Passive and Localized Corrosion of Alloy 22 Modeling and ... · Alloy 22 is controlled by outward...

CNWRA 2005-02Revision 1

PASSIVE AND LOCALIZED CORROSION OFALLOY 22—MODELING AND EXPERIMENTS

Prepared for

U.S. Nuclear Regulatory CommissionContract NRC–02–02–012

Prepared by

D.S. DunnO. PensadoY.-M. Pan

R.T. PabalanL. YangX. He

K.T. Chiang

Center for Nuclear Waste Regulatory AnalysesSan Antonio, Texas

December 2005

ii

PREVIOUS REPORTS IN SERIES

Number Name Date Issued

CNWRA 91-004 A Review of Localized Corrosion of High-Level April 1991Nuclear Waste Container Materials—I

CNWRA 91-008 Hydrogen Embrittlement of Candidate Container June 1991Materials

CNWRA 92-021 A Review of Stress Corrosion Cracking of High-Level August 1992Nuclear Waste Container Materials—I

CNWRA 93-003 Long-Term Stability of High-Level Nuclear Waste February 1993Container Materials: I—Thermal Stability of Alloy 825

CNWRA 93-004 Experimental Investigations of Localized Corrosion of February 1993High-Level Nuclear Waste Container Materials

CNWRA 93-006 Characteristics of Spent Nuclear Fuel and CladdingRelevant to High-Level Waste Source Term May 1993

CNWRA 93-014 A Review of the Potential for Microbially Influenced June 1993Corrosion of High-Level Nuclear Waste Containers

CNWRA 94-010 A Review of Degradation Modes of Alternate Container April 1994Designs and Materials

CNWRA 94-028 Environmental Effects on Stress Corrosion Cracking of October 1994Type 316L Stainless Steel and Alloy 825 As High-LevelNuclear Waste Container Materials

CNWRA 95-010 Experimental Investigations of Failure Processes of May 1995High-Level Radioactive Waste Container Materials

CNWRA 95-020 Expert-Panel Review of the Integrated Waste September 1995Package Experiments Research Project

CNWRA 96-004 Thermal Stability and Mechanical Properties of May 1996High-Level Radioactive Waste Container Materials:Assessment of Carbon and Low-Alloy Steels

CNWRA 97-010 An Analysis of Galvanic Coupling Effects on the August 1997Performance of High-Level Nuclear Waste ContainerMaterials

iii

PREVIOUS REPORTS IN SERIES (continued)


CNWRA 98-004 Effect of Galvanic Coupling Between Overpack Materials March 1998of High-Level Nuclear Waste Containers—Experimentaland Modeling Results

CNWRA 98-008 Effects of Environmental Factors on Container Life July 1998

CNWRA 99-003 Assessment of Performance Issues Related to September 1999Alternate Engineered Barrier System Materials andDesign Options

CNWRA 99-004 Effects of Environmental Factors on the Aqueous September 1999Corrosion of High-Level Radioactive WasteContainers—Experimental Results and Models

CNWRA 2000-06 Assessment of Methodologies to Confirm January 2001Revision 1 Container Performance Model Predictions

CNWRA 2001-003 Effect of Environment on the Corrosion of September 2001Waste Package and Drip Shield Materials

CNWRA 2002-01 Effect of In-Package Chemistry on the Degradation October 2001of Vitrified High-Level Radioactive Waste and SpentNuclear Fuel Cladding

CNWRA 2002-02 Evaluation of Analogs for the Performance Assessment March 2002of High-Level Waste Container Materials

CNWRA 2003-01 Passive Dissolution of Container Materials—Modeling October 2002and Experiments

CNWRA 2003-02 Stress Corrosion Cracking and Hydrogen October 2002Embrittlement of Container and Drip Shield Materials

CNWRA 2003-05 Assessment of Mechanisms for Early Waste March 2003Package Failures

CNWRA 2004-01 Effect of Fabrication Processes on Materials Stability— October 2003Characterization and Corrosion

CNWRA 2004-02 Natural Analogs of High-Level Waste Container January 2004Materials—Experimental Evaluation of Josephinite

iv

PREVIOUS REPORTS IN SERIES (continued)


CNWRA 2004-03 The Effects of Fabrication Processes on the July 2004Mechanical Properties of Waste Packages—Progress Report

CNWRA 2004-08 A Review Report on High Burnup Spent September 2004Nuclear Fuel—Disposal Issues

CNWRA 2005-01 Microbially Influenced Corrosion Studies of October 2004Engineered Barrier System Materials

CNWRA 2005-02 Passive and Localized Corrosion of Overpack November 2005Revision 1 Materials—Modeling and Experiments

CNWRA 2005-03 Microstructural Analyses and Mechanical Properties March 2005of Alloy 22

CNWRA 2006-001 Crevice Corrosion Penetration Rates of December 2005Alloy 22 in Chloride-Containing Waters—Progress Report

v

ABSTRACT

Long lifetime of the waste package is identified by the U.S. Department of Energy (DOE) as akey system attribute for the performance of the potential high-level waste repository at YuccaMountain, Nevada. Degradation processes including dry-air oxidation, uniform corrosion,disruption of the passive film, localized corrosion, and stress corrosion cracking are consideredimportant degradation processes that may influence the lifetimes of the waste packages. Thesedegradation processes are dependent on several factors including the metallurgical condition ofthe alloy and environmental conditions. In support of the U.S. Nuclear Regulatory Commission(NRC) prelicensing activities on issues important to the postclosure performance of the potentialrepository, the Center for Nuclear Waste Regulatory Analyses (CNWRA) is conducting anindependent technical assessment of effects of degradation processes on the performance ofthe engineered barrier materials. This report presents results of the CNWRA experiments onAlloy 22 of dry-air oxidation, uniform corrosion rate, passive film analyses, localized corrosionsusceptibility and propagation. The CNWRA investigations indicate that dry-air oxidation ofAlloy 22 is controlled by outward diffusion of chromium to form an external chromium oxide. Dry-air oxidation rates under conditions expected in the potential repository are low, andsignificant oxidation of the Alloy 22 waste package outer container is not expected. Uniformcorrosion rates under passive conditions are low and are not strongly dependent onenvironmental conditions or the metallurgical condition of the alloy. Localized corrosion ofAlloy 22 is limited to environmental conditions characterized by oxidizing or acidic, concentratedchloride containing solutions with low concentrations of inhibiting oxyanions. Propagation ratemeasurements suggest that stifling may limit crevice corrosion damage to the waste packageunder conditions where localized corrosion is possible. A performance assessment abstraction,conservatively ignoring this stifling mechanism, indicates that localized corrosion of wastepackages is possible, mainly along welded areas, if waste packages are contacted by seepingwater at high temperatures {80 °C [176 °F]}. This performance assessment abstraction is beingimplemented in a revised version of the Total-system Performance Assessment code.

vii

CONTENTS

Section Page

PREVIOUS REPORTS IN SERIES . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . iiABSTRACT . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . vFIGURES . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . ixTABLES . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . xiiiACKNOWLEDGMENTS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . xvEXECUTIVE SUMMARY . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . xvii

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

1.1 Objective . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1-11.2 Scope and Organization of the Report . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1-21.3 Relevant DOE and NRC Agreements . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1-2

2 PASSIVE AND LOCALIZED CORROSION OF OVERPACK MATERIALS—MODELING AND EXPERIMENTS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-1

2.1 Dry-Air Oxidation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-12.2 Experimental . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-22.3 Oxidation Rate Measurements . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-22.4 Oxidation Morphology and Assessment of Intergranular Oxidation . . . . . . . . 2-5

2.4.1 Scale Morphology and Composition at 850 °C [1,562 °F] . . . . . . . . . 2-52.4.2 Scale Morphology and Composition at 1,100 °C [2,012 °F] . . . . . . . . 2-5

2.5 Analysis of Results and Discussion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-9

3 OXIDE FILMS AND PASSIVE DISSOLUTION . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-1

3.1 Test and Analysis Methods . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-13.2 Passive Corrosion Rate Measurements . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-43.3 Surface Analyses of Alloy 22 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-11

3.3.1 Short-Term Tests in 0.028 M NaCl Solution . . . . . . . . . . . . . . . . . . . 3-113.3.2 Long-Term Tests in 4 M NaCl Simulated Groundwater . . . . . . . . . . 3-18

4 LOCALIZED CORROSION . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-1

4.1 Test Methods . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-14.2 Corrosion Potential Measurements . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-54.3 Repassivation Potential Measurements . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-64.4 Crevice Corrosion Initiation Tests . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-134.5 Localized Corrosion Propagation Rate Measurements . . . . . . . . . . . . . . . . 4-154.6 Coupled Multielectrode Array Sensors . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-19

viii

CONTENTS (continued)

Section Page

5 MODELS FOR WASTE PACKAGE DEGRADATION . . . . . . . . . . . . . . . . . . . . . . . . . 5-1

5.1 Corrosion Potential Model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5-15.2 Localized Corrosion Model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5-75.3 General Corrosion Model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5-105.4 Environmental Model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5-115.5 Conservative Assumptions in the Localized Corrosion Model . . . . . . . . . . . 5-20

6 SUMMARY AND CONCLUSIONS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6-1

6.1 Dry-Air Oxidation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6-16.2 Passive Dissolution and Oxide Film Analyses . . . . . . . . . . . . . . . . . . . . . . . . 6-16.3 Localized Corrosion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6-26.4 Passive Dissolution and Localized Corrosion Model . . . . . . . . . . . . . . . . . . . 6-2

7 REFERENCES . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7-1

ix

FIGURES

Figure Page

2-1 Oxidation Kinetics of Alloy 22 in Air at 850 °C [1,562 °F] and 1,100 °C [2,012 °F]:(a) Mass Change Versus Time Data; (b) Mass Change Because of Oxidation . . . . . 2-4

2-2 Cross-Section Through the Scale Produced on Alloy 22 After 120 HoursOxidation at 850 °C [1,562 °F]: (a) Scanning Electron MicroscopeBackscattered . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-6

2-3 Cross-Section Through the Scale Produced on Alloy 22 After 24 Hours Oxidationat 1,100 °C [2,012 °F]: (a) Scanning Electron Microscope Backscattered . . . . . . . . 2-7

2-4 Oxide Morphology for Alloy 22 Oxidized in Air for 120 Hours at1,100 °C [2,012 °F]: (a) Scanning Electron Microscope BackscatteredElectron Image . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-8

2-5 Distribution of Oxygen, Chromium, Iron, Molybdenum, and Tungsten in the AreaOutlined in Figure 2-4a: (a) Oxygen X-Ray Map; (b) Chromium X-Ray Map . . . . . . 2-10

3-1 Illustration of the Cylindrical and Crevice Corrosion Test Specimens . . . . . . . . . . . . 3-23-2 Short-Term Potentiostatic Passive Dissolution Rate Measurements for Welded

and Mill-Annealed Alloy 22 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-43-3 Measured Current Density for an Alloy 22 Crevice Specimen Under Potentiostatic

