed at a lower relative humidity (55 percent) in 

the NaCl–dust mixture compared to the NaCl–quartz mixture. At a relative humidity of 

80 percent, for both the 0.4 percent NaCl–99.6 percent dust and 100 percent dust tests, the 

measured current reached a peak then decayed with time. In contrast, the current from the test 

with NaCl–quartz continued to increase and resulted in higher current compared to the tests 

involving a dust component. 

The results suggest that the presence of dust reduced the corrosion current by a capillary 

retention effect and/or by the leaching of a corrosion-inhibiting species from the dust. However, 

the results also indicate that the presence of dust did not prevent brine from contacting the 

metal surface. 

 

10% NaCl 

90% Dust 1% NaCl 

99% Dust 

0.1% NaCl 

99.9% Dust 100% Dust 
2-24 

Figure 2-12. Maximum Corrosion Current Measured From the Test 3 Carbon Steel 

Probes. Test 3 Was Conducted at 70 °C [158 °F] With the Following Solid Phase 

Compositions: (i) 0.4 Percent NaCl–99.6 Percent Quartz on Probe 1, (ii) 100 Percent 

Quartz on Probe 2, (iii) 0.4 Percent NaCl–99.6 Percent Dust on Probe 3, and 

(iv) 100 Percent Dust on Probe 4.

2.4.2.4 Test 4—Effect of Dust Particle Sizes 

To understand the effect of particle size on the capability of dusts to retain moisture by 

capillary retention, Test 4 was conducted with 1 percent NaCl–99 percent dust mixtures and the 

following dust particle size ranges: (i) >149 μm [>5.87 mil] on Probe 1, (ii) 149–74 μm 

[5.87–2.9 mil] on Probe 2, (iii) 74–44 μm [2.9–1.7 mil] on Probe 3, and (iv) <44 μm [<1.7 mil] on 

Probe 4. Figure 2-13 shows the corrosion current from all four probes. Consistent with 

previous tests, brines formed at a relative humidity of 55 percent. Except for the largest dust 

particles {>149 μm [>5.87 mil]}, no significant difference in corrosion current between the tests 

was observed, which suggests a negligible effect of particle size on capillary retention of brine. 

Figure 2-13 shows that large size particles delayed the current increase. This delay could result 

from the smaller surface area of large particles to absorb moisture compared to other smaller 

particles less than 149 μm [5.87 mil]. Approximation showed that the total surface areas of 

particles from Probes 1, 2, and 3 were at the same order; however, the total surface area of 

particles from Probe 1 was one order of magnitude smaller than other smaller particles. 

2.4.2.5 Test 5—Alloy 22 and Carbon Steel Corrosion in NaCl–NaNO3–KNO3 Dust 

Mixture at 125 °C [257 °F] 

This test assessed the corrosion at high temperature of Alloy 22 in comparison with carbon 

steel. The test used a mixture of NaCl, NaNO3, and KNO3 salts to allow deliquescence to occur 

above 100 °C [212 °F]. Because of the limited operational temperature range of the humidity 

chamber, a temperature of 125 °C [257 °F] was selected for the test. At this temperature, the 

0 10 20 30 40 50

Time (days)

1E-010

1E-009

1E-008

1E-007

1E-006

1E-005

Carbon steel corrosion current (A)

-40

0

40

80

Relative humidity (%)

0.4% NaCl and 99.6% Quartz, Probe 1 

100% Quartz, Probe 2

0.4% NaCl and 99.6% Dust, Probe 3

100% Dust, Probe 4

40%

55%

68%

80%

20%

100% Quartz

100% Dust

0.4% NaCl 

99.6% Dust

0.4% NaCl

99.6% Quartz
2-25 

0 10 20 30 40

Time (days)

1E-010

1E-009

1E-008

1E-007

1E-006

1E-005

Carbon steel corrosion current (

A)

-40

0

40

80

Relative humidity (%)

Dust particles > 149 μm, Probe 1

Dust particles: 14974 μm, Probe 2

Dust particles: 7444 μm, Probe 3

Dust particles < 44 μm, Probe 4

40%

55%

68%

80%

20%

Figure 2-13. Maximum Corrosion Current Measured From the Test 4 Carbon Steel 

Probes. Test 4 Was Conducted at 70 °C [158 °F] Using Mixtures of 1 Percent 

NaCl–99 Percent Dust. The Dust Had the Following Particle Size Ranges: (i) >149 μm 

[5.87 mil] on Probe 1, (ii) 149–74 μm [5.87–2.9 mil] on Probe 2, (iii) 74–44 μm [2.9–1.7 mil] 

on Probe 3, and (iv) <44 μm [<1.7 mil] on Probe 4. 

mutual deliquescence relative humidity for a NaCl–NaNO3–KNO3 mixture calculated using 

StreamAnalyzer Version 2.0 is 24.5 percent (Figure 2-4), and the calculated composition of the 

deliquescence brine is listed in Table 2-14. The calculated brine composition was used to 

estimate the weight ratios of reagent grade NaCl, NaNO3, and KNO3 salts that were used in the 

test (column 4 in Table 2-14). 

Four probes were used in the test—two with Alloy 22 electrodes and another two with carbon 

steel electrodes. Two weight ratios of dust to salt mixtures (3:7 and 5:5) were used, which 

yielded the following four probes: (i) 30 percent dust–70 percent salt mixture on carbon steel 

Probe 1, (ii) 50 percent dust–50 percent salt mixture on carbon steel Probe 2, (iii) 30 percent 

dust–70 percent salt mixture on Alloy 22 Probe 3, and (iv) 50 percent dust–50 percent salt 

mixture on Alloy 22 Probe 4. 

Figure 2-14(a) shows the corrosion current measured from the four probes. Data from the 

day-15 to day-40 portion of the test are not shown in this figure, because the humidity chamber 

did not function well during that period. The results show that under the same test conditions, 

the carbon steel corrosion current was much higher than that measured from Alloy 22 probes. 

The corrosion rate calculated for carbon steel was about 20 times that of Alloy 22. Because 

the coupled multielectrode probe measures the rate of nonuniform corrosion, which 
2-26 

Table 2-14. Chemical Composition of Brine Formed by Deliquescence of a NaCl–NaNO3–

KNO3 Salt Mixture at a Temperature of 125 °C [257 °F] and Weight Ratio of Reagent Grade 

Salts in Salt Mixture Used in Test 5 

Salt Moles Per kg H2O Mass (g) Per kg H2O Weight Ratio in Salt Mixture 

NaCl 2.69 157 0.018 

NaNO3 39.2 3,332 0.381 

KNO3 52.0 5,252 0.601 

includes localized corrosion, the low measured currents from the Alloy 22 probe also 

indicate that the penetration rate due to localized corrosion of Alloy 22 in the mixture of dust and 

NaCl–KNO3–NaNO3 was very low at 125 C [257 F]. 

Figure 2-14(b) shows the posttest probes. No localized corrosion was observed on Alloy 22 

with the unaided eye and under an optical microscope with H 20 magnification, which is 

consistent with the low nonuniform corrosion current measured with the probe. Table 2-13 

shows that the molar ratio of chloride plus sulfate to chloride in the dust leachate is less than or 

equal to 0.30. However, previous tests in bulk solution at 110 C [230 F] (Dunn, et al., 2005) 

indicated that localized corrosion is inhibited when the nitrate to chloride molar ratio is at least 

0.3. Thus, the lack of evident localized corrosion on the Alloy 22 specimen from this test could 

be due to the short duration of the test and the small amount of brine formed during the test. 

2.5 Summary and Conclusions 

Available information on the geochemistry of dust samples taken from Yucca Mountain and the 

vicinity was reviewed. Dusts contain relatively small fractions of soluble salts, ranging from 

10 percent as the maximum in remote locations to less than 0.1 percent in Yucca Mountain 

surface dusts. Leachates typically contain sodium, potassium, and calcium as major cations 

and chloride, nitrate, and sulfate as major anions. Nitrate-to-chloride ratios exceeded 0.1 for all 

samples. Feldspars, quartz, and calcite were among the minerals typically found in 

Yucca Mountain dusts. 

The possible assemblage of deliquescent salts present in Yucca Mountain dusts was 

determined by simulating the evaporation of aqueous solutions derived from leaching dust 

samples from Yucca Mountain. The results of the thermodynamic simulations indicated that the 

most likely salt assemblages comprise NaCl–NaNO3–KNO3 and NaCl–KNO3 mixtures, with 

relatively uncommon occurrence of assemblages containing calcium nitrate salts. 

Thermodynamic calculations also were conducted to determine the mutual deliquescence 

relative humidity of the salt assemblages and the chemistry of the brines that could form by 

deliquescence. The results indicate that the mutual deliquescence relative humidity of the salt 

assemblages is low, particularly at temperatures above 100 °C [212 °F], such that if these salts 

were present in the potential repository, brines would likely form by salt deliquescence even if 

the waste packages were relatively hot. However, the brines that are likely to form at 

temperatures above 120 °C [248 °F] would have molar ratios of nitrate to chloride higher than 5. 

At temperatures between 100 and 120 °C [212 and 248 °F], the nitrate-to-chloride molar ratio 

may be low (less than 0.1), but the molar ratio of total inhibitors (nitrate + sulfate + bicarbonate 

+ carbonate) to chloride would still be higher than 3.3 for the KCl–KNO3 salt assemblage. 

 
2-27 

Figure 2-14. (a) Maximum Corrosion Current Measured From Test 5 Probes. Test 5 Was 

Conducted at 125 °C [257 °F] and Relative Humidity of 25 Percent Using the Following 

Solid Phase Compositions and Probe Types: (i) 30 Percent Dust–70 Percent Salt Mixture 

on Carbon Steel Probe 1, (ii) 50 Percent Dust–50 Percent Salt Mixture on Carbon Steel 

Probe 2, (iii) 30 Percent Dust–70 Percent Salt Mixture on Alloy 22 Probe 3, and 

(iv) 50 Percent Dust–50 Percent Salt Mixture on Alloy 22 Probe 4. (b) Posttest 

Photographs of the Probes. 

 

0 10 40 50

Time (days)

1E-010

1E-009

1E-008

1E-007

1E-006

1E-005

Corrosion current (

A)

Carbon steel, 3 g dust and 7 g salt mixture, Probe 1

Carbon steel, 5 g dust and 5 g salt mixture, Probe 2

Alloy 22, 3 g dust and 7 g salt mixture, Probe 3

Alloy 22, 5 g dust and 5 g salt mixture, Probe 4

Alloy 22

Carbon steel

(a) 

(b) 
2-28 

The effect of capillary retention by dusts on the corrosivity of deliquescence brines was 

evaluated. The experiments were conducted at 70 °C [158 °F] in a temperature-and-humidity 

controlled chamber using solid mixtures with varying ratios of NaCl to quartz or rock dust. The 

results indicate that nondeliquescent quartz or rock dust reduced the amount of deliquescence 

brine that contacted the metal surface. However, even at a dust-to-NaCl weight ratio of 249, a 

brine solution still contacted and caused the corrosion of the carbon steel metal surface. No 

significant effect of dust particle size on capillary retention was observed, except for the largest 

dust particles {>149 μm [>5.87 mil]} used in the test. 

Additional experiments assessed the corrosion of Alloy 22 in brines formed by deliquescence of 

salts mixed with rock dust. The tests were conducted using Alloy 22 coupled multielectrode 

probes at a temperature of 125 °C [257 °F] and at humidity near the deliquescence point of a 

NaCl–KNO3–NaNO3 salt mixture. The results indicate that the corrosion rate of Alloy 22 was 

much lower than that of carbon steel and no visible localized corrosion was observed on 

Alloy 22 probes. Because the coupled multielectrode probe measures the rate of nonuniform 

corrosion, which includes localized corrosion, the low measured currents from the Alloy 22 

probe also indicate that the penetration rate due to localized corrosion for Alloy 22 in the mixture 

of dust and NaCl–KNO3–NaNO3 was low at 125 C [257 F] under the test conditions. 

2.6 References 

Bechtel SAIC Company, LLC. “Analysis of Dust Deliquescence for FEP Screening.” 

ANL–EBS–MD–000074. Rev. 01, Addendum 01. Las Vegas, Nevada: Bechtel SAIC 

Company, LLC. 2007. 

——. “Analysis of Dust Deliquescence for FEP Screening.” ANL–EBS–MD–000074. Rev. 01. 

Las Vegas, Nevada: Bechtel SAIC Company, LLC. 2005. 

——. “Environment on the Surfaces of the Drip Shield and Waste Package Outer Barrier.” 

ANL–EBS–MD–000001. Rev. 01. Las Vegas, Nevada: Bechtel SAIC Company, LLC. 2004. 

Bryan, C. “Evolution of Waste Package Environments in a Repository at Yucca Mountain.” 

Nuclear Waste Technical Review Board Meeting. September 25–26, 2006, Las Vegas, 

Nevada. 2006. <http://www.nwtrb.gov/meetings/2006/sept/bryan.pdf>. (7 April 2008). 

Dunn, D.S., O. Pensado, Y.-M. Pan, R.T. Pabalan, L. Yang, X. He, and K.T. Chiang. “Passive 

and Localized Corrosion of Alloy 22—Modeling and Experiments.” CNWRA 2005-02. 

San Antonio, Texas: CNWRA. 2005. 

Felker, S., P.D. Hailey, T. Lian, K.J. Staggs, and G. Gdowski. “Alloy 22 Localized Corrosion 

Susceptibility in Aqueous Solutions of Chloride and Nitrate Salts of Sodium and Potassium at 

110–150 °C.” UCRL–TR–218195. Livermore, California: Lawrence Livermore National 

Laboratory. 2006. 

Gerbino, A. “A Guide for Using the OLI Analyzers.” Morris Plains, New Jersey: OLI Systems, 

Inc. 2006. <http://support.olisystems.com/Documents/Manuals/TricksOfTheTrade.pdf>. 

(7 April 2008). 

Gruszkiewicz, M.S., D.A. Palmer, R.D. Springer, P. Wang, and A. Anderko. “Phase Behavior of 

Aqueous Na–K–Mg–Ca–Cl–NO3 Mixtures: Isopiestic Measurements and Thermodynamic 

Modeling.” Journal of Solution Chemistry. Vol. 36. pp. 723–765. 2007. 
2-29 

Pabalan, R.T., L. Yang, and L. Browning. “Deliquescence Behavior of Multicomponent Salts: 

Effects on the Drip Shield and Waste Package Chemical Environment of the Proposed Nuclear 

Waste Repository at Yucca Mountain, Nevada.” P. McGrail and G.A. Cragnolino, eds. 

Scientific Basis for Nuclear Waste Management XXV. Symposium Proceedings 713. 

Warrendale, Pennsylvania: Materials Research Society. pp. 37–44. 2002. 

Pensado, O. “Corrosion Model to Support Total System Performance Assessments.” Nuclear 

Waste Technical Review Board Meeting, September 25–26, 2006, Las Vegas, Nevada. 2006. 

<http://www.nwtrb.gov/meetings/2006/sept/pensado.pdf>. (7 April 2008). 

Peterman, Z. “Effects of Temperature on the Compositions of Soluble Salts in Dust at Yucca 

Mountain.” Nuclear Waste Technical Review Board Meeting, January 16, 2008, Las Vegas, 

Nevada. 2006. <http://www.nwtrb.gov/meetings/2008/jan/peterman.pdf>. (7 April 2008). 

Rard, J.A. “Results from Boiling Temperature Measurements for Saturated Solutions in the 

Systems NaCl+Ca(NO3)2+H2O, NaNO3+KNO3+H2O, and NaCl+KNO3+H2O, and Dry Out 

Temperatures for NaCl+NaNO3+KNO3+Ca(NO3)2+H2O.” UCRLTR–217415. 

Livermore, California: Lawrence Livermore National Laboratory. 2005. 

–––––. “Results from Boiling Temperature Measurements for Saturated Solutions in the 

Systems NaCl+KNO3+H2O, NaNO3+KNO3+H2O, and NaCl+NaNO3+KNO3+H2O.” 

UCRL–TR–207054. Livermore, California: Lawrence Livermore National Laboratory. 2004. 

Reheis, M.C. “Dust Deposition in Nevada, California, and Utah, 1984-2002.” U.S. Geological 

Survey Open-File Report 03-138. 2003. 

Wang, P., A. Anderko, and R.D. Young. “A Speciation-Based Model for Mixed-Solvent 

Electrolyte Systems.” Fluid Phase Equilibria. Vol. 203. pp. 141–176. 2002. 

Yang, L., N. Sridhar, and G. Cragnolino. “Comparison of Localized Corrosion of Fe–Ni–Cr–Mo 

Alloys in Chloride Solutions Using a Coupled Multielectrode Array Sensor.” CORROSION/2002. 

Paper No. 02545. Houston, Texas: NACE International. 2002a. 

Yang, L., R.T. Pabalan, and L. Browning. “Experimental Determination of the Deliquescence 

Relative Humidity and Conductivity of Multicomponent Salt Mixtures.” Scientific Basis for 

Nuclear Waste Management XXV Symposium. Proceedings 713. P. McGrail and 

G.A. Cragnolino, eds. Warrendale, Pennsylvania: Materials Research Society. pp. 135–142. 

2002b. 
3-1 

3 CORROSIVITY OF DELIQUESCENCE BRINES TO ALLOY 22 AT 

ELEVATED TEMPERATURES 

General corrosion rates of Alloy 22 exposed to large quantities of concentrated brines 

containing NaCl, NaNO3, and KNO3 in closed systems have been reported in the literature. To 

obtain the general corrosion rates, measurements were conducted at temperatures above 

100 C [212 F] in autoclaves under pressurized conditions. 

Because the potential repository drift would not be sealed (Sandia National Laboratory, 2007a) 

and the water vapor pressure could not be higher than the prevailing ambient pressure, the 

general corrosion rate of Alloy 22 exposed to large quantities of brines containing NaCl, NaNO3, 

and KNO3 were also measured in open systems at the Center for Nuclear Waste Regulatory 

Analyses (CNWRA®). Under such conditions, the brine would have a similar composition as the 

brines that may be formed by deliquescence of the salt mixtures containing NaCl, NaNO3, and 

KNO3. Whereas some of the data from CNWRA have been reported, additional data have been 

recently obtained. This chapter reviews the Alloy 22 corrosion rates that were reported in the 

literature and presents data recently obtained in the open systems. 