Conditions (!200 mVSCE) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-53-4 Measured Current Density for an Alloy 22 Crevice Specimen Under Potentiostatic

Conditions (!100 mVSCE) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-63-5 Passive Corrosion Rate for Mill-Annealed, As-Welded and Thermally Aged

Alloy 22 in 0.028 M NaCl As a Function of Temperature . . . . . . . . . . . . . . . . . . . . . . 3-73-6 Corrosion Rate of Alloy 22 in 0.028 M NaCl and 35-Percent MgCl2

(7.5 M Chloride) As a Function of Temperature . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-83-7 Activation Energy for Alloy 22 Corrosion Rates in 0.028 M NaCl and 35-Percent

MgCl2 (7.5 M Chloride) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-83-8 Passive Corrosion Rate for Mill-Annealed Alloy 22 in 0.028 M NaCl As a

Function of Temperature . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-93-9 Activation Energy for the Passive Corrosion Rate for Mill-Annealed Alloy 22 in

0.028 M NaCl As a Function of Temperature . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-103-10 Passive Corrosion Rate for Mill-Annealed Alloy 22 in 4 M NaCl As a Function of

Temperature . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-103-11 Concentration Depth Profiles of Alloy 22 Electrochemically Treated in Deaerated

0.028 M NaCl at 95 °C [203 °F] with an Applied Potential of (a) 100 mVSCE . . . . . . . 3-123-12 Scanning Electron Microscope Images of Alloy 22 Electrochemically

Treated in Deaerated 0.028 M NaCl at 95 °C [203 °F] with an Applied Potential . . . 3-143-13 High Resolution Spectrum Curve Fits of the Cr 2p3/2, Ni 2p3/2, and O 1s Regions

of Alloy 22 Passivated in Deaerated 0.028 M NaCl at 95 °C [203 °F] . . . . . . . . . . . 3-153-14 Species Concentration Depth Profiles of (a) Alloy 22 Electrochemically

Treated in Deaerated 0.028 M NaCl at 95 °C [203 °F] with an Applied Potential . . . 3-173-15 Concentration Depth Profiles of Alloy 22 Electrochemically Treated in

Deaerated 4 M NaCl Multi-Ionic Solution at 95 °C [203 °F] . . . . . . . . . . . . . . . . . . . 3-193-16 Species Concentration Depth Profiles of Alloy 22 Electrochemically Treated in

Deaerated 4 M NaCl Multi-Ionic Solution at 95 °C [203 °F] . . . . . . . . . . . . . . . . . . . 3-20

x

FIGURES (continued)

Figure Page

3-17 Concentration Depth Profiles in the Crevice Area of Alloy 22 ElectrochemicallyTreated in Deaerated 4 M NaCl Multi-Ionic Solution at 95 °C [203 °F] . . . . . . . . . . . 3-22

3-18 Species Concentration Depth Profiles in the Crevice Area of Alloy 22Electrochemically Treated in Deaerated 4 M NaCl Multi-Ionic Solution at95 °C [203 °F] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-23

4-1 Glass Cell Used for Alloy 22 Localized Corrosion PropagationRate Measurements . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-2

4-2 Schematic Diagram (a) and Different Configurations (b) of Multielectrode ArraySensor Probes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-4

4-3 Alloy 22 Corrosion Potential Measured at Temperatures from 25 to 95 °C[77 to 203 °F] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-5

4-4 Corrosion Potentials of Alloy 22 As a Function of Solution pH . . . . . . . . . . . . . . . . . . 4-64-5 Corrosion Potential of Alloy 22 As a Function of Temperature . . . . . . . . . . . . . . . . . . 4-74-6 Crevice Corrosion Repassivation Potentials for Mill-Annealed, Thermally Aged,

As-Welded and Welded-Plus-Solution Annealed Alloy 22 in Chloride Solutionsat 95 °C [203 °F] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-8

4-7 Crevice Corrosion Repassivation Potentials for Mill-Annealed and ThermallyAged Alloy 22 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-10

4-8 Crevice Corrosion Repassivation Potential Measurements Showing the Effect ofFluoride and Sulfate on the Localized Corrosion Susceptibility of Alloy 22 . . . . . . . 4-10

4-9 Crevice Corrosion Repassivation Potential Measurements Showing the Effect ofCarbonate and Bicarbonate on the Localized Corrosion Susceptibility . . . . . . . . . . 4-11

4-10 Results of Speciation Calculations of Solutions Containing SodiumBicarbonate and Sodium Chloride at 95 °C [203 °F] . . . . . . . . . . . . . . . . . . . . . . . . 4-12

4-11 Results of Speciation Calculations for Solutions Containing Sodium Carbonateand Sodium Chloride at 95 °C [203 °F] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-12

4-12 Speciation Calculations Showing the Maximum Anion Solubilities in SodiumChloride Solutions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-13

4-13 Initiation Time for Crevice Corrosion As a Function of Potential forAlloys 825, 625, and 22 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-14

4-14 Open-Circuit Corrosion Test for Alloy 22 in 5 M NaCl with the Addition of CuCl2at 100 °C [212 °F] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-15

4-15 Open-Circuit Corrosion Test for Alloy 22 in 5 M NaCl with the Addition of CuCl2at 100 °C [212 °F] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-16

4-16 Measured Current Density and Potential for Alloy 22 SpecimensGalvanically Coupled to Platinum at Various Durations in 4 M MgCl2 Solution . . . . 4-17

4-17 Photographs Showing Localized Corrosion Features of Alloy 22 Single-CreviceSpecimens After Corrosion in 4 M MgCl2 Solution at 110 °C [230 °F] . . . . . . . . . . . 4-18

4-18 Maximum Penetration Depths Measured on Alloy 22 Specimens CreviceCorroded in 4 M MgCl2 for Various Propagation Durations at 110 °C [230 °F] . . . . . 4-20

4-19 Typical Standard Deviation Signals from a 25-Electrode Coupled MultielectrodeArray Sensor Probe Made of Type 304 SS in Different Environments . . . . . . . . . . . 4-21

4-20 Standard Deviation Signals of the Alloy 22 and Three Other Probes in0.1 M Ferric Chloride Solution at Different Temperatures . . . . . . . . . . . . . . . . . . . . 4-22

xi

FIGURES (continued)

Figure Page

4-21 Relationship Between the Standard Deviation Signals from an Alloy 22 andThree Other Probes in 0.1 M Ferric Chloride Solution and Temperature . . . . . . . . . 4-23

4-22 Posttest Examination of the Surface of an Electrode in the Alloy 22 Probe . . . . . . . 4-244-23 Posttest Examination of the Surface of an Electrode in the Type 304 SS Probe . . . 4-254-24 Posttest Examination of the Surface of an Electrode in the Type 316 SS Probe . . . 4-254-25 Responses of the Standard Deviations of the Currents from Alloy 22 and Three

Other Probes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-26

5-1 Comparison of Computed Corrosion Potentials to Experimental DataAs a Function of pH and Temperature . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5-8

5-2 Ternary (Ca-SO4-HCO3) Phase Diagram Plotting Yucca Mountain UnsaturatedZone Porewater Compositions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5-13

5-3 Box Plots Showing the Range in pH and Concentrations of Chloride, Nitrate,Total Carbonate, and Sulfate of the Three Brine-Types . . . . . . . . . . . . . . . . . . . . . . 5-15

5-4 Estimated Concentrations at 110 °C [230 °F] (a) pH, (b) Chloride, (c) Nitrate,(d) Carbonate-Bicarbonate, and (e) Sulfate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5-16

5-5 (a) Distribution Functions for the Corrosion Potential and Critical Potential forLocalized Corrosion Resulting from the Adopted Stochastic Sampling Approach . . 5-18

5-6 Distribution Functions for the Corrosion Potential and Critical Potential forLocalized Corrosion Resulting from the Adopted Stochastic Sampling Approach . . 5-19

xiii

TABLES

Table Page

1-1 DOE and NRC Agreements Related to This Report . . . . . . . . . . . . . . . . . . . . . . . . . . 1-3

2-1 Composition of Alloys 22, 625, and 825 Heats and Alloy 622 Filler Metal . . . . . . . . . 2-3

3-1 Calculated Species Contributions for Various Core-Level Signals of Alloy 22Passivated at 100 mVSCE in Deaerated 0.028 M NaCl at 95 °C [203 °F] . . . . . . . . . 3-16

4-1 Parameters for the Localized Corrosion Regression Equations for Mill-Annealedand Thermally Aged Alloy 22 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-9

4-2 Maximum Penetration Depths of Alloy 22 Crevice Specimens Propagated atVarious Durations in 4 M MgCl2 Solution at 110 °C [230 °F] . . . . . . . . . . . . . . . . . . 4-19

4-3 Chemical Compositions (%wt) of the Metal Wires Used in the MultielectrodeArray Sensor Probes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-21

4-4 Pitting Resistance Equivalent (PRE) Numbers Calculated According to theElemental Composition of the Multielectrode Array Sensor Alloys . . . . . . . . . . . . . . 4-23

5-1 Parameters to Simulate the Corrosion Potential As a Function of Temperature andpH . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5-6

5-2 Parameters and Distributions Used to Define the Critical Potential As a Function ofthe Temperature and Chloride Concentration in the TPA Code . . . . . . . . . . . . . . . . . 5-9

5-3 Values of the Reference Ratios to Compute the Term )Ercrev in the TPA CodeAbstraction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5-10

5-4 Rank Correlations to Be Used in a Performance Assessment Abstraction . . . . . . . 5-14

xv

ACKNOWLEDGMENTS

This report was prepared to document work performed by the Center for Nuclear WasteRegulatory Analyses (CNWRA) for the U.S. Nuclear Regulatory Commission (NRC) underContract No. NRC–02–02–012. The activities reported here were performed on behalf of theNRC Office of Nuclear Material Safety and Safeguards, Division of High-Level WasteRepository Safety. This report is an independent product of the CNWRA and does notnecessarily reflect the view or regulatory position of the NRC.

The authors gratefully acknowledge Mr. Brian Derby for conducting the laboratory tests, thetechnical reviews of Drs. V. Jain and S. Mohanty, the programmatic review of B. Sagar, and theeditorial review of P. Mackin. Appreciation is due J. Gonzalez for assistance in the preparationof this report.

QUALITY OF DATA: Sources of data are referenced in each chapter. CNWRA-generateddata contained in this report meet quality assurance requirements described in the CNWRAQuality Assurance Manual. Data from other sources, however, are freely used. The respectivesources of non-CNWRA data should be consulted for determining levels of quality assurance.