3.1 Corrosion Rate of Alloy 22 in Nitrate–Chloride Salt Systems 

Measured in Closed Systems 

Orme, et al. (2004) used the weight loss method to measure the corrosion rate of Alloy 22 with 

foil specimens exposed to both the vapor and the liquid phases of a system containing a 

nitrate–chloride brine inside an autoclave at temperatures between 120 and 220 C [248 and 

428 F]. The total concentration of the salts in the brines varied from 6.7 to 9.6 molal, and the 

nitrate-to-chloride ratio varied from 0.05 to 6.7. The autoclaves were made of Alloy 276, and the 

specimens were assembled on metal racks that were fabricated from Alloy 22. Because the 

brines were relatively dilute {purposely controlled so that the brines were saturated at 90 C

[194 F]}, the experiments were conducted under pressurized conditions to maintain liquid 

phase in addition to the vapor phase. Because the autoclave was sealed, the oxygen initially in 

the air space of the autoclave might have been partially consumed by the reaction with the 

metal surfaces inside the autoclave after the autoclave was heated up to test temperature. 

Therefore, the corrosion potential was probably low, even though the system was not purged 

with an inert gas at the start of the test. The test durations varied from 120 to 240 days. 

The measured corrosion rates varied from !45 to 150 nm/yr [!0.0018 to 0.0059 mil/yr]. The 

negative corrosion rates were due to the difficulties in removing, by acid cleaning after the tests, 

the oxides formed on the specimen surfaces. These oxides led to weight gains, instead of 

weight losses. The accuracy of the weighing system {±30 μg [±0.001 oz]} may have introduced 

an error in the corrosion rate of ±30 nm/yr [±0.0012 mil/yr] for the foil specimens. 

The maximum corrosion rate measured in the liquid phase under all test conditions was less 

than 150 nm/yr [±0.0059 mil/yr] with an uncertainty of ±60 nm/yr [±0.0024 mil/yr]; the maximum 

corrosion rate measured in the vapor phase under all test conditions was less than 70 nm/yr 

[0.0027 mil/yr] with an uncertainty of ±40 nm/yr [±0.0016 mil/yr]. 

Surface examination with cross-sectional transmission electron microscopy revealed that the 

oxide thickness varied from 0 to 60 nm [0 to 0.0024 mil] for specimens exposed at 140 C 

[284 F] for up to 270 days, but it was as high as 500 to 600 nm [0.020 to 0.024 mil] for 

specimens exposed at 220 C [428 F] for 270 days, which corresponds to corrosion rates from 
3-2 

667 to 800 nm/yr [0.026 to 0.031 mil/yr]. These high rates are not consistent with the maximum 

corrosion rate obtained from the weight loss method {150 nm/yr [0.0059 mil/yr] at 220 C 

[428 F]}, and the discrepancy was probably caused by the incomplete removal of the oxide 

layer with the acid cleaning method used prior to the weight loss measurement. 

The experiments conducted under similar conditions (9.6 to 21 molal of salt concentration and 

0.5 to 7.4 nitrate-to-chloride ratio) at temperatures ranging from 160 to 220 C [320 to 428 F] 

yielded near zero corrosion rates {8±12 nm/yr for specimens exposed to the liquid phase and 

5±11 nm/yr [0.0002±0.00043 mil/yr] for specimens exposed to vapor phase} using the weight 

loss method (Dixit, et al., 2006). Auger depth profile analysis, however, revealed that the oxide 

thickness for selected specimens that were exposed to the liquid phase of the system at 220 C 

[428 F] varied from 100 to 370 nm [0.004 to 0.016 mil], which corresponds to corrosion rates 

ranging from 133 to 493 nm/yr [0.0052 to 0.019 mil/yr]. 

Considering the large uncertainties in the removal of the oxide layer on the specimens, the 

weight loss measurements may underestimate the corrosion rate. Because the thickest oxide 

layer observed was 600 nm [0.024 mil] in a 270-day test, which corresponds to a corrosion rate 

of 800 nm/yr [0.031 mil/yr], the maximum uncertainty is probably less than 1,000 nm/yr 

[0.04 mil/yr]. Therefore, the measured corrosion rates of Alloy 22 in the concentrated solution 

containing NaCl, NaNO3, and KNO3 as Orme, et al. (2004) and Dixit, et al. (2006) reported were 

less than 1 μm/yr [0.04 mil/yr] at temperatures from 120 to 220 C [248 to 428 F]. 

3.2 Corrosion Rate of Alloy 22 in Nitrate–Chloride Salt Systems 

Measured in Open Systems 

Because the potential repository drift would not be sealed (Sandia National Laboratory, 2007a), 

rock cracks and air spaces within rubbles would allow air and water vapor to move to or from 

the emplacement drift during the high temperature period. Air may be present in some sections 

of the drift where insufficient groundwater could be available in the boiling zone as the water 

supply for steam generation would continuously flow outward and keep the air from entering into 

the drift. Two sets of experiments were conducted at CNWRA in 2005 and 2006, respectively, 

under conditions that simulated the unsealed system. Experiments conducted in 2005 were 

summarized in a previous report (Yang, 2006) and described partially in a conference paper 

(Yang, et al., 2007). In 2006, another set of experiments was conducted to verify the 

experimental results measured in 2005. This section describes the experiments conducted in 

2006 and analyzes the results from the two sets of experiments. 

3.2.1 Test Specimens 

Alloy 22 specimens with different metallurgical conditions were used in the experiments 

conducted in both 2005 and 2006. The experiments include both creviced and uncreviced 

specimens. Alloy 22 was tested in mill-annealed, thermally aged, and as-welded conditions. 

The metallurgical conditions and geometry of the different specimens are provided in Table 3-1. 

The shape of the creviced specimens is similar to that which Dunn, et al. (2005) used and is 

shown in Figure 3-1. The chemical compositions of these specimens are given in Table 3-2. All 

specimens were polished with 600-grit silicon carbide paper before use. 

The weight loss measurements were conducted according to ASTM G1–03 

(ASTM International, 2005). The posttest specimens were first cleaned with soft brushes and 
3-3 

Table 3-1. Specimens Used in the Experiments 

Specimen

s Code

Heat 

Treatment

Heat Number 

of Base 

Metal

Heat 

Number of 

Filler Metal

Geometry (cm3

) 

[in3

]

Surface 

Area 

(cm2

) [in2]

Thermally 

aged* 

Thermally aged 

at 870 C [1,598 F] 

for 30 minutes 

2277-1-3150 Not applicable 1.27 H 0.787 H 0.152 

[0.5 H 0.31 H 0.06] 

or 

1.27 H 1.27H 0.152† 

[0.5 H 0.5 H 0.06] 

2.66 

[0.212] 

or 

3.99†

[0.618] 

Annealed* Mill-annealed 

specimens 

2277-3-3266 Not applicable 1.27 H 2.54 H 0.25 

[0.5 H 1.0 H 0.1] 

8.48 

[1.31] 

Welded* Welded specimens 2277-3-3292 WN 813 1.27 H 2.54 H 0.25 

[0.5 H 1.0 H 0.1] 

or 

1.27 H 2.54 H 0.15†

[0.5 H 1.0 H 0.06] 

8.48 

[1.31] 

or 

7.60†

[1.18] 

Annealed, 

creviced 

Mill-annealed 

specimen 

2277-3-3266 Not applicable Figure 3-1 12.9 

[2.00] 

Welded, 

creviced 

Welded specimen 2277-3-3292 WN 813 Figure 3-1 12.9 

[2.00] 

*Specimens have a 2-mm [0.08-in]-diameter hole drilled at top for hanging in the test cells. 

†Used in the experiment conducted in 2005 only.

Figure 3-1. Schematic Diagrams for the Creviced Specimens Used in the Tests 

(PTFE—Polytetrafluoroethlene: 1 mm = 0.04 in) 

 

Alloy 276

Washer

Alloy 276

nut

Alloy 276

bolt

Serrated PTFE

crevice washer

Side view 

of specimen

Crevice assembly

Specimen

19 mm

~44.5 mm

~19.7 mm
3-4 

Table 3-2. Chemical Compositions of Specimen and Weld Fillers (in Weight Percent) 

Heat 

Number Ni* Cr* Co* Fe* Mo* Mn* W* V* Si* C* S* P*

2277-1-3150 Bal† 21.1 1.70 4.5 13.2 0.35 2.9 0.16 0.03 0.003 0.004 0.008 

2277-3-3266 Bal† 21.4 1.19 3.75 13.30 0.23 2.81 0.14 0.03 0.005 0.004 0.008 

2277-3-3292 Bal† 21.22 1.32 3.69 13.64 0.23 2.96 0.13 0.02 0.004 0.003 0.005 

WN813 Bal† 22.24 0.41 2.37 13.7 0.34 3.13 0.01 0.02 0.003 0.001 0.003 

*Ni–Nickel, Cr–Chromium, Co–Cobalt, Fe–Iron, Mo–Molybdenum, Mn–Manganese, W–Tungsten, V–Vanadium, 

Si–Silicon, C–Carbon, S–Sulfur, and P–Phosphorus 

†Bal–Balance 

rinsed with water and acetone, and then cleaned with a dilute HCl solution in an ultrasonically 

agitated bath. The specimens were then brushed with a soft brush and rinsed with acetone and 

water. The corrosion rates were calculated according to Eq. (3-1) 

CR W

ADt

  8.76 107 (3-1) 

where 

CR — corrosion rate (μm/yr) 

W — weight loss (g) 

A — surface area (cm2

) 

t — time (hr) 

D — density (g/cm3

) 

The microbalance used for the weight loss measurement had a precision of ±0.025 mg 

[3.52 H 10!5

 oz], which corresponds to an error of ±20 nm/yr [0.00079 mil/yr] in the corrosion 

rate for the specimens with the smallest surface area as shown in Table 3-1. 

3.2.2 Test Vessels for Weight Loss Corrosion Rate Measurements 

The corrosion experiments were conducted in the same type of vessels described in the 

previous report (Figure 3-2) (Yang, 2006). Approximately 1,000 g [2.2 lb] of a salt mixture 

containing equimolar quantities of NaCl, NaNO3, and KNO3 were placed in the glass vessel. 

The vessel was heated to a test temperature ranging from 130 to 220 C [266 to 428 F]. The 

system was kept at ambient pressure and vented to the open air through the vent bubbler, 

which was used to trap the vapor exiting the test vessel. The water in the bubbler was sampled 

periodically for pH and chemical analyses. The extra water in the bubbler that was used for 

chemical analysis or pH measurement was recycled back into the vessel in the experiment 

conducted in 2006, but not recycled in the experiment conducted in 2005. 

At the start of the experiment, deionized water was added to the test vessel so that 

approximately half of the salts dissolved. A small amount of water {10 to 20 mL [0.34 to 

0.68 oz]} was periodically added to the vessel to make up the water that evaporated and to 

maintain approximately the same liquid level in the test vessel. The water vapor pressure over 
3-5 

Figure 3-2. Schematic of Typical Glass Vessels Used for the Weight Loss Experiment 

{Total Volume of the Vessel Is Approximately 3 L [102 oz]} 

the liquid phase in the test vessels was close to the atmospheric pressure because of the 

continuous and slow evaporation of the water; the chemical composition in the liquid phase is 

determined by the phase equilibrium. 

Because the vessels were not sealed, the test systems are called “open systems.” They are 

designed to simulate the environmental conditions in the emplacement drift filled with water 

vapors produced by evaporation of groundwater and air diffusing or penetrating from the outside 

drift through rock cracks or rubble spaces (Sandia National Laboratories, 2007a,c). Note that 

such systems are not fully open, because widely open systems would not maintain the moisture 

required to simulate potential Yucca Mountain drift conditions. 

Test specimens were placed in the vessels in contact with the vapor and liquid phases 

(Figure 3-2). The vessels were heated with a heating mantle, and thermocouples were used to 

measure the temperatures of the liquid and vapor phases. 

At the end of the test and while the system was still at the test temperature, samples of known 

volume and weight were taken from the liquid phase for all the vessels for the experiment 

conducted in 2006, but for only one vessel for the experiment conducted in 2005. Because the 

samples crystallized as soon as the temperature decreased, the samples were redissolved and 

diluted with deionized water approximately 10 times the mass of the sample. The diluted 

solution, referred to as diluted samples in this report, was later analyzed for chemical 

composition and pH. 

3.2.3 Test Vessels for Characterization of Vapor Phase Acidity 

The vapor phase acidity was characterized by measuring the pH of vapor phase condensates 

obtained by using a distillation flask system shown in Figure 3-3. The two-stage liquid–solid 

Salt solution Liquid phase 

specimen

Vent 

bubbler

Liquid phase 

thermocouple

Vapor phase 

thermocouple

Vent

Vapor phase 

Specimen

Solid salt

Cooling 

water

Cooling 

water

Condenser

Salt solution Liquid phase 

specimen

Vent 

bubbler

Liquid phase 

thermocouple

Vapor phase 

thermocouple

Vent

Vapor phase 

Specimen

Solid salt

Cooling 

water

Cooling 

water

Condenser
3-6 

Figure 3-3. Schematic of Typical Glass Flask Used for the Characterization of Vapor 

Phase pH {Volume of the Flask Is Approximately 1 L [34 oz]} 

traps were used to prevent solid particles or liquid drops formed in the test flask from entering 

the condenser and possibly contaminating the vapor phase condensates. During the test, the 

valve located between the upper trap and the liquid collector was normally open. This valve 

was only closed when a sample from the collector was being taken. 

3.2.4 Weight Loss Results 

Table 3-3 shows the general corrosion rates obtained at 130, 160, 190, and 220 C [266, 320, 

374, and 428 F] from the weight loss measurements. Values are listed for the different types of 

specimens: (i) welded, (ii) annealed, (iii) thermally aged, (iv) annealed and creviced, and 

(v) welded and creviced. Note that two tests were conducted at 160 C [320 F] using Vessels 

B and E. The corrosion rates varied roughly from 0 to 500 nm/yr [0 to 0.020 mil/yr] for the 

specimens exposed to the liquid phase and from 0 to 80 nm/yr [0 to 0.0031 mil/yr] for the 

specimens exposed to the vapor phase. The small negative numbers listed in the table are due 

to experimental errors arising from weight loss measurements {±20 nm/yr [0.00079 mil/yr]} or 

from incomplete removal of the oxides by the acid cleaning method after exposure as discussed 

in Section 3.1. 

Table 3-3 also lists the corrosion rates averaged at each temperature for specimens exposed to 

either the liquid phase or the vapor phase in the same test vessel and at the same temperature 

and the average vent bubbler pH (see below for more information). For the specimens exposed 

to the liquid phase, the average corrosion rate ranged from 10.8 to 462 nm/yr [0.00043 to 

0.018 mil/yr]. For these samples, the average corrosion rate was low at 130 °C [266 °F], but 

significantly higher at 190 and 220 °C [374 and 428 °F]. Experiments conducted in Vessels B 

Vent 

bubbler

Vent

Sampler

Test flask

Liquidphase

Thermocouple

Solidliquid

trapper

Cooling 

water

Cooling 

water

Bypass

Liquid 

collector

Condenser

Valves

Vent 

bubbler

Vent

Sampler

Test flask

Liquidphase

Thermocouple

Solidliquid

trapper

Cooling 

water

Cooling 

water

Bypass

Liquid 

collector

Condenser

Valves
3-7 

Table 3-3. Corrosion Rates of Alloy 22 Specimens in the Liquid and Vapor Phases and t

Water in the Vent Bubbler Conducted in 2006 

Vessel A 

130 C [266 F] 

Vessel B 

160 C [320 F] 

Vessel E 

160 C [320 F] 

Vessel C 

190 C [374 F] 

Coupons 

Coupon 

ID 

CR-Liq* 

(nm/yr) 

CR-Vap† 

(nm/yr) 

Coupon 

ID 

CR-Liq* 

(nm/yr) 

CR-Vap† 

(nm/yr) 

Coupon 

ID 

CRLiq* 

(nm/yr) 

CRVap† 

(nm/yr) 

Coupon 

ID 

CRLiq* 

(nm/yr) 

CR-Va

(nm/y

Welded AW3 15.6 — BW3 22.0 — — — — CW3 130 — 

AW4 19.4 — BW4 25.2 — — — — CW4 172 — AW7 — 7.5 BW7 — 25.2 — — — CW7 — 36.7 AW8 — 24.2 BW8 — 30.4 — — — CW8 — 24.9 

Annealed A3 9.7 — B3 27.8 — — — — C3 128 — 

A4 !2.7 — B4 36.2 — — — — C4 153 — A7 — 11.3 B7 — 24.1 — — — C7 — 0.0 A8 — 3.2 B8 — 19.9 — — — C8 — !2.2 

Thermally aged 

G44 !15.4 — G48 23.4 — G40 — !4.8 G52 269 — 

G45 — — G49 40.1 — G41 — 1.6 G53 323 — G46 — — G50 — 40.1 G42 425 — G54 — 79.5 G47 — — G51 — 21.7 G43 499 — G55 — 65.5 

Annealed Creviced 

A9 — 30.9 B9 — 4.5 — — — C9 — !4.3 

A10 — 46.3 B10 24.5 — — — — C10 162 — 

Welded Creviced 

AW9 — 37.2 BW9 — 21.0 — — — CW9 — 33.3 

AW10 20.2 — BW10 22.4 — — — — CW10 269 — 

Average‡ — 10.8 23.0 — 27.7 23.4 — 462 0.8 — 201 30.0 Days — 184 184 — 189 189 — 197 197 — 135 — Points — 6 7 — 8 8 — 2 2 — 8 8 Average Bubbler pH 

7.2 7.8 2.9 6.6 

*CR-Liq―stands for corrosion rate in liquid phase †CR-Vap—stands for corrosion rate in vapor phase ‡Calculated after all negative values that are within experimental error {±20 nm/yr [0.00079 mil/yr]} were set to zero. Note: 1 mil/yr = 25,400 nm/yr 
3-8

and E at 160 °C [320 F] resulted in very different average corrosion rates {27.7 and 462 nm/yr 

[0.0011 and 0.018 mil/yr], respectively}. In light of the possible effect of the incomplete surface 

oxide removal as Dixit, et al. (2006) reported, the true corrosion rate may have been higher than 

the value measured from the weight loss measurement. To be consistent with the data reported 

previously (Yang, 2006) and with the long-term corrosion rate data reported for the Yucca 

Mountain Project (Sandia National Laboratories, 2007c), the effect of the incomplete oxide 

removal is ignored in this analysis. 