ANALYSES AND CODES: Mathematical routines for computations reported in Chapter 5 wereprogrammed in Mathematica 4.1 (Wolfram Research, Inc., 1999) and are included in qualityassurance records accompanying this report. The computer software LabAnalyzer Version 1.3and StreamAnalyzer Version 1.3 were used in the analyses contained in this report. LabAnalyzer Version 1.3 and StreamAnalyzer Version 1.3 (not under CNWRA configurationcontrol) are commercial software and only object codes are available to the CNWRA.Corresponding electronic result files are also part of the quality assurance records. Documentation can be found in scientific notebook numbers 185E, 288, 366, 485, 498, 505,520, 534, 578, 607, 611, 617, 637, 638, 670, 672, and 659.

REFERENCE

OLI Systems, Inc. “LabAnalyzer and Stream Analyzer.” Version 1.3. Morris Plains, NewJersey: OLI Systems, Inc. 2004.

Wolfram Research, Inc. The Mathematica Book. 4th Edition. Champaign, Illinois: WolframMedia and Cambridge University Press. 1999.

xvii

EXECUTIVE SUMMARY

The U.S. Department of Energy is preparing a license application for permanent disposal ofhigh-level radioactive waste at Yucca Mountain, Nevada. Radionuclide containment by theengineered barriers is important to protect the public from long-term risks. Two importantcomponents of the engineered barrier subsystem are the waste package and the drip shield.Degradation processes that can affect waste package performance include dry-air oxidation,uniform corrosion, localized corrosion, and stress corrosion cracking. These degradationmodes are considered in Subissues 1, 2, and 6 of the Container Life and Source Term KeyTechnical Issue and Subissue 3 of the Total System Performance Assessment and IntegrationKey Technical Issue. Under the new work breakdown structure, all degradation modes areincorporated into the Degradation of Engineered Barriers Integrated Subissue.

The Center for Nuclear Waste Regulatory Analyses (CNWRA) is conducting an independenttechnical assessment to evaluate degradation of the Alloy 22 (Ni-22Cr-13Mo-4Fe-3W) outercontainer material by dry-air oxidation, uniform corrosion, and localized corrosion. Results ofthe independent assessment conducted at CNWRA are used to develop abstractions for wastepackage performance implemented in the Total-system Performance Assessment code. Thisreport provides results of recent CNWRA experimental work conducted to determine the effectsof dry-air oxidation, uniform corrosion rates under passive conditions, passive film chemistry,localized corrosion susceptibility, and localized corrosion propagation rates for the Alloy 22waste package outer barrier. Models for the passive corrosion rate and localized corrosionsusceptibility are developed to evaluate the effect of environmental conditions on wastepackage performance.

The consequence of dry-air oxidation over extended periods on the passive behavior of Alloy 22is analyzed via extrapolation of short-term experiments at high temperatures. Dry-air oxidationkinetics and oxide morphology studies were conducted at 850 °C [1,562 °F] and 1,100 °C[2,012 °F] for up to 120 hours. At 1,100 °C [2,012 °F], the oxidation mechanism involvesexternal Cr2O3 scale formation and evaporation, and internal oxidation. At 850 °C [1,562 °F], theoxidation mechanism is controlled predominantly by outward diffusion of chromium ions to formexternal Cr2O3 scale.

Under environmental conditions where passivity is maintained, the uniform aqueous corrosionrate of Alloy 22 is low, not strongly dependent on solution composition, and follows andArrhenius dependence on temperature with an activation energy in the range of 33.5 to49.6 kJ/mol [8.0 to 11.9 kcal/mol]. The passive dissolution rate decreased with time,attributable to improved corrosion resistance of the oxide film. Short-term passive dissolutionrates for mill-annealed Alloy 22 were measured to be in the range of 6 × 10!6 to 2 × 10!4 mm/yr[2.4 × 10!4 to 7.9 × 10!3 mpy] over the temperature range of 25 to 95 °C [77 to 203 °F]. Afterexposures of 2 months, the passive dissolution rate decreased to a value as low as3.5 × 10!5 mm/yr [1.4 × 10!3 mpy] at 95 °C [203 °F].

Characterization of the oxide films on Alloy 22 under passive and transpassive conditionsrevealed that the thickness and chemical composition of the surface films is a function ofenvironmental conditions and time. For short-term exposures (< 1 day) in a dilute chloridesolution, a thin oxide film {approximately 5.4 nm [2.12 × 10!4 mils]} was formed in the passiveregion consisting primarily of Cr2O3, whereas the surface film under transpassive conditions isthick and in the form of elongated patches. After long-term exposures (> 1 year) in

xviii

concentrated chloride solutions, the surface layers formed on Alloy 22, both inside and outsidethe crevice sites, had an outer silica deposit layer as a result of glassware dissolution. A distinctCr2O3 passive film was also identified on the crevice sites. These analyses support aconclusion that a stable Cr2O3 film controls the low corrosion rates of chromium-containingalloys such as Alloy 22.

Crevice corrosion of Alloy 22 is possible in a limited range of conditions characterized by thecorrosion potential exceeding the crevice corrosion repassivation potential. The corrosionpotential of Alloy 22 is dependent on solution pH and temperature. The crevice corrosionrepassivation potential is dependent on the metallurgical condition of the alloy, temperature,chloride concentration, and the relative concentration of inhibiting anions to the chlorideconcentration. At sufficient concentrations, defined in this report, oxyanions such as nitrate,carbonate, bicarbonate, and sulfate can inhibit localized corrosion.

Propagation of crevice corrosion can occur only when the corrosion potential remains above therepassivation potential. Corrosion potential drops as a result of the activation of crevicecorrosion were measured. The reduction in the corrosion potential occurring as a result ofcrevice corrosion initiation may promote repassivation and significantly limit both propagationrates and penetration depths. Tests using alternate methods to determine crevice corrosionpropagation rates support the application of the critical potential model for localized corrosioninitiation and propagation.

Corrosion processes that can potentially compromise the capability of the waste package tocontain radioactive waste are related to the establishment of environmental conditions leadingto crevice corrosion or loss of passivity. Seepage waters are naturally dilute. Thousands ofyears after repository closure, when the thermal output from the waste package is not sufficientto cause significant evaporation, in-drift waters are likely to remain dilute and unlikely to inducecrevice corrosion. Also, when the waste package thermal output is high enough to move wateraway from the drift (within the first thousand years after repository closure), crevice corrosion isnot likely because of the abundance of nitrate as evidenced by dust samples gathered in theYucca Mountain vicinity (nitrate is an effective localized corrosion inhibitor). Crevice corrosioncould occur if seepage contacts the waste package and evaporates sufficiently to yieldconcentrated solutions. Probability distribution functions for concentrations of relevant ionicspecies and pH under these conditions are proposed in this report to estimate compositions ofsolutions potentially contacting waste packages (assuming waste packages are not fullyprotected from seepage by drip shields). This report suggests correlation coefficients forstochastic sampling of these parameters in performance assessment modeling. If seepagecontacts waste packages at temperatures close to 110 °C [230 °F], a 26-percent probability forthe waste package welded area and a 3-percent probability for the mill-annealed wastepackage surface to exhibit crevice corrosion, respectively is estimated.

1-1

1 INTRODUCTION

Proper performance of the engineered barriers after waste emplacement is extremely importantto protect the public from unacceptable long-term risk, as recognized by the U.S. Department ofEnergy (DOE) (DOE, 2002). As an independent regulatory agency, the U.S. NuclearRegulatory Commission (NRC) has published regulations for disposal of high-level wastes in thepotential repository at Yucca Mountain, Nevada (10 CFR Part 63). 10 CFR Part 63 requires thatthe engineered barrier subsystem must be designed so that, working in combination with naturalbarriers, radiological exposures to the reasonably maximally exposed individual are within thelimits specified in 10 CFR 63.113. Therefore, performance of the waste package and dripshield, the two main components of the engineered barrier subsystem, are considered importantto long-term performance (DOE, 2002). The reference waste package design in the DOE siterecommendation (DOE, 2002) consists of an outer container of highly corrosion-resistantnickel-chromium-molybdenum alloy, Alloy 22 (Ni-22Cr-13Mo-4Fe-3W), and an inner container ofType 316 nuclear grade stainless steel (low C-high N-Fe-18Cr-12Ni-2.5Mo). Additionally, aninverted U-shaped drip shield, fabricated from Titanium Grade 7 (Ti-0.15Pd) and TitaniumGrade 24 (Ti-6Al-4V-0.08Pd), would extend the length of the emplacement drifts to enclose thetop and sides of the waste packages. For undisturbed repository conditions, corrosion isexpected to be the primary degradation process limiting the life of the waste packages and thedrip shields. Loss of containment as a result of corrosion would allow the release ofradionuclides to the environment surrounding the waste packages.

The corrosion-related processes considered important in the potential degradation of the wastepackage and the drip shield include dry-air oxidation, humid-air and uniform aqueous (general)corrosion, localized (pitting, crevice, and intergranular) corrosion, microbially influencedcorrosion, stress corrosion cracking, and hydrogen embrittlement. The Alloy 22 proposed forthe waste package outer container is a corrosion resistant alloy that forms a chromium-richoxide film that is stable over a wide range of environmental conditions. Under conditions wherethis passive film is stable the uniform corrosion rate of the waste package is primarily dependenton environmental factors such as temperature and the chloride concentration of solutions incontact with the waste package. Localized corrosion of Alloy 22 can be initiated in oxidizingchloride solutions at elevated temperatures. In addition, the fabrication processes can influencethe localized corrosion susceptibility of Alloy 22. This report focuses on the dry-air oxidation,uniform, and localized corrosion of the Alloy 22 waste package outer container. Models for theperformance of the proposed Alloy 22 outer container are based on the results of tests tomeasure the dry-air oxidation, uniform corrosion rates, and localized corrosion susceptibility asa function of environmental and metallurgical conditions.

1.1 Objective

The Center for Nuclear Waste Regulatory Analyses (CNWRA) is developing performanceassessment models for the degradation of the engineering barriers to support NRC prelicensingactivities on topics important to the postclosure performance of the potential repository. Modelparameters are based on the results of independent evaluations of uniform corrosion rate andlocalized corrosion susceptibility of Alloy 22. An assessment of the passive corrosion rates ofcontainer materials is provided in Pensado, et al. (2002). The effects of fabrication processeson the uniform and localized corrosion of Alloy 22 have been reported by Dunn, et al. (2003a). This report summarizes the uniform corrosion rate and localized corrosion susceptibility models

1-2

used to assess the proposed waste package outer container and presents results of recentexperimental work conducted at CNWRA.

1.2 Scope and Organization of the Report

Degradation processes potentially important in the degradation of the engineered barriers havebeen reviewed in the Integrated Issue Resolution Status Report (NRC, 2004). Fabricationprocesses may alter the mechanical properties, the range of passive film stability, and thelocalized corrosion resistance of the Alloy 22 outer container, which could lead to earlythrough-wall penetration of the waste package.