In the liquid phase at 190 and 220 °C [374 and 428 °F], the corrosion rates from the thermally 

aged specimens are approximately two times higher than those of the annealed specimens; 

the corrosion rates of the welded creviced specimens are also approximately two times higher 

than those of the annealed creviced specimens. However, significantly different corrosion 

rates were not observed among the different types of specimens tested at 130 and 160 °C 

[266 and 320 °F]. 

For the specimens exposed to the vapor phase, the average corrosion rate was roughly the 

same at different temperatures {0.8 to 30 nm/yr [0 to 0.0012 mil/yr]}. The small negative value 

at 160 °C [320 °F] for Vessel E (see Table 3-3) was within the error that the weight loss 

measurement may cause {±20 nm/yr [0.00079 mil/yr]} and was treated as zero for deriving the 

average value. Unlike the specimens exposed to the liquid phase, the specimens exposed to 

the vapor phase in this vessel exhibited the lowest value. Because the corrosion rates of 

specimens exposed to the vapor phase are insignificant {<0.1 μm/yr [0.004 mil/yr]} compared 

to those exposed to the liquid phase and corrosion rates near 10 to 100 nm/yr [0.0004 to 

0.004 mil/yr] are the approximate lower detection limits for the weight loss method under the 

experiment conditions, the following discussions focus on the corrosion of Alloy 22 exposed to 

the liquid phase. 

Figure 3-4 shows the posttest appearance of typical specimens tested at 160, 190, and 220 °C 

[320, 374, and 428 °F] for the experiments conducted in 2006. The original polishing marks on 

the three specimens [Figure 3-4(a), (c), and (d)] that showed corrosion rates lower than 

200 nm/yr [0.008 mil/yr] are clearly visible, whereas the polishing marks on the specimen 

that showed a higher corrosion rate [Figure 3-4(b)] have completely disappeared. In addition, 

the specimens in Figure 3-4(c) and (d) show visible differences between the welded fusion zone 

and the base metal. The polishing marks over the welded fusion zone are less visible than 

those over the base metal. Specimens in Figure 3-4(a) and (b) have very different appearances 

although they were exposed to the same temperature, confirming the high corrosion rate for 

specimens in Vessel E [Figure 3-4(b)] and the relatively low corrosion rate for specimens in 

Vessel B [Figure 3-4(a)]. 

Table 3-4 lists the corrosion rates obtained in the experiments conducted in 2005 (Yang, 2006; 

Yang, et al., 2007) and the average pH of the solution in the vent bubbler (Section 3.2.5 for 

more information). Optical and scanning electron microscopy results from the 2005 experiment 

showed significant corrosion for all specimens exposed to the liquid phase of the system 

(Yang, 2006; Yang, et al., 2007). Compared with the corrosion rate results shown in Table 3-4 

{ranging from 1 to 6 μm/yr [0.04 to 0.3 mil/yr]} for the specimens exposed to the liquid phase, 

the corrosion rates listed in Table 3-3 are significantly lower, except for some deviation at 

160 C [320 F]. No reliable temperature dependence can be drawn from the data shown in 

Table 3-4 due to the limited number of tests. 

 
3-9

Figure 3-4. Optical Appearance of the Specimens After the Tests in Liquid Phase For 

(a) Welded in Vessel B at 160 C [320 F], (b) Thermally aged in Vessel E at 160 C 

[320 F], (c) Welded in Vessel C at 190 C [374 F], and (d) Welded in Vessel D at 220 C 

[428 F] From the Experiment Conducted in 2006. 

[Polishing Marks in (b) Completely Disappeared.] 

 

(a) (b)

(c) (d)
3-10

Table 3-4. Corrosion Rates of Alloy 22 Specimens in the Liquid and Vapor Phases and the pH of the Water in 

the Vent Bubbler Conducted in 2005* 

 

Vessel A' 

150 C [302F] 

Vessel B'

160 C [320 F] 

Vessel C' 

180 C [354 F] 

Coupons 

Coupon 

ID 

CR-Liq 

(nm/yr) 

CR-Vap 

(nm/yr) 

Coupon 

ID 

CR-Liq 

(nm/yr) 

CR-Vap 

(nm/yr) 

Coupon 

ID 

CR-Liq 

(nm/yr) 

CR-Vap 

(nm/yr) 

Welded 1' — 290 8' — 0.6 3' — 221 

2' 1623 — 9' 3970 — 4' 1096 — 

Thermally 

Annealed 

1 — 197 9 — 344 5 — 474 

2 1869 — 10 4125 — 6 1702 — 

3 — 553 11 — 172 7 — 368 

4 2066 — 12 3094 — 8 1035 — 

Annealed 

Creviced 

 — — — C 8512 — A 1968 — 

 — — — CV — 1235 B — 954 

Welded 

Creviced 

plus 

Solution 

Annealed 

 — — — 12M 9402 — AV† 1190 — 

— — — 12T — 1374 9T† — 671 

Average — 1853 375 — 5821 781 — 1398 617 

Days — 80 80 — 49 49 — 60 60 

Average 

Bubbler pH 

3.0 2.0 3.5 

* Yang, L. “Corrosion of Alloy 22 in Concentrated Nitrate and Nitrate and Chloride Sale Environments at Elevated 

Temperature—Progress Report.” CNWRA 2006–02. San Antonio, Texas: Center for Nuclear Waste Regulatory 

Analyses. 2006. 

†Specimens were exposed for 32 days only. 

The experimental conditions for Tables 3-3 and 3-4 were nearly identical, except for 

the following: 

1. The exposure times for the data in Table 3-3 were longer (80 to 140 days) than the 

exposure times for the data in Table 3-4 (50 to 80 days). 

2. Some of the reagent grade chemicals were from different batches. 

3. Some water from the vent bubbler was recycled back into the test vessel as makeup 

water in the experiments listed in Table 3-3, while no vent bubbler water was recycled 

back as makeup water in the experiments listed in Table 3-4. 

3.2.5 Effect of pH on Corrosion Rate 

As discussed previously, the corrosion rates in Tables 3-3 and 3-4 exhibited no reasonable 

correlation with temperature, immersion time, or metallurgical conditions. For the 2006 data, the 

corrosion rates varied from 28 to 462 nm/yr [0.0011 to 0.018 mil/yr] for the experiments that 

were conducted at the same temperature {160 °C [320 °F]} and with the same batch of chemical 

reagents (Table 3-3). 

During the experiments, the water in the vent bubbler was frequently sampled and its pH was 

measured. Figure 3-5 shows the pH values for the water collected from the vent bubbler. The 
3-11

Figure 3-5. pH of the Vent-Bubbler Water and the Deionized Makeup Water That Was 

Added to the Test Vessels and Vent Bubblers During the Tests and pH 

of the Diluted Sample From Four Test Vessels at the End of the Tests for the Experiments 

Conducted in 2006 {T [F] = 1.8T [C] + 32}

pH values of the deionized water that was used to fill the bubbler and to add to the test vessels 

as makeup water are also shown in Figure 3-5. The slightly acidic pH (~5.8) was probably due 

to the absorption of atmospheric CO2. The pH of the water in the vent bubbler connected to 

Vessel E ranged from 2.2 to 3.2 and was significantly lower than that in the vent bubblers 

connected to the other test vessels. In the tests conducted at 190 °C [374 F] (Vessel C) and 

220 °C [428 F] (Vessel D), the pHs of the water from the vent bubblers were high (7 to 8.6) 

initially, but decreased slowly to slightly acidic values (4.5 to 6) after 100 days. The difference 

between the pH of the vent-bubbler water and the pH of the deionized water would likely be due 

to the gases exiting the test vessels. Therefore, Vessel E produced a significant amount of acid 

gases during the course of the test. The acid gases in Vessel E could have been caused by the 

impurities in the reagent chemicals used or the components in the PyrexTM glass vessels, such 

as silica or boron (see below for details), but the cause of the relatively high pH of the vapor 

phase in the other vessels is not known. 

Figure 3-5 also shows the pH for the samples taken from the liquid phase of four of the test 

vessels at the end of each experiment (one of the vessels, Vessel C, broke before the liquid 

sample was taken). The samples were diluted 10 times by weight with deionized water (diluted 

sample) prior to the pH measurement. In general, the lower the pH in the vent bubbler, the 

lower the pH of the diluted sample. This observation is expected according to Henry’s Law for 

the distribution of the acidic components in the liquid phase and in the gas phase. Note that the 

pH of the diluted samples from all the different vessels is in a narrow range (from 5.3 to 6.3) and 

none is above 7, which may be because of the pH of the added deionized water (~5.8). 

Figure 3-6 shows the pH of the vent bubblers and the pH of a diluted sample from Vessel A' for 

the 2005 tests conducted. 

2

3

4

5

6

7

8

9

40 70 100 130 160 190 220

Time (day)

pH

130 A―130°C

160 B―160°C

160 E―160°C

190 C―190°C

220 D―220°C

Added water

pH of final 

liquid diluted 

by 10X
3-12

Figure 3-6. pH of the Vent-Bubbler Water During the Tests Conducted in Vessels A', B', 

and C' and pH of the Diluted Sample From Vessel A' at the End of the Test for the 

Experiments Conducted in 2005 {T [F] = 1.8T [C] + 32} 

The average pH for the vent-bubbler water connected to each test vessel is also given in 

Table 3-3. Figure 3-7 shows the corrosion rates plotted against the pH of the vent-bubbler 

water for the 2006 and 2005 experiments. There is a strong correlation between the corrosion 

rate and the pH of the vent-bubbler water regardless of the test temperature. When the 

vent-bubbler water was acidic (pH close to or less than 4), the corrosion rate of the Alloy 22 was 

high. The test temperatures are also indicated in Figure 3-7. It appears that the corrosion rates 

are higher for previous tests than for the present tests for the same vent-bubbler water pH. This 

may be because the water that collected in the vent-bubbler was partially recycled during the 

2006 experiments, but the water that collected in the bubbler was not recycled during the 2005 

experiments. Bubbler water recycling during the 2006 experiments might have lowered the 

vent-bubbler pH for the liquid in the vessel with the same pH. In another words, the pH of the 

liquid in the vessels for the 2006 work should be higher than the pH of the liquid in the vessels 

for the 2005 data for the same vent bubbler pH. Because the corrosion rate in the liquid phase 

of the vessels is more directly related to the pH of the liquid rather than the pH of the vent 

bubbler, the corrosion rate was lower in the present work, even though one of the vent bubbler 

pHs was as low as that in the 2005 work. 

For most of the experiments conducted in 2006, samples were taken directly from the liquid 

phase at the end of the tests and the pH values for the liquid samples were measured. 

However, only one value for the pH of the liquid phase sample was available from the 

experiments conducted in 2005 (Yang, et al., 2007). 

In a short-term experiment to characterize the chemistry in the vapor phase of the test vessels, 

samples were taken from the liquid phase of the test flask after a relatively short period 

(7–10 days) and from the liquid collector for the condenser water. Figure 3-8 shows the pH of 

1

1.5

2

2.5

3

3.5

4

4.5

5

-10 10 30 50 70

Time (day)

pH

A'―150°C

B'―160°C

C'―180°C

pH of final 

liquid diluted 

by 10X
3-13

Figure 3-7. Average Corrosion Rates for Alloy 22 as a Function of pH of the Vent-Bubbler 

Water [Vent-Bubbler Water Was Recycled During the 2006 Work, but Not Recycled During 

the 2005 Work (Yang, 2006)] {T [°F] = 1.8T [°C] + 32} 

the diluted sample taken directly from the liquid phase of the test flasks as a function of the pH 

of the condenser water for the short-term tests and the one data point from the 2005 

experiments (Yang, et al., 2007). For the short-term tests conducted in 2006, the pH of the 

condenser water and the pH of the diluted sample are linearly related, which is consistent with 

Henry’s Law. Figure 3-8 also shows that the pH of the diluted samples from the short-term tests 

is significantly lower than that of the 2005 experiment for the same bubbler water pH. This is 

probably because the liquid samples for the 2005 experiments were taken 80 days after the 

solution was reacted in the vessel and the samples for the vent bubbler pH were taken during 

the course of the 80-day test (the average pH value). Conversely, the samples from the 

short-term tests were taken only about 10 days after they were reacted in the vessels. The 

acidic components of the liquid in the 2005 long-term tests might have decreased with time by 

processes such as degassing. According to Henry’s Law, it is reasonable to assume that the 

slope of the linear correlation between the pH of the diluted samples at the end of the 2005 

experiments and the pH of the vent-bubbler water averaged over the duration of the previous 

tests is the same as the slope of the linear correlation between the pH of the diluted samples 

and the pH of the condenser water from the short-term tests. Based on this assumption, the pH 

of the diluted sample for two of the 2005 tests was estimated using the single available data 

point (bubbler water pH = 3.0 and diluted sample pH = 4.4). Figure 3-8 shows that the 

estimated pH values for the diluted samples are 4.7 and 3.9 for the bubbler water pH of 3.5 and 

2.0, respectively. 

With these two estimated data points for the 2005 tests and the four data points in Figure 3-6 for 

the 2006 tests, the corrosion rates of Alloy 22 were plotted against the pH of the diluted samples 

for both the 2006 and the 2005 tests, as shown in Figure 3-9. The two sets of corrosion rate 

data exhibit an exponential relationship with pH. When the pH of the diluted samples is higher 

0

1

2

3

4

5

6

7

0 2 4 6 8 10

pH of vent bubbler

Corrosion Rate (μm/y)

2005 Data

2006 Data 160°C

160°C

220°C

150°C

180°C

190°C

130°C

160°C
3-14

Figure 3-8. Relationship Between the pH of the Diluted Sample From the Liquid Phase 

and the pH of the Bubbler Water. Two Values for the pH of the Diluted Sample for the 

2005 Data (3.9 and 4.7) Were Estimated Using the Linear Relationship 

for the Short-Term Data From the 2006 Work. 

Figure 3-9. Relationship Between Corrosion Rates of the Alloy 22 Specimens Exposed to 

the Liquid Phase and the pH of the Diluted Samples From the Liquid Phase (pH Values 

for Two of the 2005 Data Were Estimated As Shown in Figure 3-8) 

4.4

3.9

4.7

2

2.5

3

3.5

4

4.5

5

5.5

6

13579

pH of Bubbler Water

pH of Diluted Sample

2006 Short-term data

2005 data

Estimated 2005 data

0

1

2

3

4

5

6

7

8

345678

pH of Diluted Sample

Corrosion Rate (μm/y)

2005 data

2006 Data

(

Exponential 

regression
3-15

than 6, the corrosion rates are low, but when the pH of the diluted samples is lower than 6, the 

corrosion rates are significantly higher. Therefore, the high corrosion rates for Alloy 22 in the 

liquid phase observed in the 2005 tests are likely to have been caused by the low pH in the 

liquid phase (diluted sample pH <4.8), where the low corrosion rates observed in the 2006 

experiments were due to the relatively high pH of the liquid phase (diluted sample pH 5.8).

3.3 Verification of the pH Effect on Alloy 22 Corrosion and 

Understanding of the pH Variations in the NaCl–NaNO3–KNO3

Systems 

Experiments were conducted to verify the effect of pH on Alloy 22 corrosion rates observed 

using the weight loss method in the NaCl–NaNO3–KNO3 system under elevated temperature 

conditions. Simulation calculations were also conducted to understand the variations in pH of 

the NaCl–NaNO3–KNO3 system at test temperatures. 

3.3.1 Corrosion Rates Measured With Coupled Multielectrode 

Array Sensors 

The experiments for verifying the corrosion rates observed with the weight loss method were 

conducted using a coated Alloy 22 coupled multielectrode array sensor. The sensor has nine 

electrodes made with mill-annealed Alloy 22 wires {1 mm [40 mil] in diameter and 150.2 mm 

[6 in] long}. The wires were coated with an inorganic coating to prevent the formation of a 

crevice between the wire and the surrounding insulating material (Chiang and Yang, 2007). A 

schematic diagram of the sensor with coated electrodes is shown in Figure 3-10(a), and a 

photograph of the sensor after the corrosion tests is shown in Figure 3-10(b). 

Figure 3-10. (a) Schematic Diagram of the Coated Multielectrode Array Sensor and 

(b) Photograph of a Coated Alloy 22 Multielectrode Array Sensor After the Corrosion Test 

Shown in Figure 3-11 

Coated

Sensing Sensing 

Electrode Electrode

Epoxy

Insulation Insulation

Multielectrode

Probe View 

from Bottom from Bottom 

Coated

Sensing Sensing 

Electrode Electrode

Epoxy

Insulation Insulation

Multielectrode

Probe View 

from Bottom from Bottom 

(a) (b)
3-16

During the test, the coupled multielectrode array sensor probe was immersed in the liquid phase 

of a vessel system similar to that shown in Figure 3-2. Reagent grade NaOH and HNO3 were 

used to adjust the pH of the liquid in the test vessel. 

Figure 3-11 shows the measured maximum corrosion rate [the rate derived from the current that 

was measured from the most anodic electrode (Yang, 2008; Yang, et al., 2002)] and the pH of 

the diluted samples. The temperature varied from 124 to 134 °C [255 to 273 F] and 

occasionally reached 145 °C [293 F] during the test. The general corrosion rate, as measured 

with the weight loss method and shown in Figures 3-7 and 3-9, is anticipated to be lower than 

the maximum corrosion rate from the multielectrode array sensor. This is because the 

maximum sensor current is derived from a small electrode {surface area = 0.00785 cm2 

[0.00122 in2]} experiencing the most corrosion (Yang, et al., 2002), whereas the corrosion rate 

from the weight loss tests is averaged over the entire surface area {>2.66 cm2 [0.412 in2

]}. The 

test vessel solution pH was initially adjusted with a dilute NaOH solution so that the diluted 

sample pH was about 8.5 and the measured maximum corrosion rate of Alloy 22 stabilized at 

values around 1 μm/yr [0.04 mil/yr] as shown in Figure 3-11. Then dilute nitric acid was added 

to the test vessel from time to time, and such additions are reflected by the instantaneous 

decrease in the diluted sample pH. The solution pH increased slowly after each sharp 

decrease, probably due to the evaporation of the acid in the system. After the diluted sample 

pH was lowered below 5, the maximum corrosion rate started to increase every time dilute nitric 

acid was added to the vessel. Figure 3-12 shows the relationship between the maximum 

corrosion rates and the pH of the diluted sample from the liquid phase. The maximum corrosion 

rate remained low {about 1 μm/yr [0.04 mil/yr]} over a wide pH range near neutral (from 4.8 to 

8.5) and increased exponentially with a decrease in pH when the diluted sample pH became 

lower than 4.8. 

The maximum corrosion rate, as shown in Figure 3-12, is consistent with the general corrosion 

rate shown in Figure 3-9. Both figures show low corrosion rates for Alloy 22 at pH values 

greater than 4.8 for the diluted sample and a sharp increase in the corrosion rate of Alloy 22 

when the diluted sample pH was lower than 4.5. 

Because the current from the most anodic electrode of the coupled multielectrode array 

sensor indicates nonuniform corrosion, the low rates in Figure 3-12 also show that the localized 

corrosion rate of Alloy 22 exposed to large quantities of saturated NaCl–KNO3–NaNO3 brine is 

low {#1 μm/yr [0.04 mil/yr]} at 130 C [266 F] when the pH of the diluted sample is higher 

than 4.8. 

3.3.2 Causes for the Variations in the pH of the NaCl–NaNO3–KNO3 Systems 

As discussed earlier, the pH of the vent-bubbler water varied significantly even for the 

vent-bubblers connected to vessels containing the same salts at the same temperature 

(Vessels B and E). The CO2 in the air and impurities in the chemical reagents used for the test 

and their interactions with the salts and the components of the glass vessel (silica and boron) 

were suspected to cause the pH variations. The experimental setup shown in Figure 3-3 was 

used to assess the effect of CO2. To assess the effect of the impurities and glass components, 

preliminary model simulations were conducted to understand the influence of these chemical 

species on the pH of the NaCl–NaNO3–KNO3 solutions. 
3-17

Figure 3-11. Maximum Corrosion Rates for Alloy 22 Measured With a Coupled 

Multielectrode Array Sensor at Different pH in the NaCl–NaNO3–KNO3 System at 130 C 

[266 F] (Maximum Corrosion Rate Was Derived From the Highest Anodic Currents 

Measured From All the Electrodes of the Sensor) 

Figure 3-12. Relationship Between the Maximum Corrosion Rates Measured With the 

Alloy 22 Coupled Multielectrode Array Sensor and the pH of the Diluted Samples From 

the Liquid Phase of the NaCl–NaNO3–KNO3 System 

0.1

1

10

100

08/01/07 08/06/07 08/11/07 08/16/07 08/21/07 08/26/07

Date

Maximum Rate (μm/yr)

1

2

3

4

5

6

7

8

9

10

pH of Diluted Sample 

Corr Rate

pH

0

5

10

15

20

25

30

35

40

3456789

pH of Diluted Sample

Maximum Rate (μm/yr)

Exponential

regression

(low pH only)
3-18

Effect of CO2 in Air 

Bechtel SAIC Company, LLC (2005) reported that salts containing sodium, chloride, and nitrate 

ions are not stable in CO2-containing environments at temperatures above 100 C [212 F]. In 

such environments, hydrochloric acid or nitric acid will be formed according to the 

following reactions: 

(s)CONaOH(g)CONO22Na (g)2HNO

(s)CONaOH(g)COCI22Na 2HCI(g)

3 2 2 32 3

2 2 32

 

 

 

 

Figure 3-13 shows the condenser water pH as a function of time using the experimental setup 

shown in Figure 3-3. The venting line was connected to the bubbler, as shown in Figure 3-3, 

during the first 23 days of the test and opened to the air for the remaining duration of the test. 

The pH of the condenser water was high initially (7.9), but decreased to 6.9 within the first 

15 days. The pH of the condenser water continued to decrease after the venting line was 

opened to air, but showed no acceleration in the pH decrease. It is not evident from Figure 3-13 

that the CO2 in the open air has caused the pH decrease; moreover, opening the venting line to 

air did not cause the bubbler pH to decrease to 2 or 3, as observed in Vessel E (Table 3-3). 

The final pH was close to that of the added deionized water. 

To provide a preliminary understanding of the pH changes in a solution containing NaCl, 

NaNO3, and KNO3 at elevated temperatures, thermodynamic modeling of the chemistry of the 

salt system was performed, using software developed by OLI Systems, Inc. (see Section 2.3 for 

more information). 

Model Simulated Effect of Silica 

Silica is a primary component of the glass container (about 81 percent SiO2) used during the 

test. Figure 3-14 shows the result of thermodynamic modeling that evaluated the effect of 

adding SiO2 on the pH of the NaCl–NaNO3–KNO3 solution. The calculated initial solution pH 

was 8.0. As SiO2 was added, the pH decreased and reached 6.6 when SiO2 reached its 

solubility limit (337 ppm), which is expected under the test conditions because of silica 

abundance in the glass and the relatively long test duration. The results indicate that silica did 

not cause the low pH observed in the liquid phase for some of the tests. 

Model Simulated Effect of Boron 

Boron is a major component of the glass container (about 13 percent B2O3) used in the tests. 

Figure 3-15 shows the result of thermodynamic modeling that evaluated the effect of adding 

boron [in the form of B(OH)3] on the pH of the NaCl–NaNO3–KNO3 solution. The solution pH 

decreased from 8.0 to 7.0 when the added B(OH)3 increased to 210 ppm. The solution pH 

continued to decrease as more boron was added and reached about 6.5 at a boron content of 

500 ppm. The results indicate that boron leached from the glass container could not cause the 

NaCl–NaNO3–KNO3 solution to attain an acidic pH. 

Model Simulated Effect of Phosphorous 

Phosphorous is one of the impurities listed in the specification sheet for the reagents used in the 

experiments (e.g., 4 ppm PO4

!

 in one of the NaCl batches). Figure 3-16 shows the calculated 

effect of phosphorous addition (added in the form of P2O5) on the pH of the NaCl–NaNO3–KNO3
3-19

Figure 3-13. Effect of Atmospheric Components on the pH of the Condenser Water 

Above the Test Flasks Containing Wet NaCl, NaNO3, and KNO3 Salt Mixtures at 

130 C [266 F]

Figure 3-14. Calculated Effect of Addition of SiO2 and Dissolved SiO2 Concentration 

on the pH of the NaCl–NaNO3–KNO3 Solution at 160 C [320 F] 

 

2

3

4

5

6

7

8

9

07/12/06 08/11/06 09/10/06 10/10/06 11/09/06 12/09/06 01/08/07

Date

pH of condensor water

Vessel outlet was 

open to ambient

Vessel was 

closed
3-20

Figure 3-15. Calculated Effect of the Addition of B(OH)3 and Dissolved B(OH)3

Concentration on the pH of NaCl–NaNO3 Solution at 160 C [320 F] 

Figure 3-16. Calculated Effect of the Addition of P2O5 and the Dissolved P2O5

Concentration on the pH of NaCl–NaNO3–KNO3 Solution at 160 C [320 F] 

 
3-21

solution. The solution pH decreased sharply from 8.0 to 5.0 with a P2O5 content as low as 

2.5 ppm. The results indicate that the presence of phosphorus impurity, most likely from the 

reagent grade chemicals used in the experiments, could have caused the low pH of some of the 

NaCl–NaNO3–KNO3 solutions. 

Other Impurities 

The reagents also contain other impurities that may affect the pH of the solution. For example, 

the KNO3 used in the experiment contains 20 ppm calcium and 10 ppm magnesium. If these 

impurities were in a soluble form such as chlorides or nitrates, they would undergo hydrolysis to 

form acid at the test temperatures (Pulvirenti, et at., 2004, 2003). In addition, the combination 

of the impurities mentioned previously (SiO2, B, and P) may increase the effect on the pH of 

the solution. 

3.4 Summary 

Orme, et al. (2004) and Gray, et al. (2006) reported the corrosion rate of Alloy 22 in 1 m NaCl 

and 4 m NaCl solutions containing different concentrations of hydrochloric, sulfuric, or nitric acid 

at 90 C [194 F] measured using electrochemical linear polarization resistance and 

electrochemical impedance spectroscopy. Figure 3-17(a) shows the dependence of the 

corrosion rate on the solution pH in 4 m NaCl solution. The solution pH values in Figure 3-17(a) 

were the theoretical values calculated with EQ3/6 thermodynamic software (Gray, et al., 2006). 

The corrosion rate was lower {less than 1 μm/yr [0.04 mil/yr]} at pH values higher than 2, but 

started to increase as the pH decreased below 2. The increase in corrosion rate was moderate 

for nitric acid—increasing to approximately 10 to 25 μm/yr [0.4 to 1 mil/yr] when pH ranged 

between 1.5 and 0.5 and to approximately 40 μm/yr [1.6 mil/yr] when pH was !0.2—but 

significant for hydrochloric and sulfuric acids {from less than 1 μm/yr [0.04 mil/yr] to 3,000 μm/yr 

[120 mil/yr] at pH 1.0}. The relatively low corrosion rate observed in the presence of nitric acid 

was attributed to the passive oxide film formed on the Alloy 22 specimen by the nitric acid. The 

passive behavior of the oxide film formed in the presence of nitric acid is also indicated by the 

high values of the open circuit potential [Figure 3-17(b)]. 

In the immersion tests and the test conducted with coupled multielectrode array 

probes, the critical pH at which the corrosion rate increased significantly in the NaCl–KNO3–

NaNO3 system was 4.5 for the solution that was diluted tenfold by weight and cooled to room 

temperature (Figures 3-9 and 3-11). The pH obtained from the diluted sample cannot be 

directly compared with the pH reported by Orme, et al. (2004), but the pH values for the 

undiluted solutions at the testing temperatures can be measured directly with a high 

temperature pH probe or calculated using thermodynamic software such as EQ3/6. 

Pulvirenti, et al. (2004, 2003) also conducted corrosion rate measurements for Alloy 22 in 

concentrated brines obtained by evaporating solutions containing Mg2+, Ca2+, Na+

, K+, Cl!, and 

NO3

!

. Because of the hydrolysis of the alkaline earth metals, Cl2, NO, NO2, HCl, and HNO3

were detected in the vapor phase of the system during evaporation. The pH of the vapor 

condensate and the residue liquid were close to 1 or less. The corrosion rates of Alloy 22 were 

on the order of 5 to 100 μm/yr [0.2 to 4 mil/yr] over the temperature range of 110 to 145 °C 

[230 to 293 °F]. Pitting corrosion was also observed in some of the specimens. The high 

corrosion rates reported by Pulvirenti, et al. (2004, 2003) are consistent with the corrosion rates 

measured in the low pH range, as shown in Figures 3-9 and 3-12. 
3-22

Figure 3-17. (a) Corrosion Rates and (b) Open Circuit Potential Versus pH for Alloy 22 in 

Hydrochloric (Circles), Sulfuric (Squares), and Nitric (Triangles) With 4 m NaCl at 

90 C [194 F] 
3-23

Manufacturers of Alloy 22 have reported the corrosion rate of Alloy 22 below or at boiling 

temperatures slightly over 100 °C [212 °F] in a wide range of chemical solutions 

(Haynes International, 2002). According to Haynes International, the corrosion rate of Alloy 22 

in dilute nitric acid (10 weight percent) at temperatures up to the boiling temperature is relatively 

low {25 μm/yr [1 mil/yr]}. However, the general corrosion rate of Alloy 22 in hydrochloric 

solutions varies from 75 to 10,000 μm/yr [3 to 400 mil/yr] when the concentration changes from 

1 to 10 wt percent at boiling temperatures. In a solution containing 4.9 wt percent HNO3 and 

0.125 wt percent HCl, the corrosion rate of Alloy 22 was 50 μm/yr [2 mil/yr] at 

boiling temperature. 

The experimental results reported in this chapter conclude that the NaCl–NaNO3–KNO3 system 

is not corrosive toward Alloy 22, exhibiting a general corrosion rate less than 1 μm/yr [.0.4 mil], 

unless the pH is lower than 4.8 as measured at room temperature for the brine solution that was 

diluted 10 times by weight. Only at a pH lower than 4.8 can high corrosion rates be expected. 

3.5 References 

ASTM International. “Standard Practice for Preparing, Cleaning, and Evaluating Corrosion Test 

Specimens.” G1–03. West Conshohocken, Pennsylvania: ASTM International. 2005. 

Bechtel SAIC Company, LLC. “Analysis of Dust Deliquescence for FEP Screening.” Report 

ANL–EBS–MD–00007. Rev. 01. Las Vegas, Nevada: Bechtel SAIC Company, LLC. 2005. 

Chiang, K.T. and L. Yang. “Development of Crevice-Free Multielectrode Sensors for Elevated 

Temperature.” Paper No. 07376. CORROSION/2007. Houston, Texas: NACE International. 

2007. 

Dixit, S., S. Roberts, K. Evans, T. Wolery, and S. Carroll. “General Corrosion and Passive Film 

Stability.” UCRL–TR–217393. Livermore, California: Lawrence Livermore National Laboratory. 

2006. 

Dunn, D.S., O. Pensado, Y.-M. Pan, R.T. Pabalan, L. Yang, X. He, and K.T. Chiang. “Passive 

and Localized Corrosion of Alloy 22—Modeling and Experiments.” CNWRA 2005–02. San 

Antonio, Texas: CNWRA. 2005. 

Gray, J.J., J.R. Hayes, G.E. Gdowski, B.E. Viani, and C.A. Orme. “Influence of Solution pH, 

Anion Concentration, and Temperature on the Corrosion Properties of Alloy 22.” Journal of the 

Electrochemical Society. Vol. 153. No. 3. pp. B61–B67. 2006. 

Haynes International. “HASTELLOY C–22 Alloy.” Bulletin H–2019F. Kokomo, Indiana: 

Haynes International. 2002. 

Orme, C.A., J. Gray, J. Hayes, L. Wong, R. Rebak, S. Carroll, J. Harper, and G. Gdowski. 

“Fiscal Year 2004 Summary Report: General Corrosion and Passive Film Stability.” 

UCRL–TR–208588. Livermore, California: Lawrence Livermore National Laboratory. 2004. 

Pulvirenti, A., K. Needham, M. Adel-Hadadi, A. Barkatt, C. Marks, and J. Gorman. “Multi-Phase 

Corrosion of Engineered Barrier Materials.” Proceedings of the CORROSION 2004 

Conference. Paper No. 694. Houston, Texas: NACE International. 2004. 
3-24

Pulvirenti, A., K. Needham, M. Adel-Hadadi, E. Bishop, A. Barkatt, C.R. Marks, and 

J.A. Gorman. “Corrosion Behavior of Alloy 22 Under Conditions of Dynamic Groundwater 

Composition.” Proceedings of the CORROSION 2003 Conference. Paper No. 693. Houston, 

Texas: NACE International. 2003. 

Sandia National Laboratories. “Analysis of Dust Deliquescence for FEP Screening.” 

ANL–EBS–MD–000074. Rev. 01 AD 01. Las Vegas, Nevada: Sandia National Laboratories. 

2007a. 

–––––. “Multiscale Thermohydrologic Model.” ANL–EBS–MD–000049. Rev. 03 AD01. 

Las Vegas, Nevada: Sandia National Laboratories. 2007b. 

–––––. “General and Localized Corrosion of Waste Package Outer Barrier.” 

ANL–EBS–MD–000003. Rev. 03. Las Vegas, Nevada: Sandia National Laboratories. 2007c. 

Wang, P., R.D. Springer, A. Anderko, and R.D. Young. “Fluid Phase Equilibria.” Proceedings 

of 15th Symposium on Thermophysical Properties, Boulder, Colorado, June 22–27, 2003. 

Vol. 222/223. Boulder, Colorado: University of Colorado. p. 11. 2004. 

Yang, L. “Multielectrode Systems.” Techniques for Corrosion Monitoring. Chapter 8. 

L. Yang, ed. Cambridge, England: Woodhead Publishing. 2008. 

Yang, L. “Corrosion of Alloy 22 in Concentrated Nitrate and Nitrate and Chloride Salt 

Environments at Elevated Temperature—Progress Report.” CNWRA 2006–02. 

San Antonio, Texas: CNWRA. 2006. 

Yang, L., D. Dunn, G. Cragnolino, X. He, Y.-M. Pan, A. Csontos, and T. Ahn. “Corrosion 

Behavior of Alloy 22 in Concentrated Nitrate and Chloride Salt Environments at Elevated 

Temperatures.” CORROSION/2007. Paper No. 07580. Houston, Texas: NACE International. 