This report is focused on dry-air oxidation, uniform corrosion, and localized corrosion and isorganized into seven chapters including this Chapter 1 introduction. Dry oxidation of Alloy 22 isdiscussed in Chapter 2, which includes results of recent dry oxidation rate measurements andan assessment of intergranular oxidation of Alloy 22. Chapter 3 summarizes passive corrosionrate measurements and characterization results for the passive oxide films. Chapter 4 providesresults of tests performed to characterize localized corrosion susceptibility and propagation oflocalized corrosion for Alloy 22. Chapter 5 discusses models for the performance of theproposed waste package for undisturbed conditions. Chapter 6 summarizes conclusions andrecommendations for future technical assistance prior to the submission by DOE of anypotential Yucca Mountain repository license application, and referenced literature is listed inChapter 7.

1.3 Relevant DOE and NRC Agreements

Degradation processes for waste packages are considered in Subissues 1 and 2 of theContainer Life and Source Term Key Technical Issue (NRC, 2001) and are incorporated in theDegradation of Engineering Barriers Integrated Subissue (NRC, 2004). Through prelicensingconsultation for issue resolution between DOE and NRC, these two subissues are consideredclosed-pending. Agreements pertaining to degradation of container materials and their statusare listed in Table 1-1.

1-3

Table 1-1. DOE and NRC Agreements Related to This Report

Agreement Agreement Statement Current Status

CLST.1.01* Provide the documentation for Alloy 22 and titanium for thepath forward items listed on slide 8 [establish crediblerange of brine water chemistry; evaluate effect ofintroduced materials on water chemistry; determine likelyconcentrations and chemical form of minor constituents inYM waters; characterize YM waters with respect to theparameters which define the type of brine which wouldevolve; evaluate periodic water drip evaporation] DOE willprovide the documentation in a revision to AMR“Environment on the Surfaces of the Drip Shield and WastePackage Outer Barrier” by LA.

Complete

CLST.1.02* Provide the documentation for the path forward items listedon slide 12. (Surface elemental analysis of alloy testspecimens is necessary for determination of selectivedissolution; surface analysis of welded specimens forevidence of dealloying; continue testing including simulatedsaturated repository environment to confirm enhancementfactor). DOE will provide the documentation in a revision tothe Analysis Model Report General and LocalizedCorrosion of Waste Package Outer Barrier by licenseapplication.

Complete

CLST.1.03* Provide documentation that confirms the linear polarizationresistance measurements with corrosion ratemeasurements using other techniques. DOE will providethe documentation in a revision to AMR “General andLocalized Corrosion of Waste Package Outer Barrier” byLA.

Complete

CLST.1.04* Provide the documentation for Alloy 22 and titanium for thepath forward items listed on slide 14 [continue testing in theLTCTF; add new bounding water test environments toLTCTF (SSW & BSW); install thinner coupons in LTCTFwith larger surface area/volume rations; install highsensitivity probes of Alloy 22 in some of the LTCTFvessels; materials testing continues during performanceconfirmation] DOE will provide the documentation in arevision to AMR “ANL–EBS–MD–000003 andANL–EBS–MD–000004" by LA.

Complete

1-4

Table 1-1. DOE and NRC Agreements Related to This Report (continued)


CLST.1.05* Provide additional details on sensitivities, resolution ofmeasurements, limitations, and deposition of silica for thehigh sensitivity probes. DOE will document the results ofthe sensitivity probes including limitation and resolution ofmeasurements as affected by silica deposition in the Alloy22 AMR and Ti Corrosion AMR (ANL–EBS–MD–000003and ANL–EBS–MD–000004) prior to LA.

Complete

CLST.1.06* Provide the documentation on testing showing corrosionrates in the absence of silica deposition. DOE willdocument the results of testing in the absence of silicadeposits in the revision of Alloy 22 AMR(ANL–EBS–MD–000003) prior to LA.

Complete

CLST.1.07* Provide the documentation for the alternative methods tomeasure the corrosion rate of the waste package material(e.g., ASTM G–102 testing) or provide justification for thecurrent approach. DOE will document the alternativemethods of corrosion measurement in the revision of Alloy22 AMR (ANL–EBS–MD–000003), prior to LA.

Complete

CLST.1.08* Provide the documentation for Alloy 22 and titanium for thepath forward items listed on slide 16 and 17 [calculatepotential-pH diagrams for multi-component Alloy 22; growoxide films at higher temperatures in autoclaves, in airand/or electrochemically to accelerate film growth forcompositional and structural studies below; resolve kineticsof film growth: parabolic or higher order, whether filmgrowth becomes linear, and if, as film grows it becomesmechanically brittle and spalls off; determine chemical,structural, and mechanical properties of films, includingthicken films; correlate changes in Ecorr measured inLTCTF with compositional changes in passive film overtime; perform analyses on cold-worked materials todetermine changes in film structural properties; performexamination of films formed on naturally occurringJosephinite; compare films formed on Alloy 22 with othersimilar passive film Alloys with longer industrialexperience]. DOE will provide the documentation in arevision to AMRs (ANL–EBS–MD–000003 andANL–EBS–MD–000004) prior to LA.”

Complete



1-5

CLST.1.09* Provide the data that characterize the passive film stability,including the welded and thermally aged specimens. DOEwill provide the documentation in a revision to Analysis andModel Reports (ANL–EBS–MD–000003 andANL–EBS–MD–000004) prior to license application.

Complete

CLST.1.10* Provide the documentation for Alloy 22 and titanium for thepath forward items listed on slide 21 and 22 [measurecorrosion potentials in the LTCTF to determine any shift ofpotential with time toward the critical potentials for LC;determine critical potentials on welded and welded andaged coupons of Alloy 22 vs those for basemetal—particularly important if precipitation or severesegregation of alloying elements occurs in the welds;separate effects of ionic mix of specimens in YM waters oncritical potentials—damaging species from potentiallybeneficial species; determine critical potentials inenvironments containing heavy metal concentrations] DOEwill provide the documentation in a revision to AMRs(ANL–EBS–MD–000003 and ANL–EBS–MD–000004) priorto LA.

Complete

CLST.1.11* Provide the technical basis for the selection of the criticalpotentials as bounding parameters for localized corrosion,taking into account MIC. DOE will provide thedocumentation in a revision to AMRs(ANL–EBS–MD–000003 and ANL–EBS–MD–000004) priorto LA.

Complete



1-6

CLST.1.12* Provide the documentation for Alloy 22 and titanium for thepath forward items listed on slides 34 and 35 [qualify andoptimize mitigation processes; generate SCC data formitigated material over full range of metallurgicalconditions; new vessels for LTCTF will house many of theSCC specimens; continue SSRT in same types ofenvironments as above, specimens in the same range ofmetallurgical conditions; determine repassivation constantsneeded for film rupture SCC model to obtain value for themodel parameter ‘n’; continue reversing direct currentpotential drop crack propagation rate determinations insame types of environments and same metallurgicalconditions as for SSRT and LTCTF tests; evaluate SCCresistance of welded and laser peened material vs non-welded unpeened material; evaluate SCC resistance ininduction annealed material; evaluate SCC resistance offull thickness material obtained from the demonstrationprototype cylinder of Alloy 22] DOE will provide thedocumentation in a revision to AMRs(ANL–EBS–MD–000005 and ANL–EBS–MD–000006) priorto LA.

Complete

CLST.1.13* Provide the data that characterize the distribution ofstresses due to laser peening and induction annealing ofAlloy 22. DOE will provide the documentation in a revisionto AMR (ANL–EBS–MD–000005) prior to LA.

Complete

CLST.1 15* Provide the documentation for Alloy 22 and titanium for thepath forward items listed on slide 39 [install specimens cutfrom welds of SR design mockup in LTCTF and in otherSCC test environments—determine which specimengeometry is most feasible to complement SCC evaluation;evaluate scaling and weld process factors between thincoupons and dimensions in actual welded waste packagecontainers—including thermal/metallurgical structuraleffects of multi-pass weld processes; providerepresentative weld test specimens for MIC work, thermalaging and localized corrosion evaluations] DOE will providedocumentation for Alloy 22 and Ti path forward items onslide 39 in a revision to the SCC and general and localizedcorrosion AMRs (ANL–EBS–MD–000003,ANL–EBS–MD–000004, ANL–EBS–MD–000005) by LA.

Complete



1-7

CLST.1.16* Provide the documentation on the measured thermal profileof the waste package material due to induction annealing. DOE stated that the thermal profiles will be measuredduring induction annealing, and the results will be reportedin the next SCC AMR (ANL–EBS–MD–000005) prior to LA.

Complete

CLST.2.04* Provide information on the effect of the entire fabricationsequence on phase instability of Alloy 22, including theeffect of welding thick sections using multiple weld passesand the proposed induction annealing process. DOEstated that the aging studies will be expanded to includesolution annealed and induction annealed Alloy 22 weldand base metal samples from the mock-ups as well aslaser peened thick, multi-pass welds. This information willbe included inrevisions of the AMR Aging and PhaseStability of the Waste Package Outer Barrier,ANL–EBS–MD–000002, before LA.

Complete

CLST.2.05* Provide the Aging and Phase Stability of Waste PackageOuter Barrier, AMR, including the documentation of thepath forward items listed in the Subissue 2: Effects ofPhase Instability of Materials and Initial Defects on theMechanical Failure and Lifetime of the Containerspresentation, slides 5 & 6. [data input to current models isbeing further evaluated and quantified to reduceuncertainty; aging of Alloy 22 samples for microstructuralcharacterization, tensile property test, and Charpy impacttest is ongoing; theoretical modeling will be employed toenhance confidence in extrapolating aging kinetic data torepository thermal conditions and time scale—modeling willutilize thermodynamic principles of the processes; Alloy 22samples for SCC compact tension test are being added toaging studies; test program will be expanded to includewelded and cold worked materials; effects of stressmitigation techniques such as laser peening and inductionannealing on phase instability will be investigated; agingtest facility will be expanded to include aging at lowertemperatures] DOE stated that the Aging and PhaseStability of the Waste Package Outer Barrier AMR,ANL–EBS–MD–000002, Rev. 00 was issued 3/20/00. ThisAMR will be revised to include the results of the pathforward items before LA.

Complete



1-8

CLST.2.07* Provide documentation for the fabrication process,controls, and implementation of the phases which affect theTSPA model assumptions for the waste package (e.g., fillermetal, composition range). DOE stated that updates of thedocumentation on the fabrication processes and controls(TDR–EBS–ND–000003, Waste Package OperationsClosure Weld Tech. Guidelines Fabrication Process Reportand TDP–EBS–ND–000005, Waste Package OperationsFY-00 Closure Weld Technical Guidelines Document) willbe available to the NRC in January 2001.