2007. 

Yang, L., N. Sridhar, O. Pensado, and D. Dunn. “An In-situ Galvanically Coupled 

Multi-Electrode Array Sensor for Localized Corrosion.” Corrosion. Vol. 58. p. 1,004. 2002. 
4-1

4 LOCALIZED CORROSION OF ALLOY 22 IN DUST 

DELIQUESCENCE BRINES 

This chapter focuses on understanding the localized corrosion of Alloy 22 in dust deliquescence 

brines. As reported in Chapter 2, dust deliquescence brines are expected to be rich in nitrate 

ions. In seepage water brines, nitrate acts as a localized corrosion inhibitor (Dunn, et al., 2005). 

This chapter summarizes the reported experimental data relating to localized corrosion of Alloy 

22 in dust deliquescence brines. The review is primarily based upon the reports by Felker, et al. 

(2006) and Dixit, et al. (2006). Additional references are also included where applicable. The 

experimental data show a strong tendency to localized corrosion inhibition of the alloy with 

increasing brine temperature. This result is based upon the values of corrosion potential and 

repassivation potentials that are reported in Sections 4.1 and 4.2. 

A mathematical model of the localized corrosion process simulating the cathodic side of the 

process is presented in Section 4.3. The current generated at the cathodic site is referred to as 

cathodic capacity. Model parameters used for estimating the cathodic capacity are discussed, 

and the cathodic capacity of Alloy 22 in various dust deliquescence brines is calculated for 

varying brine amounts at different temperatures. Finally, the work reported in this chapter and 

conclusions drawn from that work are provided in Section 4.4. 

4.1 Electrochemical Test Data on Localized Corrosion in Large 

Quantity of Brines 

Both corrosion and repassivation potentials are important electrochemical parameters with 

respect to localized corrosion. It has been long established that localized corrosion in the 

form of crevice corrosion could initiate if the corrosion potential is greater than the 

repassivation potential after long exposure of the metal in a test solution (Shukla, 2008). The 

studies on the localized corrosion susceptibility of Alloy 22 in aqueous solutions of chloride and 

nitrate salts of sodium and potassium in the temperature range of 110–150 °C [230–302 °F] 

by measuring the repassivation and corrosion potentials were reported by Felker, et al. (2006). 

As Felker, et al. (2006) reported, a three-electrode electrochemical cell consisting of a PyrexTM

corrosion cell, a working electrode, an auxiliary electrode, and a suitable reference electrode 

[saturated silver/silver chloride (SSC)*

 or saturated calomel electrode (SCE)†] through a Luggin 

capillary was used. Carbon dioxide free air was purged through a dispenser immersed in the 

solution. A condenser was used to condense the vapors formed due to evaporation of the test 

solution. Specimens were cleaned sequentially with distilled water, hexane, acetone, and 

methanol (or ethanol) prior to testing. Nonwelded Alloy 22 prism specimens were used in the 

experimental studies. Standard ASTM procedures (G 5 and G 59) were followed for conducting 

the potentiostatic, potentiodynamic, and corrosion potential measurements. 

The aqueous solutions were made using the chloride and nitrate salts of sodium and potassium. 

As stated in the report by Felker, et al. (2006), the chloride-dominated aqueous solutions 

existed only below 116 °C [240.8 °F]. In these solutions, the maximum nitrate salt solubility was 

determined. For solutions in which nitrate ions dominate the solutions, the maximum amount of 

soluble chloride was determined. It was observed that the nitrate-dominated solutions can 

 

*

Saturated silver/silver chloride reference electrode is referenced frequently throughout this chapter; therefore, the 

acronym SSC will be used. This acronym is predominantly used as a subscript. 

†

Saturated calomel electrode is referenced frequently throughout this chapter; therefore, the acronym SCE will 

be used. This acronym is predominantly used as a subscript. 
4-2

exist at temperatures much higher than 116 °C [240.8 °F] because of the very high boiling 

points of the mixed nitrate salt solutions. 

The corrosion potential measurements were conducted in the test solution containing 25m NO3



+ 1m Cl! at 110 °C [230 °F]. The pH of the solution was adjusted and varied between 2 and 7. 

These results are shown in Figure 4-1. As seen in the figure, the Ecorr values for the solution 

with pH equal to 2 are several hundred mV higher than the Ecorr values in solutions with higher 

pH. Ecorr is about !100 mVSSC in the neutral solution and increases by 100 mV for solutions with 

pH equal to 3. However, there is a sudden increase in Ecorr for the test solution with pH equal to 

2. At this pH, the nitrate becomes a dominant reductive species in solution. 

The values of the repassivation potential Erp reported by Felker, et al. (2006) are summarized in 

Table 4-1. The repassivation potential Erp versus NO3

!

/Cl! for different temperatures is 

graphically represented in Figure 4-2 where open and partially open symbols represent Erp of 

Alloy 22 in test solutions with neutral and acidic pH, respectively. As seen in the Figure 4-2, Erp

increases with increasing NO3

!

/Cl! at a fixed temperature. Also note that Erp values vary within 

a range of 0.383 to 0.522 VSSC at NO3

!

/Cl! ratio and temperature equal to 1 and 110 C [230 F], 

respectively, and Erp values for solution pH equal to 4 lie within 0.383 to 0.522 VSSC. Similar 

variations can be observed for Erp values at NO3

!

/Cl! equal to 10.5 where Erp values lie between 

0.54 and 0.683 VSSC. In this case, the Erp value for solution pH equal to 4.0 is at the lower end 

of the range. Therefore, the solution pH does have a statistically significant effect on Erp values. 

Furthermore, Erp reaches the asymptotic value 0.773 and 0.685 VSSC in the test solution 

containing only nitrate ions at 110 and 125 °C [230 and 257 °F], respectively. 

 

Figure 4-1. Corrosion Potential Versus Time of Alloy 22 in 25m NO3

!

 + 1m Cl! Solution 

With the pH Adjusted to Different Values at 110 °C [230 °F]. The Legend on the Graph 

Represents the Test Solution Identification Numbers (JE2915, JE2914, 

JE2906, and JE2902) and Corresponding Solution pH (Felker, et al., 2006).

Ecorr (mVSCC)
4-3

Table 4-1. Reported Value of Repassivation Potential Erp * 

Temperature 

(°C) [°F] pH 

Cl!

 

(molal) 

NO3

!

(molal) 

NO3

!

/Cl!

(molal ratio) Erp VSCC

110 [230] Neutral 8 0 0 !0.200 

110 [230] Neutral 8 0 0 !0.182 

110 [230] Neutral 8 0.8 0.1 !0.047 

110 [230] Neutral 8 0.8 0.1 !0.096 

110 [230] Neutral 8 1.6 0.2 !0.017 

110 [230] Neutral 8 2.4 0.3 0.320 

110 [230] Neutral 8 4 0.5 0.328 

110 [230] Neutral 6 6 1 0.383 

110 [230] Neutral 6 6 1 0.522 

110 [230] Neutral 2 42 21 0.674 

110 [230] Neutral 1 42 42 0.690 

110 [230] Neutral 0 42 NA 0.773 

110 [230] 4 8 0 0 !0.207 

110 [230] 4 8 0.8 0.1 !0.097 

110 [230] 4 8 4 0.5 0.443 

110 [230] 4 6 6 1 0.444 

110 [230] 4 6 6 1 0.480 

110 [230] 4 4 42 10.5 0.654 

110 [230] 4 2 42 21 0.661 

110 [230] 4 1 42 42 0.661 

125 [257] Neutral 4 42 10.5 0.622 

125 [257] Neutral 4 42 10.5 0.683 

125 [257] Neutral 2 42 21 0.595 

125 [257] Neutral 1 42 42 0.608 

125 [257] Neutral 0 42 NA 0.685 

125 [257] 4 4 42 10.5 0.540 

125 [257] 4 2 42 21 0.590 

125 [257] 4 1 42 42 0.685 

140 [284] 4 1 72 72 0.640 

140 [284] Neutral 1 72 72 0.725 

140 [284] Neutral 3 76 25.3 0.623 

140 [284] Neutral 3 76 25.3 0.656 

150 [302] Neutral 1 100 100 0.671 

*Felker, S., P.D. Hailey, T. Lian, K.J. Staggs, and G. Gdowski. “Alloy 22 Localized Corrosion Susceptibility in 

Aqueous Solutions of Chloride and Nitrate Salts of Sodium and Potassium at 110–150 °C.” UCRL–TR–218195. 

Livermore, California: Lawrence Livermore National Laboratory. 2006. 

Erp values lie above 0.383 VSSC irrespective of temperature and solution pH for NO3

!

/Cl! ratios 

greater or equal to 1. In test solutions with such high values of Erp, the localized corrosion could 

initiate only if solution pH is close to 2 (see Figure 4-1). This conclusion is based upon 

observations that the corrosion potential of Alloy 22 in test solutions containing 25m NO3

 and 

1m Cl! with pH equal to 2 could reach as high as 0.7 VSSC. Similarly, Erp values lie below 0 VSSC

if the NO3

!

/Cl! ratios are equal to or less than 0.2, and localized corrosion is possible 

(see Figure 4-1) if the solutions pH is close to 3. 

 
4-4

Figure 4-2. Graphical Representation of the Erp Data Tabulated in Table 4-1 

Rebak (2006) presented an additional experimental data point on repassivation potential in 

100 m NO3

!

 and 3 m Cl! test solution at 150 °C [302 °F]. This data point is not included in the 

Felker, et al. (2006) report. The Rebak (2006) reported value of repassivation potential of 

Alloy 22 in the test solution is close to 0.7 VSCC. The pH of the solution was not reported. 

Yang (2006) reported the short-term corrosion potential Ecorr for mill-annealed and welded 

plus solution-annealed Alloy 22 specimens in a brine saturated by a NaCl–KCl–NaNO3

salt mixture under semiopen conditions. The corrosion potential Ecorr values lie in the range of 

0.17 and 0.37 VSCE at 160 C [320 F]. Although there is a difference of 0.045V between SCE 

and SSC, the difference is insignificant as compared to measurement uncertainties in corrosion 

and repassivation potentials. 

As discussed in Chapter 2, the brines that are likely to form during the deliquescence period 

contain a high NO3

!

/Cl! molar ratio (higher than 5) at or above 120 C [248 F]. Therefore, dust 

deliquescence brine-induced localized corrosion is unlikely based on the measurements of 

corrosion and repassivation potentials, unless the solution has a sustainable low pH close to 2. 

At temperatures between 100 and 120 C [212 and 248 F], however, some brines that may 

develop during the deliquescence period may have nitrate-to-chloride ratios below 0.1, but the 

ratio of total inhibitors (nitrate + sulfate + bicarbonate + carbonate) to chloride is still high 

(higher than 3.3 for the KCl–KNO3 system). Based on the measurements of corrosion and 

repassivation potentials in pure nitrate and chloride solutions, localized corrosion is possible for 

nitrate-to-chloride molar ratios at or below 0.2. It is not clear, however, whether the 

repassivation potential will be higher than the corrosion potential for systems containing low 

nitrate but high concentrations of the other anion such as sulfate, bicarbonate, and carbonate, 

which were shown to be localized corrosion inhibitors at 95 C [203 F] (Dunn, et al., 2005). 
4-5

4.2 Immersion Test Data on Localized Corrosion in Large Quantity 

of Brines 

Dixit, et al. (2006) conducted localized corrosion tests using crevice assemblies specimens 

partially submerged in Na–K–Cl–NO3 brines. The experiments were conducted in a 2 L 

[0.44 gal] Alloy 276 autoclave under a slight pressure at 160 and 220 C [320 and 428 F]. A 

specimen holder made of Alloy 22 was used inside each autoclave. The Alloy 22 specimens 

were not electrically isolated from the Hastelloy C–276. Specimens were held in a vertical 

position to minimize settling of secondary precipitates on the Alloy 22 specimens. 

The experimental setup was assembled by first adding 1 L [0.22 gal] of brine solution in the 

autoclave at 90 C [194 F]. The assembled specimen holder was placed such that half of the 

specimen was fully immersed in liquid and the other half was fully exposed to the vapor phase 

above the liquid. The autoclave was then sealed, and the solution was purged with nitrogen for 

1 hour at 100 C [212 F] to remove oxygen and carbon dioxide gases. Dixit, et al. (2006) 

reported that at high temperatures and high nitrate concentrations, nitrate will be the most 

readily available oxidant for Alloy 22 corrosion. Test solutions were prepared by mixing the 

sodium nitrate and sodium chloride such that the NO3

!

/ Cl! molal ratio was 7.4 at 160 C 

[320 F], and NO3

!

/ Cl! ratio of 0.5 and 7.4 at 220 C [428 F]. Three autoclaves were used to 

test these three solutions, and 30 specimens were tested in each autoclave. The 30 specimens 

for each autoclave consisted of 8 pucks (of which 4 were immersed in the solution and 

4 were exposed to the vapor phase), 12 weight loss foils, and 10 creviced foils. The autoclaves 

were heated to get the desired temperature, and the experiment ran continuously for 

267 days (≈ 9 months). 

Dixit, et al. (2006) observed crevice corrosion on specimens immersed in the solution and on 

specimens held in the vapor phase at all conditions {i.e., in test solutions with NO3

!

/ Cl! molal 

ratio equal to 7.4 at 160 C [320 F]} and in test solutions with NO3

!

/ Cl! molal ratio equal to 

7.4 and 0.5 at 220 C [428 F]. All the creviced specimens in the three autoclaves showed 

crevice corrosion initiation. Crevice corrosion penetration depths were not quantified due to 

experimental difficulties. 

Lee and Solomon (2006) also reported localized corrosion for Alloy 22 disc specimens held at 

200 C [392 F] under dripping conditions in ambient pressure to simulate the potential 

Yucca Mountain condition. The dripping solution was the 10-time diluted standard simulated 

saturated water which contains NO3

!

, Cl!, Na+, and K+ (with a molar ratio of 21.1:3.6: 21.1:3.6). 

Unlike the tests reported by Dixit, et al. (2006), which were conducted in closed autoclaves, the 

Lee and Solomon (2006) dripping test was conducted in an open system where air was 

expected to be present. A liquid pool was observed on the top surface of the specimen during 

the dripping test, and the measured temperature of the pool was 143 C [289.4 F]. Because 

the relative humidity was low under the test conditions, the solution above the Alloy 22 

specimen was expected to be rich in NO3

!

. However, pits were observed on the Alloy 22 

surface that was in contact with the solution during the test. A scanning electron 

microscope image reveals that the size of the pits was as large as 10 μm [0.39 mil] after 3 hours 

of exposure. 

As is evident from these studies, localized corrosion did initiate in immersion tests in both closed 

autoclaves and open dripping environments for Alloy 22 specimens in the nitrate-rich solution. 

Crevice corrosion was also observed for specimens exposed to the vapor phase of the 

autoclave containing the nitrate-rich brine solutions. These observations appear to contradict 
4-6

the electrochemical test results reported in the previous section, which concluded that the 

localized corrosion initiation is inhibited in nitrate-rich brine solutions. 

4.3 Localized Corrosion Model for Alloy 22 in Dust Deliquescence 

Brines Under Potential Repository Conditions 

In localized corrosion processes, metal corrodes at certain preferential sites at an accelerated 

rate. The localized corrosion of Alloy 22 in the dust deliquescence brines can occur in the form 

of crevice corrosion when the corrosion potential is greater than the repassivation potential. 

This condition can be represented by the following expression 

corr EE rp  0 (4-1) 

where 

Ecorr — corrosion potential [V] 

Erp — repassivation potential for localized corrosion initiation [V] 

It is further assumed that the crevice corrosion instantaneously initiates when the corrosion 

potential exceeds the repassivation potential. The choice of repassivation potential for crevice 

corrosion initiation is based upon several experimental observations reported in the literature for 

nickel-based alloys (Dunn, et al., 2000; Kolts and Sridhar, 1984). 

 

The localized corrosion sites will remain active as long as the corrosion potential exceeds the 

repassivation potential (Dunn, et al., 2005, 2003; Hua, et al., 2004) and there is sufficient 

cathodic current available to balance the anodic currents generated by metal dissolution 

reactions (Cui, et al., 2005; Kelly, et al., 2006). In the absence of either condition, localized 

corrosion could stifle and even repassivation may occur. For example, if the corrosion potential 

of Alloy 22 in a given dust deliquescence brine is greater than the repassivation potential, but 

the available cathodic capacity is of the order of the passive current density, then the localized 

corrosion sites are likely to corrode at the rate of general corrosion. Therefore, the sufficient 

cathodic current is necessary for achieving the high dissolution rates required to maintain the 

localized corrosion sites. 

In contrast to general corrosion, the cathodic and anodic reaction sites are physically separated 

in the crevice corrosion process. A schematic diagram of the crevice corrosion process 

depicting the physical separation of the anodic and cathodic regions is presented in Figure 4-3. 

Cathodic reactions dominate in the cathodic area, and the excess current generated in the 

cathodic area is referred to as cathodic capacity. The oxygen reduction reaction is the dominant 

cathodic reaction within the expected range of environmental conditions at the potential 

repository drift environment, and a sufficient cathodic capacity is needed for sustained 

propagation of crevice corrosion at a sufficiently high rate. 

The cathodic capacity is affected by the following factors: (i) electrochemical conditions, such 

as corrosion potential and repassivation potential; (ii) environmental conditions such as 

temperature, pH, and chemical composition of brines; (iii) quantity of available electrolytes; and 

(iv) oxygen reduction kinetics. 
4-7

Figure 4-3. Schematic Diagram of the Crevice Corrosion Process. The Anodic Reactions 

Predominantly Occur in the Crevice Region, Whereas Cathodic Reactions Dominate in 

the Cathodic Region Located Adjacent to the Crevice. 