Complete

TSPAI.3.01† Propagate significant sources of uncertainty intoprojections of waste package and drip shield performanceincluded in future performance assessments. Specificsources of uncertainty that should be propagated (or strongtechnical basis provided as to why it is insignificant)include: (1) the uncertainty from measured crevice andweight-loss samples general corrosion rates and thestatistical differences between the populations, (2) theuncertainty from alternative explanations for the decreasein corrosion rates with time (such as silica coatings thatalter the reactive surface area), (3) the uncertainty fromutilizing a limited number of samples to define thecorrection for silica precipitation, (4) the confidence in theupper limit of corrosion rates resulting from the limitedsample size, and (5) the uncertainty from alternativestatistical representations of the population of empiricalgeneral corrosion rates. The technical basis for sources ofuncertainty will be established upon completion of existingagreement items CLST.1.4, 1.5, 1.6, and 1.7. DOE willthen propagate significant sources of uncertainty intoprojections of waste package and drip shield performanceincluded in future performance assessments. Thistechnical basis will be documented in a future revision ofthe General and Localized Corrosion of Waste PackageOuter Barrier AMR (ANL–EBS–MD–000003) expected tobe available consistent with the scope and schedules forthe specified CLST agreements. The results of the AMRanalyses will be propagated into future TSPA analyses forany potential license application.

Complete



1-9

TSPAI.3.04† Provide the technical basis that the representation of thevariation of general corrosion rates (if a significant portionis “lack of knowledge” uncertainty) does not result in riskdilution of projected dose responses (ENG1.3.3). DOE willprovide the technical basis that the representation of thevariation of general corrosion rates results in reasonablyconservative projected dose rates. The technical basis willbe documented in an update to the WAPDEG Analysis ofWaste Package and Drip Shield Degradation AMR(ANL–EBS–PA–000001). This AMR is expected to beavailable to NRC in FY 2003. These results will beincorporated into future TSPA documentation for anypotential license application.

Complete

TSPAI.3.05† Provide the technical basis for the representation ofuncertainty/variability in the general corrosion rates inrevised documentation. This technical basis shouldprovide a detailed discussion and analyses to allowindependent reviewers the ability to interpret therepresentations of 100% uncertainty, 100% variability, andany intermediate representations in the DOE model(ENG1.3.6). DOE will provide the technical basis for therepresentation of uncertainty/variability in the generalcorrosion rates. This technical basis will include the resultsof 100% uncertainty, 100% variability, and selectedintermediate representations used in the DOE model. These results will be documented in an update to theWAPDEG Analysis of Waste Package and Drip ShieldDegradation AMR (ANL–EBS–PA–000001) or otherdocument. This AMR is expected to be available to NRC inFY 2003.

Complete

*Schlueter, J.R. “U.S. Nuclear Regulatory Commission/U.S. Department of Energy Technical Exchange andManagement on Container Life and Source Term (September 12–13, 2000).” Letter (October 4) to S. Brocoum,DOE. Washington, DC: NRC. 2000.†Reamer, C.W. “U.S. Nuclear Regulatory Commission/U.S. Department of Energy Technical Exchange andManagement on Total System Performance Assessment and Integration (August 6–10, 2001).” Letter (August 23)to S. Brocoum, DOE. Washington, DC: NRC. 2001.

2-1

2 PASSIVE AND LOCALIZED CORROSION OF OVERPACKMATERIALS—MODELING AND EXPERIMENTS

2.1 Dry-Air Oxidation

Dry-air oxidation is assumed to occur during the initial period after waste package emplacementwhen the radioactive decay heat keeps moisture away from the emplacement drifts (BechtelSAIC Company, LLC, 2003a). According to the DOE model (Bechtel SAIC Company, LLC,2003a; Welsch, et al., 1996), dry-air oxidation of Alloy 22 results in the formation of an adherent,protective Cr2O3 film of uniform thickness. The Cr2O3 film growth rate, oxidation time, andtemperature relationship are described in Eqs. (2-1) through (2-3). The Cr2O3 film growth ratefollows a parabolic rate law, where the film thickness or mass gain is proportional to the squareroot of time

Δm/A = (Kpt)½ + C (2-1)where

Δm/A — mass gain per unit area in time, t (mg cm!2)Kp — parabolic rate constant (mg2cm!4h!1)C — a constant (mg cm!2)

The temperature dependence of the oxidation rate is contained in the parabolic rate constant

Kp = Ko exp (!Qoxid /RT) (2-2)

where

Ko — a constant (mg2cm!4h!1)Qoxid — activation energy for oxide growth (J mol!1)R — the ideal gas constant (J mol!1 K!1)T — absolute temperature (K)

Taking the logarithm on both sides of Eq. (2-2), the equation can be rewritten as

Log Kp = Log Ko - Qoxid /RT (2-3)

The activation energy for oxide growth Qoxid can be determined experimentally by performingoxidation tests at elevated temperatures. The temperature dependence of Kp is Arrhenius type. For the Arrhenius temperature relationship, the oxidation rate of Alloy 22 for long timeexposures at lower temperatures can be extrapolated from short-term tests athigher temperatures.

In this investigation the oxidation kinetics and oxide morphology of Alloy 22 in dry-air werestudied at elevated temperatures of 850 °C [1,562 °F] and 1,100 °C [2,012 °F] and extrapolatedto expected repository temperatures. Special attention was paid to the phases of oxides thatformed, as well as the extent of outward diffusion of chromium and inward penetration ofoxygen. The results are analyzed and the potential effect of dry-air oxidation on waste packagelife is discussed.

1Cahn TherMax, 700 Microbalance, Thermo Electron Corporation, Newington, New Hampshire.

2JEOL, JXA–8600 Electron Microprobe, JEOL USA, Inc., Peabody, Massachusetts.

2-2

The purpose of this work was to identify if there are correlations that allow oxide growth rate anddepth of oxygen penetration to be estimated as a function of time and temperature. These testsenhance fundamental understanding of oxide film properties during the thermal pulse periodand, on subsequent corrosion in an aqueous environment.

2.2 Experimental

The composition of the alloys used in this study is shown in Table 2-1. The oxidation testspecimens were machined from mill-annealed, 12.7-mm [0.5-in]-thick plate stock. The platematerial was confirmed to meet the chemical composition specified in ASTM B–575 forUNS N06022 alloy designation (ASTM International, 2004a). The machined plate was cut intooxidation coupons of dimension 15 × 13 × 1.5 mm [0.6 × 0.5 × 0.06 in]. Specimens were wetsurface ground with 600-grit silicon carbide paper and ultrasonically cleaned in a acetone andalcohol baths before oxidation exposure.

The oxidation tests were performed with a microbalance1 that enabled continuous monitoring ofweight change and temperature. After oxidation, cross sections of the specimens wereexamined optically and with an electron microprobe2 and associated wavelength x-rayspectrometers. In this analysis, a backscattered electron image was taken of the surface oxidelayer and the internal oxygen penetration region. Quantitative line scans for oxygen, chromium,iron, nickel, molybdenum, and tungsten were then run starting at the surface and penetrating in2 :m [78.7 :in] increments into the base material. For selected samples, x-ray dot maps of theelements also were made to show the elemental distribution in the oxide/alloy interface region. In this analysis, the detector was tuned to respond to x-ray wavelengths characteristic ofoxygen, chromium, nickel, molybdenum, iron, and tungsten. The density of the white dotstherefore provided information on the abundance of each element in the examined areas. Theintensities in the x-ray maps do not directly correlate quantitatively with element concentrations.

2.3 Oxidation Rate Measurements

The oxidation rate data for Alloy 22 specimens subjected to isothermal oxidation in air at 850 °C [1,562 °F] and 1,100 °C [2,012 °F] are presented in Figure 2-1(a) as mass change versus time.The oxidation mass change at a given exposure time is greater at 1,100 °C [2,012 °F] than at850 °C [1562 °F]. The results are expected because the oxidation rate is higher at highertemperatures. At both temperatures, the alloy exhibited a much more rapid growth rate duringan initial transient period of approximately 1.5 hours. The rate slowed at longer times. Themass change data are also plotted as mass change versus the square root of time inFigure 2-1(b). For the 850 °C [1,562 °F] specimens, the data fall along a straight line indicatinga parabolic rate law for oxidation kinetics.

The mass gain during oxidation is the result of the external oxide formation and the mass ofoxygen absorbed in the alloy. Wagner developed the theory of high temperature oxidation ofmetals under idealized conditions (Wagner, 1959). The theory describes the oxidation behaviorfor the case where the diffusion of ions in the metal is rate limiting, the oxidation follows a

2-3

Table 2-1. Composition of Alloys 22, 625, and 825 Heats and Alloy 622 Filler Metal

Material Ni* Cr Mo W Fe Co Si Mn V P S C OtherAlloy 22 Heat2277-8-3175

12.7-mm 57.8 21.40 13.60 3.00 3.80 0.09 0.030 0.12 0.15 0.008 0.002 0.004 —

Alloy 22 Heat2277-8-3235

12.7-mm Bal 21.40 13.47 2.87 3.94 1.31 0.023 0.24 0.17 0.008 0.001 0.003 —

622 Filler Heat

XX1045BG11Bal 20.73 14.13 3.15 3.05 0.09 0.060 0.24 0.01 0.007 0.001 0.006 —

Alloy 22 Heat059902LL238.1-mm

Bal 20.35 13.85 2.63 2.85 0.01 0.05 0.16 0.17 0.007 0.0002 0.005 —

622 Filler Heat

XX2048BGBal 20.48 14.21 3.02 2.53 0.02 0.07 0.20 0.02 0.009 <0.001 0.001 —

Alloy 22 Heat 2277-3-3266

12.7 mmBal 21.40 13.30 2.81 3.75 1.19 0.03 0.23 0.14 0.008 0.004 0.005 —

Alloy 22 Heat 2277-1-3164

25.4 mmBal 21.15 13.47 3.26 3.93 1.27 0.02 0.23 0.11 0.007 0.003 0.003 —

ERNiCrMo-10filler heatWN813 Bal 22.24 13.7 3.13 2.37 0.41 0.02 0.34 0.01 0.003 0.001 0.003 —

ERNiCrMo-10filler heat

XX1977BG11Bal 20.25 14.13 2.99 2.56 0.07 0.06 0.20 0.04 0.008 0.001 0.005 —

Alloy 625NX9936AG12.7-mm

60.9 21.70 9.01 — 3.96 0.15 0.16 0.08 — 0.009 0.001 0.02Nb: 3.48Al: 0.24Ti: 0.24

Alloy 825HH4371FG12.7-mm

41.06 22.09 3.21 — 30.41 — 0.29 0.35 — — — 0.010Cu: 1.79Al: 0.07Ti: 0.82

*Ni—nickel, Cr—chromium, Mo—molybdenum, W—tungsten, Fe—iron, Co—cobalt, Si—silicon, Mn—manganese,V—vanadium, P—phosphorus, S—sulfur, C—carbon

2-4

Figure 2-1. Oxidation Kinetics of Alloy 22 in Air at 850 °C[1,562 °F] and 1,100 °C [2,012 °F]: (a) Mass Change Versus Time

Data; (b) Mass Change Because of Oxidation Versus Square Rootof Time Data

2-5

parabolic rate law. The parabolic rate constant [the square of the slope from Figure 2-1(b)] for Alloy 22 at 850 °C [1,562 °F] is 5.1 × 10!4 mg2/cm4h. At 1,100 °C [2,012 °F], however, the datafor mass gain versus square root of time do not plot as a straight line. The slope of the plotdecrease with increasing exposure time, indicating a complex oxidation mechanism.