This section focuses on estimating the role of brine quantity in the cathodic capacity. In the 

potential repository environment, DOE expects that only a limited quantity of dust (Bechtel SAIC 

Company, LLC, 2005) would accumulate at the waste package surface after the permanent 

closure. As a result, only a limited quantity of brine would be available for localized corrosion. 

Because the available dust is expected to be uniformly distributed on the waste package 

surface, DOE assumes that the dust deliquescence brine would be also uniformly distributed. 

DOE expects that the thickness and length of the brine film would play a major role in 

determining the cathodic capacity. At constant ionic conductivity and brine film length, 

the resistance to ionic flow and the corresponding ohmic potential drop increases as the 

brine film thickness decreases. For this reason, the effect of the characteristics of brine 

films on cathodic capacity should be evaluated. To this end, a one-dimensional 

physic-chemical process model was developed to estimate the cathodic capacity. The interplay 

of oxygen reduction kinetic phenomenon with brine quantity was integrated into the model to 

calculate the electrode potential and current distribution in the cathodic region as a function of 

film height, length, and electrochemical and environmental conditions for dust deliquescence 

brines. The current distribution in the cathodic region was then integrated to obtain the cathodic 

capacity for a given set of conditions. The information generated from cathodic capacity 

calculations is used to estimate the crevice penetration rate for a fixed area of the crevice site. 

4.3.1 Model Development 

A process-level model for calculating the cathodic capacity is presented in this section. As seen 

in Figure 4-3, the anodic and cathodic regions are physically separated. The crevice region is 

anticipated to become deficient in oxidants due to the physical barrier to diffusional flow created 

by the restricted geometry of the crevice former once the oxidant is consumed. As a result, the 

metal dissolution reaction dominates in the crevice region, whereas cathodic reactions dominate 

in the area adjacent to the crevice. As depicted in Figure 4-3, the anodic dissolution reactions 

 
4-8

predominantly occur in the crevice region formed underneath the crevice former. A generic 

form of the metal dissolution reaction can be represented by the following expression: 

  

2  y  yeyHOHMOyHM (4-2) 

where metal M reacts with the water molecules to produce metal hydroxide, hydrogen ions, and 

free electrons. 

In the potential Yucca Mountain repository environment, the dominant oxidant would likely be 

oxygen. In the region outside the crevice, the oxygen reduction reaction would take place under 

mild oxidizing conditions in neutral-to-alkaline aerated solutions. This reaction can be 

expressed as 

  

22 4eO2HO 4OH (4-3) 

In acidic solutions, the dissolved oxygen molecules are reduced according to the 

following expression 

2  2O2H4e4HO 

(4-4) 

It is assumed that even in acidic solutions, the contribution of the hydrogen evolution reaction is 

negligible compared to that of oxygen reduction. This assumption is justifiable because the 

equilibrium potential of the hydrogen evolution reaction is much lower than that of the oxygen 

reduction reaction (Bard and Faulkner, 1980). For the same reason, the contribution of the 

water reduction reaction to the cathodic current in neutral-to-alkaline solutions is expected to 

be negligible. 

Nitrate reduction may also be an important reduction reaction in acidic solutions, but not in 

neutral-to-alkaline solutions. This pH dependence of nitrate reduction can be inferred from the 

results of corrosion potential measurements on Alloy 22 similar to those presented in Figure 4-1. 

In addition, the values of general corrosion rates of Alloy 22 presented in Chapter 3 indicate that 

there is a marked increase in corrosion rate in low pH solutions. This increase in corrosion rate 

in acidic solution is attributed to reduction in nitrate. 

Dunn, et al. (2005) reported corrosion studies at lower temperatures, demonstrating that nitrate 

is an important inhibitor of localized corrosion of Alloy 22 and suggesting that nitrate reduction is 

not an important reduction reaction except at very acidic pHs (< 2.0). Because the probability of 

having a low pH (i.e., pH < 6) dust deliquescence brine is much lower than having a high pH 

solution in the potential Yucca Mountain repository environment (Yang, 2006), this study 

assumes that the dominant reduction reaction [as stated by Eq. (4-3)] involves only 

dissolved oxygen. 

Under a steady-state crevice corrosion process, the sum of net current generated by the anodic 

and cathodic reactions must be equal to zero to maintain charge neutrality. This statement can 

be represented by the following equation 

I I A C   0 (4-5) 
4-9

where 

IA — net anodic current generated by metal dissolution reactions [A]

IC — net cathodic current generated by the oxygen reduction reaction [A] 

The cathodic capacity is equal to the net cathodic current, IC. For a given set of electrochemical 

conditions that would initiate localized corrosion, the net cathodic current can be calculated in 

both acidic and neutral-to-alkaline solutions because sufficient information exists on the kinetics 

of oxygen reduction reactions in both types of solutions. 

Excess cathodic current is generated in the cathodic region because of a potential gradient in 

the electrolyte solution. The excess cathodic current is equal to the current generated due to 

the oxygen reduction reaction minus the current generated by the anodic dissolution reaction in 

the cathodic region only. The electrode potential is expected to be close to the repassivation 

potential at the crevice and cathodic region interface, and the electrode potential farther into the 

cathodic region should approach the value of the corrosion potential. As a result, a gradient in 

the electrode potential is established in the cathodic region. The current density generated at a 

given location in the cathodic region is determined by the polarization curve of the oxygen 

reduction reaction. A schematic of the cathodic region with appropriate boundary conditions 

and a corresponding oxygen reduction polarization curve is presented in Figure 4-4. As 

indicated in the polarization curve, the electrode potential is equal to the repassivation potential 

at the interface of the two regions. The oxygen reduction current density near the repassivation 

potential is markedly higher than at the corrosion potential. As the electrode potential 

approaches the corrosion potential, the oxygen reduction reaction current density decreases. 

Thus, the electrode potential distribution in the cathodic region and the corresponding potential 

dependence of oxygen reduction kinetics are the sources of excess cathodic current. 

In the cathodic region, the current density due to the oxygen reduction reaction is determined by 

the interplay between the flow of current and the potential drop in the electrolyte, and the 

activation over potential as a function of the distance from the crevice mouth. The passage of 

current in the electrolyte leads to the potential drop governed by Ohm’s law, which can be 

expressed by the following equation 

i k d

dx    (4-6) 

where 

i — current density flowing through the electrolyte solution [A/cm2

] 

k — conductivity of the electrolyte solution [ohm!1

cm!1] 

 — electrostatic potential of the electrolyte at a given location [V]

x — position vector [m] 

The current flowing along the cathode surface varies with position due to the oxygen reduction 

reaction. This variation in current as a function of position can be expressed by the 

following equation 
4-10





 



   H

i)x(j

dx

di p (4-7) 

where 

j(x) — current density [A/cm2

] generated by the cathodic reduction reaction current at 

position x

H — height of the electrolyte film [m] 

ip — passive current density [A/cm2

] 

The combination of Eqs. (4-6) and (4-7) yields 

 

kH

i)x(j

dx

d  p  

2

2 (4-8) 

 

Figure 4-4. Schematic Representation of the Cathodic Region With Corresponding 

Cathodic and Anodic Polarization Curves. At the Interface of the Crevice and Cathode, 

the Potential Is Equal to the Repassivation Potential. The Excess Cathodic Capacity Is 

Generated in the Cathodic Polarization Curve When the Cathodic Region Electrode 

Potential Lies Between the Repassivation and Corrosion Potential.

O2+2H2O+4e- 4OHEcorr

M Mn+ + neO2+2H2O+4e- 4OH- O2+2H2O+4e- 4OH- O2+2H2O+4e- 4OH- O2+2H2O+4e- 4OHEcorr M Mn+ + neAlloy 22

Crevice Former

Crevice Region

(anodic reaction)

M

M(OH)+ H+

eO2 4OHCathodicregion

Brine film

Alloy 22

Crevice Former

Crevice Region

(anodic reaction)

M

M(OH)+ H+

eO2 4OHCathodicregion

Brine film

Cathodic reaction 

at the metal surface

Interface of anode 

and cathode

E=Erp

Cathodic reaction 

at the metal surface

Interface of anode 

and cathode

E=Erp
4-11

The current density generated due to the cathodic reduction reaction, j(x), can be expressed by 

the following equation 

 

   

    



 



   





 



  



corr

c

corr

c

V φ Ex

RT

α nF exp1

V φ Ex

RT

α nF

exp

xj

lim

corr

corr

i

i

i

(4-9) 

where 

Ecorr — corrosion potential [V] 

icorr — corrosion current density at the corrosion potential (equal to passive current 

density, ip) [A/cm2

]

αc — charge transfer coefficient for the cathodic reduction reaction

n — number of electrons per molecule of oxygen in the cathodic reduction 

reaction = 4

F — Faraday constant = 96,486 C/mol

R — universal gas constant = 8.3145 J/K/mol [1.987 cal/K/mol]

T — temperature of the system [K]

V — metal potential [V]

(x) — solution potential adjacent to the metal surface at position x

ilim — mass-transfer-limited current density [A/cm2

] 

The cathodic reaction can be under either activation or concentration control. If the cathodic 

reduction reaction is under concentration control, the current density is equal to the 

mass-transfer-limited current density; otherwise, the reduction reaction current density under 

pure activation control is equal to the numerator of Eq. (4-9). For a given thickness of 

electrolyte film, the mass-transfer-limited current density can be calculated by the 

following equation: 

H

b nFDC

limi  (4-10) 

where 

Cb — dissolved concentration of oxygen in the bulk electrolyte solution [mol/cm3

]

D — diffusivity of dissolved oxygen in the electrolyte solution [cm2

/sec] 

At the crevice and cathodic region interface, the electrode potential, which is equal to the 

potential drop across the metal–solution interface, is assumed to be the repassivation potential. 

This is a reasonable assumption given that very little or no cathodic current can be 

generated inside the crevice. A boundary condition for Eq. (4-8) can be represented by the 

following equation 

 E0)(xV rp (4-11) 

 
4-12

Under the free corrosion condition, the metal potential, denoted by V, is equal to zero. At the 

cathodic region and air interface, the following boundary condition for Eq. (4-8) prevails 

0

dx

d

κ

Lx

  



(4-12) 

Equation (4-12) represents the fact that no current can flow across the electrolyte and 

air interface. 

Equation (4-8) is solved subject to the boundary conditions given by Eqs. (4-11) and (4-12). 

The solution of Eq. (4-8) is obtained in terms of the solution potential, , as a function of position 

vector, x. Then the cathodic reduction current density, j(x), is calculated for each value of x

using Eq. (4-9). The net cathodic current is calculated by numerically integrating the cathodic 

reduction plus passive current density for a given length of the cathodic region according to the 

following equation 

I i j(x) dx C p  

x 0

x L

 





 (4-13) 

The net cathodic current calculated using Eq. (4-13) is in units of A/cm. Cathodic capacities are 

calculated assuming the brine film and the crevice region have a width of 1 cm [0.39 in]. This 

assumption is based upon a physical representation of the crevice region and brine film 

surrounding the crevice region where both regions extend infinitely along the width. This is a 

bounding assumption because a crevice region would be limited by the cylindrical shape of the 

waste package. As a result, only a fraction of the brine film surrounding the crevice region 

would generate excess cathodic current. The parameter values used in Eqs. (4-8), (4-9), and 

(4-10) are described in the following section. 

4.3.2 Model Parameters 

The cathodic capacity model needs the values of corrosion potential (Ecorr), passive current 

density (ipass), corrosion current density (icorr), electrolyte conductivity (κ), dissolved oxygen 

concentration (Cb), diffusivity of oxygen (D), film thickness (H), film length (L), and repassivation 

potential (Erp). It is assumed that the available brine is uniformly spread over the cathodic 

region located adjacent to the crevice (see Figure 4-3). The value of the cathodic reduction 

reaction charge transfer coefficient, denoted by αc in Eq. (4-9), is set equal to 0.49, and the 

value of oxygen diffusivity is set equal to 10!5

 cm2/sec [1.55 H 10!6 in2/sec]. These two values 

are adopted from reported experimental results on oxygen reduction kinetics on Alloy 22 

(Bryan, 2005). The values of film length, film thickness, and repassivation potential are varied, 

and their effects on cathodic capacity are studied in this report. 

The deliquescence brine properties needed to calculate the cathodic capacity of Alloy 22 are the 

brine film thickness and length, the effective ionic conductivity of the brine solution, and the 

dissolved oxygen concentration. The amount of dust on the metal surface is assumed to be 

26 mg/cm2

 [0.37 mlb/in2]. DOE considers this amount to be an upper bound value for dust that 

could deposit on the waste package surface in the potential Yucca Mountain repository (Bechtel 

SAIC Company, LLC, 2005). The dust particles are assumed to have a specific density of 1 

g/cm3

 [36.13 mlb/in3] and a spherical diameter equal to either 10 or 30 μm [0.39 or 1.18 mil], 

based on the density and upper-bound estimates of particle size reported by DOE (Bechtel 

SAIC Company, LLC, 2005). From the specific density of 1 g/cm3

 [36.13 mlb/in3], the number of 

particles per square centimeter of metal surface corresponding to 26 mg [5.73 H 10!2 mlb] of 
4-13

dust is calculated to be 4.97 H 107

 for 10-μm [0.39-mil]-diameter particles or 1.84 H 106 for 

30-μm [1.18-mil]-diameter particles. Based on geometric considerations and assuming cubic 

packing of particles, this number of particles in one square centimeter is calculated to occupy 

50-particle layers if the particle diameter is 10 μm [0.39 mil] or 17 particle layers if the particle 

diameter is 30 μm [1.18 mil]. Based on geometric considerations, the interparticle pore space is 

calculated to be 2.4 H 10!2

 cm3 [1.46 H 10!3 in3] for the 50 particle layers and 2.38 H 10!2 cm3

[1.45 H 10!3

 in3] for the 17 particle layers for 10-μm [0.39-mil] and 30-μm [1.18-mil] diameter 

particles, respectively, in 1 cm2

 [0.16 in2] of the waste package surface. These results are 

summarized in Table 4-2. 

To calculate the thickness of the brine film resulting from deliquescence, it is assumed that 

10 weight percent of the dust that deposits on the metal surface comprises deliquescent salt 

{2.6 mg/cm2

 [36.98 H 10!3 mlb/in2]}. The 10 weight percent value is based on the average 

weight percent reported in the literature (Reheis, 2003) for dust samples taken near Yucca 

Mountain. The deliquescent salt composition is assumed to be a NaCl–NaNO3–KNO3 mixture. 

At a temperature of 150 EC [302 EF], this salt mixture would deliquesce when the relative 

humidity in the environment reaches 17.7 percent and would form a saturated electrolyte 

solution with a density of 1.29 g/cm3

 [46.6 mlb/in3] according to the data shown in Table 4-3. 

The deliquescence relative humidity (or deliquescence point) and the brine density data in 

Table 4-3 are calculated using the chemical thermodynamic simulator (OLI Systems, 2005). 

The software has been validated for determining the brines properties formed due to 

deliquescence of salt mixtures with Na–K–Mg–Ca–Cl–NO3 ions (Gruszkiewicz, et al., 2007). At 

the deliquescence point of the NaCl–NaNO3–KNO3 salt mixture, the electrolyte solution is 

calculated to have a composition of 69.6, 4.07, and 61.6 molal of KNO3, NaCl, and NaNO3, 

respectively. The nitrate-to-chloride ratio for this brine is approximately equal to 32. This work 

assumes that localized corrosion initiates in this brine. 

The amount of brine that forms is limited by the amount of deliquescent salt. Given that only 

2.6 mg/cm2

 [36.98 H 10!3 mlb/in2] of salts are present on the metal surface, the calculated mass 

of water that forms upon deliquescence is 2.08 H 10!4

 g/cm2 [2.96 H 10!3 mlb/in2]. The brine 

Table 4-2. Dust Particle Size, Dust Layer Thickness, Brine Film Thickness, Ionic 

Conductivity, and Effective Ionic Conductivity of Dust Deliquescence Brine Formed by 

NaCl–NaNo3–NaNo3–KNO3 Salt Assemblage at 150 °C [302 °F] 

Particle 

Size 

(μm) 

[mil] 

Dust Layer 

Thickness 

(μm) [mil] 

Brine Film 

Thickness 

(μm) [mil] 

Ionic Conductivity of 

NaCl–KNO3–NaNO3

Brine (ohm!1

cm!1) 

[ohm!1

 in!1] 

Effective* Ionic 

Conductivity of 

NaCl–KNO3–NaNO3

Brine (ohm!1

cm!1) 

[ohm!1

 in!1] 

10 

[0.39] 

500 

[19.7] 

(50 particle 

layers) 

45.4 

[1.79] 

0.0382 

[0.0970] 

0.01834 

[0.04658] 

30 

[1.18] 

510 

[20.1] 

(17 particle 

layers) 

45.6 

[1.80] 

0.0382 

[0.0970] 

0.01826 

[0.04638] 

*Calculated using Bruggeman’s equation (Eq. 4-14). Temperature = 150 EC [302 EF]; Deliquescence Relative 

Humidity = 17.7 percent. Brine composition: 69.6 molal KNO3, 4.1 molal NaCl, and 61.6 molal NaNO3
4-14

Table 4-3. Calculated Deliquescence Relative Humidity, Dissolved Oxygen 

Concentration, and Ionic Conductivity of Deliquescence Salts at Various Temperatures 

Temperature 

(°C) [°F] Salts Present 

Deliquescence 

Relative 

Humidity 

(Percent)* 

Dissolved 

Oxygen 

Concentration 

(Parts Per 

Million) 

Ionic 

Conductivity

(ohm!1

cm!1) 

[ohm!1

in!1] 

120 [248] NaCl–NaNO3–KNO3 26.7 0.121 0.035 [0.089]

130 [266] NaCl–NaNO3–KNO3 22.2 0.092 0.033 [0.084]

140 [284] NaCl–NaNO3–

NaNO3@KNO3

18.2 0.068 0.032 [0.081]

150 [302] NaCl–NaNO3–

NaNO3@KNO3

17.7 0.065 0.038 [0.097]

160 [320] NaCl–NaNO3–

NaNO3@KNO3

16.5 0.057 0.042 [0.107]

170 [338] NaCl–NaNO3–

NaNO3@KNO3

14.6 0.044 0.042 [0.107]

*Some values reported here are slightly different from Table 2-13 due to slight differences in the input parameters for 

the different calculations. 

volume corresponding to this mass of water is 2.18 H 10!3

 cm3 [1.33 H 10!4 in3], based on a brine 

density of 1.29 g/cm3

 [46.6 mlb/in3]. 