2.4 Oxidation Morphology and Assessment of Intergranular Oxidation

2.4.1 Scale Morphology and Composition at 850 °C [1,562 °F]

A cross-section view of the oxides that formed after exposure for 120 hours at 850 °C [1,562 °F]is shown in Figure 2-2(a). The elemental line scans of oxygen, chromium, iron, nickel,molybdenum, and tungsten are presented in Figure 2-2(b). The external scale consistedprimarily of chromium and oxygen, indicating the oxide is Cr2O3. The Cr2O3 is approximately4 :m [157.4 :in] thick. In several localized areas, internal oxide precipitates were also observedunderneath the external scale. These internal oxides appeared as dark spots in thebackscattered electron micrograph in Figure 2-2(a). The maximum depth of inward oxygenpenetration is approximately 4 :m [157.4 :in]. The composition profiles also show that there isa chromium depletion zone of about 8 :m [314.9 :in] underneath the external scale. Thechromium content in the chromium depletion zone is about 16 wt% [Figure 2-2(b)], which islower than the 21.4 wt% chromium in the bulk alloy (Table 2-1). There is a chromiumconcentration gradient underneath the external Cr2O3 scale which is the driving force foroutward chromium diffusion. Nickel is slightly enriched in the chromium depletion zone becauseof the chromium deficit. No significant diffusion of iron, molybdenum, or tungsten was observed.

2.4.2 Scale Morphology and Composition at 1,100 °C [2,012 °F]

A cross-section view of the oxides that formed after exposure for 24 hours at 1,100 °C[2,012 °F] is shown in Figure 2-3(a). The elemental line scans of oxygen, chromium, iron,nickel, molybdenum, and tungsten are presented in Figure 2-3(b). A continuous layer of Cr2O3was formed [see concentration profiles of chromium and oxygen near the surface inFigure 2-3(b)]. The thickness of the external scale ranged from 2–4 :m [78.7–157.4 :in]. Oxideprecipitates were also observed along grain boundaries deep into the alloy. The chromiumconcentration around the oxide precipitates is estimated to be 14 wt%. Between 10 to 20 :m[393.7 to 787.4 :in] from the surface, there is a chromium concentration gradient, which is thedriving force for outward chromium diffusion. The maximum depth of inward oxygen penetrationis about 10 :m [393.7 :in].

The oxides that formed on Alloy 22 after the longer time exposure of 120 hours at 1,100 °C[2,012 °F] are shown in Figure 2-4(a). The elemental line scans of oxygen, chromium, iron,nickel, molybdenum, and tungsten are presented in Figure 2-4(b). The chromium and oxygenline scan show that the thickness of the external Cr2O3 scale increased to approximately 14 :m[551.2 :in]. As the measuring probe hit a precipitate in the alloy, an increase in oxygen signalwas detected. This indicates that the precipitates observed in Figure 2-4(a) are internallyoxidized particles. The maximum depth of oxygen penetration increased to 24 :m [944.9 :in]for this higher temperature.

2-6

Figure 2-2. Cross-Section Through the Scale Produced on Alloy 22After 120 Hours Oxidation at 850 °C [1,562 °F]: (a) Scanning

Electron Microscope Backscattered Electron Image; and(b) Concentration Profiles of Oxygen, Chromium, Iron, Nickel,

Molybdenum, and Tungsten Along the Line of Traverse Indicated inFigure 2-2(a).

2-7

Figure 2-3. Cross-Section Through the Scale Produced on Alloy 22After 24 Hours Oxidation at 1,100 °C [2,012 °F]: (a) Scanning

Electron Microscope Backscattered Electron Image; and(b) Concentration Profiles of Oxygen, Chromium, Iron, Nickel,

Molybdenum, and Tungsten Along the Line of Traverse Indicated inFigure 2-3(a)

2-8

Figure 2-4. Oxide Morphology for Alloy 22 Oxidized in Air for 120 Hours at1,100 °C [2,012 °F]: (a) Scanning Electron Microscope BackscatteredElectron Image; and (b) Concentration Profiles of Oxygen, Chromium,Iron, Nickel, Molybdenum, and Tungsten Along the Line of Traverse

Indicated in Figure 2-4(a)

2-9

To complete the description of oxide morphology, oxygen, chromium, nickel, molybdenum, ironand tungsten x-ray maps are presented in Figure 2-5. The region examined is the same areashown in Figure 2-4(a). The external scale is a continuous layer of Cr2O3 in contact with the gasenvironment. The inner scale region is not uniform and is composed of Cr2O3 scale withisolated islands of unoxidized alloy. The unoxidized alloy contains the metallic elementschromium, iron, nickel, molybdenum, and tungsten, but no oxygen. Numerous internal oxideprecipitates are clearly seen underneath the external Cr2O3 scale.

2.5 Analysis of Results and Discussion

Dry-air oxidation of an Alloy 22 engineering barrier is expected to occur when the relativehumidity of the repository is less than a critical relative humidity for the initiation of humid-aircorrosion. If no drift degradation occurs prior to the thermal pulse, dry-air oxidation is expectedto occur beyond 1,000 years. The maximum waste package temperature is expected to beapproximately 236 °C [457 °F] if drift degradation occurs (Fedors, et al., 2004; Manepally, et al.,2004). The rate of dry-air oxidation can be modeled by assuming that mass transport ofreacting species is limited by solid-state diffusion through a tightly adherent passiveCr2O3 oxide film.

The oxidation kinetics at 850 °C [1,562 °F] follow a parabolic rate law. The oxidationmechanism is likely predominantly controlled by the outward diffusion of chromium to form theexternal Cr2O3 film. The estimated oxidation rate is consistent with the dry-air oxidation model(Bechtel SAIC Company, LLC, 2003a).

Another mode of oxidation degradation of alloys in dry-air environment at elevated temperaturesis internal oxidation. Oxygen may diffuse inward and form internal oxides in the alloy matrix orform internal precipitates along grain boundaries. Formation of internal oxides along grainboundaries, also known as intergranular oxidation, has been reported for Fe-21Cr-32Ni alloyafter oxidation at 900 °C [1,652 °F] for 3,000 hours (Ahn, 1996; Shida, et al., 1992a,b).

Internal oxidation and intergranular oxidation is more evident at 1,100 °C [2,012 °F] than at850 °C [1,562 °F]. During early oxidation stages, chromium diffuses outward to form externalCr2O3 scale at the oxide/gas interface. Oxygen also diffuses inward to form oxide precipitates inthe alloy matrix and grain boundaries resulted in metallurgical changes in the region underneaththe external scale. Subsequent oxidation involves further oxidation around oxide precipitates inthe chromium depletion zone. At 1,100 °C [2,012 °F], thinning of Cr2O3 scale by vapor loss issignificant at the oxide/gas interface (Birks and Meier, 1983; Gulbransen and Jansson, 1970). The thinning of Cr2O3 scale by vapor loss may contribute to the decreases in parabolic rateconstant measurements in oxidation kinetics studies.

The oxidation test results in this study are consistent with the DOE model (Bechtel SAICCompany, LLC, 2003a) that indicate Alloy 22 forms an external Cr2O3 scale during dry-airoxidation with a diffusion-controlled kinetics. However, the DOE model does not include internaloxidation. Internal oxidation was observed in the current studies at both temperatures of1,100 °C [2,012 °F] and 850 °C [1,562 °F]. The amounts of internal oxide precipitatesdecreased markedly as the temperature decreased from 1,100 °C [2,012 °F] to 850 °C[1,562 °F] (see Figures 2-2a and 2-4a). There is not yet sufficient information to extrapolate theinternal oxidation results to lower temperatures over extended times. Tests are planned tocollect oxide growth rate and internal oxidation data over a range of temperatures. An

2-10

Figure 2-5. Distribution of Oxygen, Chromium, Iron, Molybdenum, and Tungsten in theArea Outlined in Figure 2-4a: (a) Oxygen X-Ray Map; (b) Chromium X-Ray Map;

(c) Nickel X-Ray Map; (d) Molybdenum X-Ray Map; (e) Iron X-Ray Map; and(f) Tungsten X-Ray Map

2-11

assessment of significance of dry-air oxidation at repository temperatures will be made afterdata at lower temperatures have been acquired and analyzed. For example, oxide growth ratescould be estimated at higher temperatures and a trend could be established as a function oftemperature. Also, the oxygen penetration depth could be estimated as a function of time andtemperature, delineating functions that could be extrapolated to lower temperatures that arerelevant to repository conditions. These temperature-time trends could be employed to estimatethe extent of alloy change in the neighborhood of the metal-oxide interface by the time anaqueous environment is established and aqueous corrosion initiated. It is argued elsewhere(Pensado, et al., 2002) and in Chapter 5 that passive dissolution, in the presence of benignenvironmental conditions, can only be disturbed by processes inducing compositional alloychanges in the region close to the metal-oxide interface. If this alloy does not change (i.e., itscomposition remains close to the initial bulk alloy composition), passive oxides are expected toform, yielding corrosion rates comparable to those determined under the controlled conditionsdiscussed in this report. Thus, if dry-oxidation does not induce significant compositionalchanges to the alloy close to the metal-oxide interface, passive dissolution is expected to prevailafter the establishment of benign aqueous environments in contact with waste packagematerials. Estimating the extent of anticipated changes to the alloy composition in theneighborhood of the metal-oxide interface is important to enhance confidence that passivedissolution can be sustained for prolonged periods, in the absence of environmental conditionsleading to localized corrosion.

3-1

3 OXIDE FILMS AND PASSIVE DISSOLUTION

Independent investigations of the passive dissolution of waste package outer containers wereconducted at the Center for Nuclear Waste Regulatory Analyses to support development ofmodel abstractions for waste package degradation modes. The results of passive dissolutionstudies have been discussed in several recent reports and presentations (Pensado, et al., 2002;Dunn, et al., 2003a). Pensado, et al. (2002) reported passive corrosion rates of Alloy 22 measured using both potentiostatic anodic passive current density and polarization resistancemethods. Corrosion rates derived from both methods compare well for 95 °C [203 °F] data andresulted in rates of approximately 5 × 10!4 mm/yr [2 × 10!2 mil/yr]. Passive corrosion rates werefound to be independent of the composition of the solution. The potentiostatic anodic passivecurrent density at 95 °C [203 °F] decreased noticeably as a function of time. A similar decreaseis not evident at lower temperatures. At temperatures of 60 °C [140 °F] and lower, the anodicpassive current density was near the resolution limit of the instrumentation, and accuratemeasurement of the passive corrosion rates was not possible. From median values of thepotentiostatic anodic passive current density data, it was noted that the dependence of thecurrent density on the temperature is of an Arrhenius form. The apparent activation energy ofthe Arrhenius expression was computed to be 44.7 kJ/mol [10.7 kcal/mol], with a standarddeviation of 5.5 kJ/mol [1.3 kcal/mol] over the temperature range of 25 to 95 °C [77 to 203 °F]. The apparent activation energy is related to the activation energy of the rate constants forfundamental charge transfer reactions controlling the alloy dissolution. Subsequent testsreported by Dunn, et al., (2003a), indicate the passive corrosion rates are similar for bothmill-annealed, thermally aged and as-welded Alloy 22.

Information reported in this chapter includes data from previous tests (Pensado, et al., 2002;Dunn, et al., 2003a). The additional tests conducted for this report were performed to reduceuncertainty and improve understanding of the passive corrosion rates. In addition tothe electrochemical measurement of passive corrosion rates, the composition of the oxidefilm on the Alloy 22 was examined to assess the evolution of passive film compositionand characteristics.

3.1 Test and Analysis Methods

The chemical compositions of the alloys used in this study are provided in Table 2-1. Passivecorrosion rate tests were conducted using only Alloy 22. Localized corrosion tests conductedwith Alloys 825 and 625 are reported in Chapter 4. Alloy 22 was tested in both mill-annealed,as-welded and thermally aged {870 °C [1,598 °F] for 5 minutes} conditions. Short-term thermalaging was used to replicate the enhanced localized corrosion susceptibility observed in the heataffected zones of welded Alloy 22.

Passive corrosion rates were measured using electrochemical impedance spectroscopy andpotentiostatic polarization. Details of the test methods have been described in previous reports(Pensado, et al., 2002; Dunn, et al., 2003a). Potentiostatic polarization tests were conductedusing the smooth cylindrical specimens and crevice corrosion specimens shown in Figure 3-1.Electrochemical impedance tests were conducted using only the cylindrical specimens shown inFigure 3-1. All specimens were polished to a 600-grit finish, cleaned ultrasonically in detergent,rinsed in deionized water, ultrasonically cleaned in acetone, and dried. Crevice corrosionspecimens were fitted with two polytetrafluoroethylene (PTFE) crevice forming washers heldwith insulated Alloy C-276 hardware. At the completion of each test, the specimens were rinsed

3-2

48.6

mm

6.2 mm

Cylindrical Specimen

19 mm

9.5 mm

C-276 bolt

C-276 nut

C-276 washer

Serrated PTFEcrevice washer

Crevice Corrosion Specimen

19 mm

Figure 3-1. Illustration of the Cylindrical and Crevice CorrosionTest Specimens

in deionized water and dried. Most specimens were cleaned ultrasonically in an inhibitedhydrochloric acid (HCl) solution that contained 2-butyne-1, 4-diol as an inhibitor. Posttestexamination was performed with an optical microscope and a scanning electron microscope.

Tests were conducted in 2-L [0.53-gal]-glass cells with a PTFE lids. The cells were fitted with awater-cooled Allihn-type condenser and a water trap to minimize solution loss at elevatedtemperatures and air intrusion. The saturated calomel, used as a reference electrode wasconnected to the solution through a water-cooled Luggin probe with a porous silica tip and thereference electrode was maintained at room temperature. A platinum flag was used as acounter electrode. All solutions were deaerated with high-purity nitrogen (99.999 percent) for atleast 24 hours prior to the start of the tests. The anodic current density was measured inpotentiostatic tests, and the specimens were maintained at potentials ranging from !200 to800 mVSCE. The resolution of the system was 1.25 × 10!10 A/cm2 [1.16 × 10!7 A/ft2] or better. Atthe conclusion of the test, the specimens were reweighed and examined for corrosion with anoptical microscope. Corrosion rates were calculated using Eq. (3-1).

Corrosion Rate (mm / yr) = corrKi EWρ

(3-1)

where

icorr — passive corrosion current density in A/cm2

EW — equivalent weightK — conversion factor (3,270 mmAgAA!1Acm!1Ayr!1)D — density in g/cm3

3-3

For Alloy 22, D is 8.69 g/cm3 [543 lb/ft3]. Assuming congruent dissolution of the major alloyingelements Ni2+, Cr3+, Mo3+, Fe2+, and W4+ within the potential range of !200 to 400 mVSCE, theequivalent weight for Alloy 22 is 26.04 (ASTM International, 2004b).

Electrochemical impedance spectra were obtained at open circuit over a frequency range of20,000 to 0.001 Hz in chloride containing solutions at temperatures ranging from 25 to 95 °C[77 to 203 °F]. The spectra were fit to an analog equivalent circuit that included components forboth a porous outer oxide layer and an inner barrier oxide. The resistive component of theanalog circuit was used as polarization resistance to calculate the corrosion rate using theapproach originally proposed by Stern and Geary (1957) where the polarization resistance, Rp,is related to the corrosion current density, icorr, using Eq. (3-2).

R = dEdi

= . ( )iE

pa c

a c corrcorr

⎛⎝⎜

⎞⎠⎟ +

β ββ2 303 β

(3-2)

where

$a — anodic Tafel slope$c — cathodic Tafel slope

For passive metals, the anodic Tafel slope can be assumed to be infinite (Epelboin, et al., 1981)and the corrosion current density is calculated according to Eq. (3-3).

i. Rcorr

c

p=

β2 303 (3-3)

The corrosion rate can then be calculated using Eq. (3-1).

The oxide films that formed on Alloy 22 under both passive and transpassive conditions werecharacterized by x-ray photoelectron spectroscopy combined with in-depth sputtering. Theobjective was to determine changes in oxide film structure and composition as functions ofenvironmental conditions and time. Both short-term and long-term electrochemical tests wereperformed in 2-L [0.53-gal]-test cells at 95 °C [203 °F] in nitrogen-deaerated solutions using acomputer controlled potentiostat. A saturated calomel electrode was used as a referenceelectrode in all tests. The short-term potentiostatic tests were conducted in 0.028 M NaClsolutions (pH 5.5) at 100 mVSCE and 600 mVSCE for approximately 2 hours using smoothrectangle specimens polished to a 2,000 grit finish. Additional long-term tests were conductedin simulated groundwater under open-circuit conditions or potentiostatic holds at 250 mVSCE forapproximately 2 years using flat crevice specimens with 600 grit polished surfaces. Thesimulated groundwater solution had a pH of 7.5 and contained 4.0 M Cl!, 1.4 mM HCO3

!,0.20 mM SO4

2!, 0.16 mM NO3!, and 0.10 mM F!, with Na+ as the single cation. These anions

are the predominant groundwater anion species in the vicinity of Yucca Mountain.

Alloy 22 specimens were analyzed using x-ray photoelectron spectroscopy. The Al K" anodeof the x-ray source was used. The take-off angle of the photoelectrons was 45 degrees withrespect to the sample surface. For each sample, a wide-energy survey was conducted with apass energy of 89.45 eV to determine surface elemental composition. The survey was followed

3-4

pH=8.0 weldedpH=2.7 welded

pH=11.0 weldedpH=8.0 base alloy

-400 -200 0 200 400 600 800Potential, mVSCE

10-9

10-8

10-7

10-6

10-5

10-4

10-3

Ano

dic

curr

ent d

ensi

ty, A

/cm

2

Figure 3-2. Short-Term Potentiostatic Passive Dissolution RateMeasurements for Welded and Mill-Annealed Alloy 22 (Dunn, et al., 2003a)

by a high-resolution scan with a pass energy of 44.75 eV that recorded spectra from fourregions corresponding to the core levels of O 1s, Ni 2p3/2, Cr 2p3/2, and Mo 3d5/2. Theconcentration for element X was calculated from the total peak area of the correspondingcore-level peak and corrected with the sensitivity factor (Moulder, et al., 1992). Thehigh-resolution spectra were curve fitted to determine the oxidation state of nickel and chromiumand the chemical-bonding environment for oxygen. The spectrometer energy scale wascalibrated using the 84.0 eV Au 4f7/2 and 932.67 eV Cu 2p3/2 peaks with a pass energy of8.95 eV at 0.05 eV per step. Concentration depth profiles were obtained after argon ionsputtering at 3 keV. The sputtering rate was calibrated by conducting depth profiling through aSiO2 layer of 100 nm [3.94 × 10!3 mils] on silicon.

3.2 Passive Corrosion Rate Measurements

Potentiostatic measurements of passive corrosion rates reported in Pensado, et al. (2002) andDunn, et al. (2003a) are included in the results reported here to allow comparison. Based onshort-term potentiostatic tests, Dunn, et al. (2003a) reported anodic current densities in therange of 10!8 to 2 × 10!8 A/cm2 [9.3 × 10!6 to 1.9 × 10!5 A/ft2], which correspond to corrosionrates in the range of 1 × 10!4 to 3 × 10!4 mm/yr [4 × 10!3 to 1 × 10!2 mpy] at applied potentialwhere the passive oxide film is stable (Figure 3-2). As previously noted, the passive corrosionrates measured using potentiostatic methods decrease with time. Reduction of the passivecurrent density can be attributed to improvement in the passive film resistance as defects areremoved during passive dissolution. In addition, the thickness of the passive film mayincrease slightly.

3-5

0 10 20 30 40Time, days

0.0x100

1.0x10-7

2.0x10-7

3.0x10-7

4.0x10-7

5.0x10-7

6.0x10-7

7.0x10-7

8.0x10-7

9.0x10-7

1.0x10-6

Cur

rent

Den

sity

, A/c

m2

Alloy 225.5 M NaCl, 100 oC-200 mVSCE

Figure 3-3. Measured Current Density for an Alloy 22 CreviceSpecimen Under Potentiostatic Conditions (!200 mVSCE). No

Crevice Corrosion Was Initiated.

The anodic current density measured on a crevice corrosion specimen over a period of 38 daysis shown in Figure 3-3. Although the specimen was fitted with a crevice forming washer andpotentiostatically held at !200 mVSCE in a 5.5 molar sodium chloride solution at 100 °C [212 °F],no crevice corrosion was initiated. The cause of the current spike that occurred at around21days is not known. It cannot be attributed, however, to stable crevice corrosion initiation. After 30 days, the porous glass bubbler used to deaerate the test solution filled with saltprecipitate, and the nitrogen flow rate decreased and eventually stopped. Without nitrogenpurging, oxygen diffused into the test cell. The decrease in the measured anodic currentdensity that occurred in this period can be attributed to increased dissolved oxygenconcentrations (Pensado, et al., 2002). After nitrogen purging was restored, the measuredanodic current density increased to values near those measured under deaerated conditions. Neglecting the unintentional upset of the cell conditions, the decrease in anodic current densityas a function of time is apparent in Figure 3-3.