The brine would occupy the interparticle pore spaces and would have a film thickness 

equivalent to 4.54 layers of 10-μm [0.39-mil]-diameter particles or 1.52 layers of 30-μm 

[1.18-mil]-diameter particles. 

A schematic representation of the spreading of the brine film layer with dust particles is 

represented in Figure 4-5, showing that the thickness of brine film is less than that of the 

dust layer. This work assumes that the capillary effect of the dust particles on brine film 

is negligible. 

The calculated ionic conductivity of the deliquescence brine at 150 EC [302 EF] using the 

software (OLI Systems, 2005) is 0.0382 ohm!1

cm!1 [0.0971 ohm!1in!1]. However, the effective 

ionic conductivity of the electrolyte is decreased by the presence of the dust particles. The 

effective ionic conductivity can be calculated using Bruggeman’s equation (Bruggeman, 1935) 

k k( ) eff

1.5   (4-14) 

where 

 



V

V V

solution

solution particle (4-15) 
4-15

Figure 4-5. Schematic Representation of Dust Deliquescence Brine Layer and Dust 

Layer on Alloy 22 Surface 

where 

Vsolution — volume occupied by the solution [cm3

] 

Vparticle — volume occupied by the particles [cm3

] 

K — ionic conductivity of the pure electrolyte solution [ohm!1

cm!1]

keff — effective ionic conductivity of the electrolyte solution in the presence of particles 

[ohm!1

cm!1] 

For the brine with a film thickness equivalent to 4.54 particle layers, the value of Θ is 

calculated to be 0.480 and the effective ionic conductivity is equal to 0.01834 ohm!1

cm!1

[0.04658 ohm!1

in!1]. For the brine with a film thickness of 1.52 particle layers, Θ is calculated to 

be 0.478 and the effective ionic conductivity is equal to 0.01826 ohm!1

cm!1 [0.04638 ohm!1in!1]. 

Because the relative difference between the two calculated values for both particle sizes is less 

than 1 percent, the effective ionic conductivity of 0.01834 ohm!1

cm!1 [0.04658 ohm!1in!1] is 

used in cathodic capacity simulations for dust deliquescence brines at 150 C [302 EF]. 

The ionic conductivity and effective ionic conductivity of dust deliquescence brines for both 

particle sizes are also summarized in Table 4-2. The effective ionic conductivity of dust 

deliquescence brines between 120 and 170 EC [248 and 338 F] is obtained by using the value 

of Θ equal to 0.480. The calculated values are listed in Table 4-4. 

Table 4-4. Effective Ionic Conductivity of Dust Deliquescence Brine Formed Between 120 

and 170 EC [248 and 338 EF] 

Temperature 

(EC) [EF] 

Effective Ionic Conductivity 

(ohm!1

cm!1) [ohm!1 in!1] 

120 [248] 0.0168 [0.0427] 

130 [266] 0.0158 [0.0402] 

140 [284] 0.0154 [0.0390] 

160 [320] 0.0202 [0.0512] 

170 [338] 0.0202 [0.0512] 
4-16

The dissolved oxygen concentration in the deliquescence brines at various temperatures is 

calculated using the model presented in Tromans (2000, 1998). This author gives equations for 

calculating the effect of electrolytes, including NaCl and NaNO3, on the dissolved oxygen 

concentration. However, Tromans (2000, 1998) did not provide any parameters for KNO3

solutions. Therefore, this study assumes KNO3 has the same effect on oxygen solubility as 

NaNO3, and the dissolved oxygen concentration in NaCl–NaNO3–KNO3 solutions is calculated 

using NaCl and NaNO3 parameters only. The dissolved oxygen concentration for various salt 

assemblages is summarized in Table 4-3. Note that the dissolved oxygen concentration does 

not change much for different salt assemblages between 140 and 160 EC [284 and 320 F]. For 

this reason, the dust deliquescence brine temperature is set to 150 EC [302 EF] for calculating 

cathodic capacity of Alloy 22 at the intermediate temperature. Cathodic capacity calculations 

are also carried in the temperature range of 120–170 EC [248–338 EF] to elucidate the 

effect of temperature. 

The corrosion potential of Alloy 22 in dust deliquescence brines is selected to be 0.6 VSHE, 

where SHE stands for the standard hydrogen reference electrode. This value is based on the 

experimental studies conducted by Yang (2006) in which the corrosion potential of Alloy 22 in 

dust deliquescence brine was found to be in the range of 0.4 to 0.65 VSHE. The corrosion 

potential of Alloy 22 in dust deliquescence brines is assumed to be independent of the 

temperature. 

At the corrosion potential, the corrosion and passive current density are calculated using the 

following equation 

i i i exp E

R

1

T

1

T corr pass a

0

ref

   





























 (4-16) 

where 

T — temperature of the system [K] 

ia

0

 — passive current density (A/cm2) at temperature Tref [K] = 10!8A/cm2 

 [6.45 × 10!8

A/in2] 

E — activation energy [J/mol] for the passive current density = 44,700 J/mol 

[10,676 cal/mol] 

R — universal gas constant = 8.3145 J/K /mol [1.987 cal/K/mol] 

Tref — reference temperature [K] for the passive current density = 368.15 K [203 EF] 

Equation (4-16) is used to calculate the passive current at 150 EC [302 EF] for dust 

deliquescence brines. Dunn, et al. (2005) originally proposed this equation, derived from their 

own experimental data, for calculating the passive current density and the corresponding 

general corrosion rate in seepage water brines at temperatures lower than 100 C [212 F]. 

Note that the corrosion rate estimated using Eq. (4-16) is also within the range of corrosion 

rates in dust deliquescence brines that Yang (2006) experimentally observed. 

4.3.3 Calculated Cathodic Capacities 

The distribution of the potential drop across the metal–solution interface (i.e., electrode 

potential) as a function of distance from the crevice region is presented in Figure 4-6(a). Four 
4-17

(a) (b) 

Figure 4-6. (a) Electrode Potential and (b) Current Density Distribution of Alloy 22 in the 

Cathodic Region. These Distributions Are Represented as a Function of Distance From 

the Crevice Region. The Electrolyte Brine Film Is Assumed To Be 0.1 m [3.94 in] Long 

and the Repassivation Potential Is Fixed at 0.1 VSHE. 

electrode potential distributions are presented corresponding to four film thicknesses: 1, 25, 45, 

and 100 μm [3.94 H 10!2

, 0.98, 1.8, and 3.94 mil]. These four simulations are performed using a 

repassivation potential equal to 0.1 VSHE at 150 EC [302 EF]. As seen in Figure 4-6(a), the 

electrode potential increases more sharply for a 1-μm [3.94 H 10!2

-mil]-thick film compared to 

the thicker films, and attains a value equal to the corrosion potential at a distance approximately 

0.06 m [2.36 in] from the crevice region. This result is consistent with the formulation of solution 

potential distribution represented by Eq. (4-8). An analysis of Eq. (4-8) indicates that solution 

potential is inversely proportional to the product of electrolyte ionic conductivity and film 

thickness. Therefore, the solution potential along the metal surface would decrease sharply 

for thin films. The corresponding current density distributions for four film thicknesses 

are presented in Figure 4-6(b). The area under the current density distribution curves is 

equal to the cathodic capacity of the film. As seen in Figure 4-6(b), the area under the 1-μm 

[3.94 H 10!2

-mil]-thick film is smaller compared to the 25-μm [0.98-mil]-thick film. The area 

under the current density distribution curve increases with increasing thickness, which indicates 

the cathodic capacity also increases with film thickness for a given film length. 

The cathodic capacities for various values of repassivation potential and film thickness are 

presented in Figure 4-7 where each data point represents cathodic capacity for one electrode 

potential and corresponding current density distribution. The corrosion potential and brine film 

length are set equal to 0.6 VSHE and 0.1 m [3.94 in], respectively, at 150 EC [302 EF]. The 

cathodic capacity is presented as a function of the difference between corrosion and 

repassivation potential (i.e., Ecorr–Erp). As seen in Figure 4-7, the cathodic capacity increases 

with increasing values of Ecorr–Erp. On other hand, the cathodic capacity increases with film 

thickness for a fixed value of Ecorr–Erp. However, for 1-μm [3.94 H 10!2

-mil]-thick film, the 

electrode potential distribution presented in Figure 4-6(a) indicates that cathodic capacity 

reaches an asymptotic value at a fixed value of Ecorr–Erp (i.e., additional increases in film length 

will not result in further increases in cathodic capacity). This occurs because the electrode 

potential attains an asymptotic value of corrosion potential close to the crevice region even for 
4-18

Figure 4-7. Cathodic Capacity of Alloy 22 in Dust Deliquescence Brine as a Function of 

the Difference Between the Corrosion Potential and the Repassivation Potential (Ecorr!Erp) 

for Various Film Thicknesses. The Corrosion Potential and Brine Film Length Are Fixed 

at 0.6 VSHE and 0.1 m [3.94 in], Respectively. A Penetration Rate of 100 μm/yr [3.94 mil/yr] 

or Higher Is Expected for 1 cm2

 [0.16 in2] of Active Crevice Region When the Cathodic 

Capacity Is Above the Horizontal Dashed Line.

large values of Ecorr–Erp. However, for thicker films and for a given value of Ecorr–Erp, a marginal 

increase in cathodic capacity may result because of an increase in film length. 

To explore the effect of brine film length on cathodic capacity, additional simulations were 

performed at a fixed film thickness but variable film length. The simulation results are presented 

in Figure 4-8, where cathodic capacity is plotted as a function of Ecorr–Erp. The brine film 

thickness and corrosion potential are set equal to 45 μm [1.8 mil] and 0.6 VSHE, respectively, at 

150 EC [302 EF]. As seen in Figure 4-8, there is a marginal variation in cathodic capacity with 

increasing film length. This result is consistent with the observation that the electrode potential 

changes sharply near the crevice region. As a result, most of the excess cathodic current 

density is generated in the cathodic region adjacent to the crevice. For this reason, a film length 

increase has only a marginal effect on cathodic capacity of Alloy 22 in dust deliquescence 

brines. In Figures 4-7 and 4-8, a horizontal dashed line is added to emphasize the 

consequence of cathodic capacity values above the dash line. Cathodic capacity greater than 

9.8 H 10!6 A/cm [2.49 H 10!5

 A/in] at a given Ecorr –Erp will result in a penetration rate of 

100 μm/year [3.94 mil/year] or higher for an active crevice area of 1 cm2

 [0.16 in2]. As seen in 

Figures 4-7 and 4-8, the cathodic capacity is below the horizontal dashed line even if Ecorr –Erp 

approaches a value of 0.5 V. 

Additional simulations were performed for a 45-μm [1.8-mil]-thick and 0.1-m [3.94-in]-long brine 

film to assess the effect of oxygen solubility and brine temperature. The corrosion potential is 

set equal to 0.6 VSHE, and the temperature of the brine is varied between 120 and 170 EC [248 

and 338 EF]. For a selected brine temperature, the dissolved oxygen concentration of the brine 

was specified using the values listed in Table 4-3. The corresponding effective ionic 

 
4-19

Figure 4-8. Cathodic Capacity of Alloy 22 in Dust Deliquescence Brines as a Function 

of the Difference Between the Corrosion Potential and the Repassivation Potential 

(Ecorr–Erp) for Various Brine Film Lengths. The Corrosion Potential and Brine Film 

Thickness Are Set Equal to 0.6 VSHE and 45 μm [1.8 mil], Respectively. A Penetration 

Rate of 100 μm/yr [3.94 mil/yr] or Higher Is Expected for 1 cm2

 [0.16 in2] of Active Crevice 

Region When the Cathodic Capacity Is Above the Horizontal Dashed Line. 

conductivity of brine is also varied according to the values listed in Table 4-4. The simulation 

results are presented in Figure 4-9, where cathodic capacity is plotted as a function of Ecorr–Erp. 

As seen in the figure, the cathodic capacity for the brine at 120 EC [248 EF] is above the 

horizontal dashed line when Ecorr –Erp is greater than 0.36 V. This result can be attributed to the 

higher oxygen solubility in the brine at 120 EC [248 EF]. For temperatures greater than or equal 

to 130 EC [266 EF], cathodic capacity varies significantly when Ecorr–Erp is less than 0.2 V. 

However, as Ecorr–Erp approaches the value of 0.5 V, the cathodic capacity values become 

closer at different temperatures with the exception of 120 C [248 F]. 

4.3.4 Discussion on Cathodic Capacity Results 

The cathodic capacity of Alloy 22 in dust deliquescence brines is determined primarily by 

the film thickness and by the difference between corrosion and repassivation potentials 

(i.e., Ecorr –Erp). As seen in Figure 4-6(a), the potential at the metal–solution interface 

approaches the corrosion potential with increasing distance from the crevice, and the gradient of 

the potential is fairly steep at the interface. As a result, most of the excess cathodic current is 

generated where the electrode potential is sufficiently cathodic with respect to the corrosion 

potential. Even though excess cathodic current also is generated farther into the cathodic 

region, the magnitude of the excess current is too small to influence significantly the cathodic 

capacity. Therefore, a sufficiently long brine film {0.1 m [3.94 in]} is needed adjacent to the 

crevice mouth for generating the excess cathodic current, and a film extending farther than 

0.1 m [3.94 in] into the cathodic region does not have much influence on the cathodic capacity. 

This observation is confirmed by the results presented in Figure 4-8, where cathodic capacity is 

plotted as a function of Ecorr –Erp for various film lengths at a film thickness equal to 45 μm 

[1.8 mil] at 150 EC [302 EF]. The calculated cathodic capacity changes only marginally with 

increasing film length for a given value of Ecorr –Erp. Therefore, the cathodic capacity of Alloy 22 
4-20

Figure 4-9. Cathodic Capacity of Alloy 22 in Dust Deliquescence Brines as a Function 

of the Difference Between the Corrosion Potential and the Repassivation Potential 

(Ecorr–Erp) for 45-μm [1.8-mil]-Thick and 0.1-m [3.94-in]-Long Brine Film at Different 

Temperatures. The Corrosion Potential Is Set Equal to 0.6 VSHE. A Penetration Rate of 

100 μm/yr [3.94 mil/yr] or Higher Is Expected for 1 cm2

 [0.16 in2] of Active Crevice 

Region When the Cathodic Capacity Is Above the Horizontal Dashed Line. 

in dust deliquescence brines is primarily dependent on Ecorr–Erp and film thickness as long as 

the film length is 0.1 m [3.94 in] or more. 

The effect of cathodic capacity on the crevice corrosion process can be understood by 

calculating the penetration rate for a given crevice site. A cathodic capacity of 9.8 H 10!6 A/cm

[2.49 H 10!5 A/in] can sustain metal dissolution at the rate of 100 μm/yr [3.94 mil/yr] for 1 cm2

[0.16 in2

] of active crevice area. Such cathodic capacity in dust deliquescence brines is attained 

at 120 EC [248 EF] when Ecorr –Erp is approximately equal to 0.36 V for a 45-μm [1.8-mil]-thick 

film (see Figure 4-9). For all other conditions, the cathodic capacity stays below 9.8 H 10!6 A/cm 

[2.49 H 10!5

 A/in]. 

An analysis similar to the one presented in the preceding paragraph can be carried out for 

evaluating the effect of cathodic capacity when the active crevice area is 0.1 cm2

 [0.16 in2] or 

less. In this case, a lower value of cathodic capacity may be sufficient to sustain a high 

penetration rate. For example, a cathodic capacity of 9.8 H 10!7 A/cm [2.49 H 10!6

 A/in] in dust 

deliquescence brines at Ecorr –Erp equal to 0.12 V for a film that is 45 μm [1.8 mil] thick at 150 EC 

[248 EF] film can sustain a penetration rate of 100 μm/yr [3.94 mil/yr] in a 0.1 cm2

 [0.016 in2] 

crevice area. However, a smaller crevice site will have less significance on the release of 

radionuclides compared to a larger site. 

In summary, the results of the cathodic capacity model for dust deliquescence brines 

indicate that localized corrosion with a penetration rate of 100 μm/year [3.94 mil/year] for 1 cm2

[0.16 in2

] of crevice can be achieved when Ecorr –Erp is greater than a critical value. 
4-21

4.4 Summary 

The experimental data on corrosion and repassivation potentials reported by Felker, et al.(2006) 

and Rebak (2006) indicate that the localized corrosion of Alloy 22 is not likely in large quantities 

of dust deliquescence-type brines that have nitrate-to-chloride ratios higher than 5, unless the 

brine pH is close to 2 at temperatures up to 150 °C [302 °F]. However, the localized corrosion is 

likely in brines having nitrate-to-chloride ratios at or below 0.2. 

According to the analysis in Chapter 2, the dust deliquescence brines that are likely to form at 

temperatures above 120 C [248 F] would have nitrate-to-chloride ratios higher than 5. Staff 

concludes that localized corrosion is not likely to occur under the potential repository conditions 

at temperatures between 120 and 150 C [248 and 302 F], based on the repassivation and 

corrosion potential data reported by Felker, et al.(2006) and Rebak (2006). However, for dust 

deliquescence brines that are likely to form between 100 and 120 C [212 and 248 F], the 

nitrate-to-chloride ratio may be low (less than 0.1) enough to initiate the localized corrosion. 

These dust deliquescence brines may contain other localized corrosion inhibiting species such 

as sulfate, bicarbonate, and carbonate. As a result, the total inhibitor (nitrate + sulfate + 

bicarbonate + carbonate) -to-chloride ratio may still be high enough to inhibit the localized 

corrosion of Alloy 22 in dust deliquescence brines formed between 100 and 120 C [212 and 

248 F]. Staff considers additional experimental data to be required in order to assess the 

susceptibility of Alloy 22 to localized corrosion in large quantities of dust deliquescence brine 

systems containing low nitrate but high concentrations of the other anions such as sulfate, 

bicarbonate, and carbonate. 

Immersion test data reported by Dixit, et al. (2006) and Lee and Solomon (2006) indicated that 

localized corrosion of Alloy 22 initiated in large quantities of nitrate-rich solutions at temperature 

equal to or greater than 140 C [284 F]. Localized crevice corrosion was also observed for 

specimens exposed to the vapor phase of the test vessel. These observations appear to 

contradict the repassivation and corrosion potential data reported by Felker, et al.(2006) and 

Rebak (2006), which indicated that localized corrosion is unlikely in nitrate-rich brine solutions. 

Additional long-term immersion tests data are required to resolve the discrepancy if the 

repassivation potential data is applied to evaluate the localized corrosion of Alloy 22 in large 

quantities of dust deliquescence brine solutions. 

The cathodic capacity results indicate that a sufficient positive value of Ecorr –Erp is needed to 

achieve sufficiently high crevice corrosion penetration rates. Based upon the experimentally 

measured values of Ecorr and Erp, it can be deduced that Ecorr –Erp should be at least equal to 

0.2 V and the maximum cathodic capacity on the order of 10!6 A/cm [2.54 H 10!6

 A/in]. A high 

crevice penetration rate over a large surface area of the crevice region cannot be achieved with 

a cathodic capacity of 10!6 A/cm [2.54 H 10!6

 A/in], even though a high penetration rate can be 

achieved over a smaller crevice region. Nevertheless, a smaller crevice site will have less 

significance on radionuclide release compared to a larger hole created by crevice corrosion. 

4.5 References 

Bard, A.J. and L.R. Faulkner. Electrochemical Methods, Fundamentals and Applications. 

New York City, New York: John Wiley & Sons. 1980. 

Bechtel SAIC Company, LLC. “Analysis of Dust Deliquescence for FEP Screening.”

ANL–EBS–MD–000074. Rev. 0. Las Vegas, Nevada: Bechtel SAIC Company, LLC. 2005. 
4-22

Bruggeman, D.A.G. “Berechnung Verschiedener Physikalischer Konstanten von Heterogenen 

Substanzen. I. Dielectrizitatskonstanten und Leitfahigkeiten der Mischkorper aus isotropen 

Substanzen.” Annalen der Physik. Vol. 24. pp. 636–679. 1935. 

Bryan, C.R. “Localized Corrosion Due to Dust Deliquescence on the Alloy 22 Waste Package 

Outer Barrier.” Nuclear Waste Technical Review Board Meeting, November 8–9, 2005, 

Las Vegas, Nevada. 2005. <http://www.nwtrb.gov/meetings/2005/nov/bryan.pdf> 

(7 April 2008). 

Cui, F., F. Presuel-Moreno, and R. Kelly. “Computational Modeling of Cathodic Limitation On 

Localized Corrosion of Wetted SS 316 at Room Temperature.” Corrosion Science. Vol. 47. 

pp. 2,987–3,005. 2005. 

Dixit, S., S. Roberts, K. Evans, T. Wolery, and S. Carroll. “General Corrosion and Passive Film 

Stability.” UCRL–TR–217393. Livermore, California: Lawrence Livermore National Laboratory. 

2006. 

Dunn, D.S., O. Pensado, Y.-M. Pan, R.T. Pabalan, L. Yang, X. He, and K.T. Chiang. “Passive 

and Localized Corrosion of Alloy 22 Modeling and Experiments.” CNWRA 2005-02. Rev. 1. 

San Antonio, Texas: CNWRA. 2005. 

Dunn, D.S., L. Yang, Y.-M. Pan, and G.A. Cragnolino. “Localized Corrosion Susceptibility of 

Alloy 22.” CORROSION/2003 Conference. Paper 03697. Houston, Texas: NACE. 2003. 

Dunn, D.S., G.A. Cragnolino, and N. Sridhar. “An Electrochemical Approach to Predicting 

Long-Term Localized Corrosion of Corrosion-Resistant High-Level Waste Container Materials.” 

Corrosion. Vol. 56. pp. 90–104. 2000. 

Felker, S., P.D. Hailey, T. Lian, K.J. Staggs, and G. Gdowski. “Alloy 22 Localized Corrosion 

Susceptibility in Aqueous Solutions of Chloride and Nitrate Salts of Sodium and Potassium at 

110–150°C.” UCRL–TR–218195. Livermore, California: Lawrence Livermore National 

Laboratory. 2006. 

Gruszkiewicz, M.S., D.A. Palmer, R.D. Springer, P. Wang, and A. Anderko. “Phase Behavior of 

Aqueous Na-K–Mg-Ca-Cl-NO3 Mixtures: Isopiestic Measurements and Thermodynamic 

Modeling.” Journal of Solution Chemistry. Vol. 36. pp. 723–765. 2007. 

Hua, F., J. Sarver, J. Jevec, and G. Gordon. “Corrosion Behavior of Alloy 22 and Ti Grade 7 in 

a Nuclear Waste Repository Environment.” Corrosion. Vol. 60, No 8. pp. 764–777. 2004. 

Kelly, R.G., A. Agarwal, F. Cui, X. Shan, U. Landau, and J. Payer. “Considerations of the Role 

of the Cathodic Regions in Localized Corrosion.” 11th International High-Level Radioactive 

Waste Management Conference, Las Vegas, Nevada, April 30–May 4, 2006. 

Kolts, J. and N. Sridhar. “Temperature Effects in Localized Corrosion in Corrosion of 

Nickel-Based Alloys.” Conference Proceedings of the International Conferences of Nickel-Base 

Alloys. Cincinnati, Ohio, October 23–25, 1984. pp. 191–198. 1984. 
4-23

Lee, S.G. and A.A. Solomon. “Localized Corrosion of Alloy C22 Nuclear Waste Canister 

Material Under Limiting Conditions.” Materials Science and Engineering. Vol. A 434. 

pp.114–123. 2006. 

OLI Systems, Inc. “A Guide to Using the OLI Software for Version 2.0 of the Analyzers.” 

Morris Plains, New Jersey: OLI Systems, Inc. 2005. 

Rebak, R. “Newer Alloy 22 Data and Their Relevance to High-Temperature Localized 

Corrosion.” 2006. Nuclear Waste Technical Review Board Meeting. September 25−26, 2006. 

<http://www.nwtrb.gov/meetings/2006/sept/rebak.pdf>. (7 April 2008). 

Reheis, M.C. “Dust Deposition in Nevada, California, and Utah, 1984–2002.” U.S. Geological 

Survey Open-File Report 03-138. 2003. 

Shukla, P. “Thermodynamics of Corrosion and Potentiometric Methods for Measuring Localized 

Corrosion.” Corrosion Monitoring Techniques. Chapter 7. L. Yang, ed. Success, 

United Kingdom: Woodhead Publishing. 2008. 

Tromans, D. “Modeling Oxygen Solubility in Water and Electrolyte Solutions.” Industrial & 

Engineering Chemistry Research. Vol. 39. pp. 805–812. 2000. 

–––––. “Temperature and Pressure Dependent Solubility of Oxygen in Water: A 

Thermodynamic Analysis.” Hydrometallurgy. Vol. 48. pp. 327–342. 1998. 

Yang, L. “Corrosion of Alloy 22 in Concentrated Nitrate and Chloride Salt Environments at 

Elevated Temperatures—Progressive Report.” CNWRA 2006-02. San Antonio, Texas: 

CNWRA. 2006.
5-1 

5 SUMMARY AND CONCLUSIONS 

5.1 Formation of Deliquescence Brines at Elevated Temperatures 

and Low Relative Humidities 

Information on the geochemistry of dust samples taken from Yucca Mountain and vicinity was 

reviewed. Dusts contain relatively small fractions of soluble salts, ranging from 10 percent at 

the maximum in remote locations to less than 0.1 percent in Yucca Mountain surface dusts. 

Leachates typically contain sodium, potassium, and calcium as major cations, and chloride, 

nitrate, and sulfate as major anions. Nitrate-to-chloride molar ratios exceeded 0.1 for all 

samples. Feldspars, quartz, and calcite were among the minerals typically found in 

Yucca Mountain dusts. 

The possible assemblage of deliquescent salts present in Yucca Mountain dusts was 

determined by simulating the evaporation of aqueous solutions derived from leaching dust 

samples from Yucca Mountain. The results of the thermodynamic simulations indicated that the 

most likely salt assemblages comprise NaCl–NaNO3–KNO3 and NaCl–KNO3 mixtures, with a 

relatively uncommon occurrence of assemblages containing calcium nitrate salts. 

Thermodynamic calculations also were conducted to determine the mutual deliquescence 

relative humidities of the salt assemblages and the chemistry of the brines that could form by 

deliquescence. The results indicate that the mutual deliquescence relative humidities of the salt 

assemblages are low, particularly at temperatures above 100 °C [212 °F]; if these salts are 

present in the potential repository, brines are likely to form by salt deliquescence even when the 

waste packages are relatively hot. However, the brines that are likely to form at temperatures 

above 120 °C [248 °F] would have molar ratios of nitrate to chloride higher than 5. At 

temperatures between 100 and 120 °C [212 and 248 °F], the nitrate-to-chloride molar ratio may 

be low (less than 0.1), but the ratio of the total amount of nitrate, sulfate, bicarbonate, and 

carbonate to chloride would still be higher than 3.3 for the KCl–KNO3 salt assemblage. 

 

Experiments were conducted to evaluate the effect of capillary retention by insoluble dusts on 

the corrosivity of deliquescence brines. The experiments were conducted at 70 °C [158 °F] in a 

temperature- and humidity-controlled chamber using solid mixtures with varying ratios of NaCl 

to quartz or rock dust particles placed on top of carbon steel coupled multielectrode array 

probes. The results indicate that even at a dust-to-NaCl weight ratio of 249, corrosion of carbon 

steel was observed, indicating that there was a significant amount of brine available on the 

metal surface. In addition, no significant effect of dust particle size on capillary retention was 

observed, except for the largest dust particles {>149 μm [>5.87 mil]} used in the test. 

Additional experiments were carried out to assess the corrosion of Alloy 22 in brines formed by 

deliquescence of salts (NaCl–KNO3 –NaNO3) mixed with rock dust. Tests were conducted 

using Alloy 22 coupled multielectrode array probes at 125 °C [257 °F] and relative humidities 

near the deliquescence point of an NaCl–KNO3–NaNO3 salt mixture. The results indicate that 

the corrosion rate of Alloy 22 was significantly lower than that of carbon steel, and no visible 

localized corrosion was observed. Because the coupled multielectrode probe measures the 

rate of nonuniform corrosion, including localized corrosion, the low currents measured with the 

Alloy 22 probe also indicate that the penetration rate of localized corrosion in the dust and 

NaCl–KNO3–NaNO3 mixture is low at 125 C [257 F] under the test conditions. 
5-2 

5.2 Corrosivity of Potential Deliquescence Brines to Alloy 22 Waste 

Package Materials 

The corrosivity of large quantities of potential deliquescence brines in contact with Alloy 22 was 

reviewed. Weight loss measurements for Alloy 22 in closed autoclaves and in open systems 

containing NaCl–KNO3–NaNO3 salt mixtures were reported for temperatures ranging from 

120 to 220 C [248 to 428 F]. The corrosion rate measured in closed autoclaves was lower 

than 1 μm/yr [0.04 mil/yr]. The measurements using Alloy 22 specimens with different 

metallurgical treatment (mill annealed, thermally aged, welded, and welded plus solution 

annealed) in a system open to the atmosphere that simulates plausible drift environments at the 

Center for Nuclear Waste Regulatory Analyses laboratory showed low corrosion rates except for 

cases where the pH of the solution was low. Only when the solution pH was lower than 4.5 as 

measured at room temperature after tenfold dilution by weight was the corrosion rate high and 

increased dramatically with decreasing pH. The dependence of the corrosion rate on different 

pH solutions was verified with an Alloy 22 coupled multielectrode array sensor. The corrosion 

rate of Alloy 22 was equal or lower than 1 μm/yr [0.04 mil/yr] in the NaCl–KNO3–NaNO3 brine at 

160 C [320 F] when the solution pH measured at room temperature after tenfold dilution by 

weight ranged from 4.8 to 8.5. The low pH observed in the test solutions was attributed to the 

impurities contained in the reagent grade salts used in the test. It is not clear whether such low 

pH brines could be formed under the potential Yucca Mountain drift conditions or whether such 

low pH brines, if formed, would persist over a long period of time and affect the integrity of the 

waste package containers. 

Corrosion rates of Alloy 22 were also reported for concentrated brines obtained by evaporation 

of solutions containing Mg2+, Ca2+, Na+

, K+, Cl!, and NO3

!

. Corrosion rates varied between 5 

and 100 μm/yr [0.2 and 4 mil/yr] within the temperature range of 110 to 145 °C [230 to 293 °F]. 

Localized corrosion in the form of pitting corrosion was also observed in the specimens. 

Because the vapor condensate and the residual liquid contacting the Alloy 22 specimen had a 

pH close to 1 or less. The hydrolysis of the alkaline earth metals (magnesium and calcium 

cations) present in the system caused the decrease in the solution pH. The observed high 

corrosion rates are consistent with the measured corrosion rates in the NaCl–KNO3–NaNO3

solutions at low pH values (< 4.5 as measured at room temperature after tenfold dilution by 

weight). However, such low pH systems may not last for long periods of time especially under 

potential repository conditions, because the hydrolysis process would eventually deplete the 

acid-producing alkaline earth metal ions. 

5.3 Localized Corrosion of Alloy 22 in Deliquescence Brines Under 

Potential Yucca Mountain Conditions 

At temperatures between 120 and 150 C [248 and 302 F], the literature data on repassivation 

and corrosion potential indicate that localized corrosion in the form of crevice corrosion in the 

presence of large quantities of dust deliquescence brines could initiate only when pH is close 

to or lower than 2. The reported values of the repassivation potential of Alloy 22 in dust 

deliquescence brines that are likely to form at these temperatures (nitrate-to-chloride molar 

ratio > 5) lie above 0.44 VSSC. For corrosion potential to exceed this value of the repassivation 

potential, the solution pH should be close to 2, which is far lower than the values normally 

expected under potential repository conditions. 

At temperatures between 100 and 120 C [212 and 248 F], however, the low nitrate-to-chloride 

molar ratio (<0.1) that may be present in the deliquescence brines from nitrate and chloride salt 
5-3 

mixtures would not be sufficient to prevent localized corrosion by raising the repassivation 

potentials. It is not clear whether high concentrations of other anions such as sulfate, 

bicarbonate, and carbonate also present in the nitrate-lean brines would be effective inhibitors 

to raise the repassivation potentials of Alloy 22 under the deliquescence brine conditions. 

Those anions were localized corrosion inhibitors for Alloy 22 at 95 C [203 F], but there are no 

data available for temperatures above 100 C [212 F]. 

The immersion test data discussed in the literature obtained from both closed autoclaves and 

open dripping environments indicated that localized corrosion of Alloy 22 specimens initiated at 

temperatures equal to or greater than 140 C [284 F] in nitrate-rich solutions where the nitrateto-chloride molar ratio was greater than 5. Localized crevice corrosion was also observed for 

specimens exposed to the vapor phase of the autoclave containing large quantities of nitraterich brine solutions. These observations appear to contradict the conclusions derived from 

repassivation and corrosion potential measurements, which indicated that localized corrosion is 

unlikely in dust deliquescence brine solutions with nitrate-to-chloride molar ratios greater than 5. 

It is not clear, however, whether the observed localized corrosion of Alloy 22 in the presence of 

large quantities of brines would propagate further with time. Additional data are needed to 

examine the discrepancy between the results of corrosion and repassivation electrochemical 

potential measurements and the immersion tests if the repassivation potential data are applied 

to evaluate the localized corrosion of Alloy 22 in large quantities of dust deliquescence brines. 

The measurements using the Alloy 22 coupled multielectrode probe indicated that the rate of 

nonuniform corrosion, including localized corrosion, was lower than 1 μm/yr [0.04 mil/yr] in large 

quantities of saturated NaCl–KNO3–NaNO3 brine at 130 C [266 F] for pH values from 4.8 to 

8.5 (as measured at room temperature for the liquid that was diluted by tenfold). 

The results of cathodic capacity estimations indicate that a sufficient positive value of Ecorr –Erp is 

needed to achieve high penetration rates. Based upon the experimentally measured values of 

Ecorr and Erp, it can be deduced that Ecorr –Erp can, at most, be equal to 0.2 V and the maximum 

cathodic capacity on the order of 10!6 A/cm [2.54 H 10!6

 A/in]. The high crevice penetration rate 

over a large surface area of the crevice region cannot be achieved with the cathodic capacity of 

10!6 A/cm [2.54 H 10!6

 A/in]. Even though a large penetration rate can be achieved over a 

smaller crevice region, a smaller crevice site would be of less risk significance to the release of 

radionuclides than a larger site. 

Considering that the amount of dust deliquescence brines on the waste package would be 

extremely small and the brine would be mixed with insoluble dust particles, the localized 

corrosion rate in the presence of dust deliquescence brines under plausible repository 

conditions is relatively low. This conclusion is supported by the low nonuniform corrosion 

current measured with the Alloy 22 coupled multielectrode probe in the mixture of dust and 

NaCl–KNO3–NaNO3 at 125 C [257 F].