The passive corrosion rate for Alloy 22 is not expected to be influenced by the dissolved oxygenunder the potentiostatic conditions used in this test. The decrease in the measured anodiccurrent density during the time when nitrogen deaeration was not present occurred becauseonly the net current can be measured under potentiostatic conditions. In a deaerated systemwith low dissolved oxygen concentration, the net current density under potentiostatic conditionsanodic to the corrosion potential approaches the anodic dissolution rate for the test specimen,and the reduction of H+ ions occurs at the counter electrode.

3-6

0 20 40 60 80 100 120Time, days

-4.0x10-9

-2.0x10-9

0.0x100

2.0x10-9

4.0x10-9

6.0x10-9

8.0x10-9

1.0x10-8

1.2x10-8

1.4x10-8

Cur

rent

Den

sity

, A/c

m2

Alloy 225.5 M NaCl, 100 oC,-100 mVSCE

Figure 3-4. Measured Current Density for an Alloy 22 Crevice SpecimenUnder Potentiostatic Conditions (!100 mVSCE). No Crevice Corrosion

Was Initiated.

Additional data under similar conditions are shown in Figure 3-4. In this test, an Alloy 22crevice specimen was maintained under potentiostatic condition for over 100 days. Initialmeasurements at short times are similar to the anodic current densities shown in Figure 3-2;however, the anodic current density decreased with time. The abrupt decrease in the anodiccurrent density after 82 days was the result of significantly reduced nitrogen flow to the testsolution; similar to the condition described for the data shown in Figure 3-3. Similarly, afternitrogen flow to the test cell was restored, the anodic current density returned to its formervalue. Neglecting the upset conditions, the average anodic current density after 60 days was3.5 × 10!9 A/cm2 [3.3 × 10!6 A/ft2], which corresponds to a passive corrosion rate of3.5 × 10!5 mm/yr [1.4 × 10!3 mpy]. This is substantially less than the short-term passivecorrosion rates measured using either potentiostatic polarization or electrochemical impedancespectroscopy (Pensado, et al., 2002; Dunn, et al., 2003a).

A comparison of the passive corrosion rates for mill-annealed, thermally aged {870 °C[1,598 °F] for 5 minutes} and as-welded Alloy 22, determined using electrochemical impedancespectroscopy, was previously reported by Dunn, et al. (2003a) and is shown in Figure 3-5. Thecorrosion rates were calculated based on the resistance of the barrier layer oxide which wasdetermined by fitting the impedance spectra to an analog circuit with two time constants and asolution resistance (Dunn, et al., 2003a). The corrosion rates shown in Figure 3-5 arecomparable to the corrosion rates calculated form the anodic current transients (Pensado, et al.,2002). In addition, it is apparent that the corrosion rates for the welded material and thethermally aged material are slightly greater than those of the mill-annealed alloy. At 25 °C

3-7

20 40 60 80 100Temperature, oC

10-6

10-5

10-4

10-3

10-2

Cor

rosi

on ra

te, m

m/y

r

Alloy 22, 0.028 M NaCl

Mill-annealed As-welded Thermally aged 870oC/5 min

Figure 3-5. Passive Corrosion Rate for Mill-Annealed, As-Weldedand Thermally Aged Alloy 22 in 0.028 M NaCl As a Function of

Temperature. Corrosion Rates Were Obtained fromElectrochemical Impedance Spectra (Dunn, et al., 2003a).

[77 °F] the corrosion rate for the as-welded and thermally aged material was about twice that ofthe mill-annealed material. Larger differences were observed at 95 °C [203 °F].

Tests were conducted in concentrated MgCl2 solutions to reach temperatures above 100 °C [212 °F] at atmospheric pressure (Dunn, et al., 2003a). In the concentrated MgCl2 solutions, thecorrosion rates for mill-annealed Alloy 22 were more than an order of magnitude greater thanthose measured in the more dilute 0.028 molar NaCl solution (Figure 3-6). The calculatedactivation energy (Figure 3-7) is 46.3 KJ mol!1 [11.1 kcal mol!1] over the temperature range from25 to 95 °C [77 to 203 °F] in 0.028 M NaCl. The activation energy for the corrosion rate as afunction of temperature in the MgCl2 was 49.6 KJ/mol [11.9 kcal mol!1]. Although the activationenergies for the 0.028 molar NaCl and the more concentrated MgCl2 solution were similar, it isapparent from Figure 3-6 that the corrosion rates in MgCl2 at 120 °C [248 °F] deviatesubstantially from the trend observed at lower temperatures. In addition, initial measurementsconducted at 20 °C [68 °F] also deviate from the established trend. Significant variability in thecorrosion rates in the MgCl2 solution observed at 40 °C [104 °F] is a result of measurementuncertainty at this temperature. From the data shown in Figure 3-6, it is apparent that the highchloride concentrations in the concentrated magnesium chloride solutions disrupt passive filmstability and significantly accelerate the corrosion rate of Alloy 22.

The corrosion rates and corresponding activation energies measured in short-term tests do notconsider the decrease in corrosion rate that can occur through the slow refinement andimprovement of the oxide film during passive dissolution. The results of additionalelectrochemical impedance tests conducted with longer exposure times are shown in

3-8

20 40 60 80 100 120Temperature, oC

10-6

10-5

10-4

10-3

10-2

10-1

Cor

rosi

on ra

te, m

m/y

r

Mill Annealed 35% MgCl2

Mill Annealed 0.028 M NaCl

Alloy 22

Figure 3-6. Corrosion Rate of Alloy 22 in 0.028 M NaCl and35-Percent MgCl2 (7.5 M Chloride) As a Function ofTemperature. Corrosion Rates Were Obtained from

Electrochemical Impedance Spectra (Dunn, et al., 2003a).

0.0024 0.0026 0.0028 0.0030 0.0032 0.0034 0.00361/T (1/K)

-14

-12

-10

-8

-6

-4

Ln (C

orro

sion

rate

, mm

/yr)

35% MgCl2

0.028 M NaCl

Mill-annealed Alloy 22

Ea = 49.6 kJ mol-1

Ea = 46.3 kJ mol-1

Figure 3-7. Activation Energy for Alloy 22 CorrosionRates in 0.028 M NaCl and 35-Percent MgCl2

(7.5 M Choride) (Dunn, et al., 2003a)

3-9


10-6

10-5

10-4

10-3

Cor

rosi

on ra

te, m

m/y

r

0.028 M NaCl


Figure 3-8. Passive Corrosion Rate for Mill-AnnealedAlloy 22 in 0.028 M NaCl As a Function of

Temperature. Corrosion Rates Were Obtained fromElectrochemical Impedance Spectra. Multiple

Measurements Were Obtained at Each Temperaturefor a Minimum of 10 Days.

Figure 3-8. The tests were started at 95 °C [203 °F]. Multiple impedance measurements wereconducted at each exposure temperature for a minimum of 10 days. When the impedancespectra for a given temperature became independent of time, the temperature was reduced andthe process was repeated. In general, the lowest corrosion rates were obtained near the end ofthe exposure. At 25 °C [77 °F], the impedance of the passive film was higher, and theuncertainty in the impedance measurements was larger. Figure 3-9 shows the log of thecorrosion rate as a function of the inverse of temperature, which was used to determine anactivation energy of 41.8 kJ/mol [10.0 kcal/mol]. The slightly lower activation energy for the datashown in Figure 3-8 compared to the previously calculated activation energy of 46.3 KJ mol!1

[11.1 kcal mol!1] is consistent with the decrease in corrosion rate at 95 °C [203 °F] in Figure 3-8compared to the measurements shown in Figure 3-5.

Corrosion rates of Alloy 22 were measured at temperatures up to 175 °C [347 °F] in 4 M NaClusing an autoclave with a solid state tungsten/tungsten oxide reference electrode. Attempts toobtain impedance spectra using an internal pressure balanced Ag/AgCl reference electrodewere not successful. The corrosion rates, measured using electrochemical impedancespectroscopy, are shown in Figure 3-10. In general, the corrosion rates measured in thesetests are higher than previous results obtained in standard test cells. In addition, the corrosionrates measured at 125 and 175 °C [257 to 347 °F] appear to deviate significantly from the trendline. It is not clear if this deviation is related to possible changes in the reference electrode atelevated temperature. The activation energy calculated using these data was 33.6 kJ/mol[8.0 kcal/mol]. The lower activation energy for these tests is consistent with the higher corrosionrates measured at low temperatures compared with the corrosion rates shown in Figure 3-8. Although the data shown in Figure 3-10 are not in agreement with previous results, they suggestthat the passivity of Alloy 22 can be maintained at elevated temperatures in relativelyconcentrated chloride solutions.

3-10

0.0026 0.0028 0.0030 0.0032 0.00341/T (1/K)

-13

-12

-11

-10

-9

-8

-7

LN (C

orro

sion

rate

, mm

/yr)

0.028 M NaCl


Ea = 41.8 kJ mol-1

Figure 3-9. Activation Energy for the PassiveCorrosion Rate for Mill-Annealed Alloy 22 in 0.028 M

NaCl As a Function of Temperature. Corrosion RatesWere Obtained from Multiple Electrochemical

Impedance Spectra.

20 40 60 80 100 120 140 160 180Temperature, oC

10-5

10-4

10-3

10-2

10-1

Cor

rosi

on ra

te, m

m/y

r

4 M NaClMill-annealed Alloy 22

Figure 3-10. Passive Corrosion Rate for Mill-AnnealedAlloy 22 in 4 M NaCl As a Function of Temperature.

Corrosion Rates Were Obtained from ElectrochemicalImpedance Spectra. Test Was Conducted in an

Autoclave Using a Tungsten/Tungsten Oxide SolidState Reference Electrode.

3-11

3.3 Surface Analyses of Alloy 22

3.3.1 Short-Term Tests in 0.028 M NaCl Solution

Alloy 22 exhibits passive behavior at a potential of 100 mVSCE in 0.028 M NaCl solution at 95 °C[203 °F] as a result of the formation of a protective oxide film, whereas transpassive behavioroccurs at 600 mVSCE when the passive film breaks down (Dunn, et al., 2001). The wide-energysurvey spectrum of the electrochemically treated Alloy 22 specimen at 100 mVSCE indicated thepresence of nickel, chromium, molybdenum, oxygen, and carbon. In addition, iron and tungstenwere also observed in the surface film formed on the Alloy 22 specimen after anodicpotentiostatic treatment at 600 mVSCE. In both specimens, the high-carbon concentration on theas-treated surface was caused by atmospheric contamination during transfer of the samplesin air.

The concentration depth profiles for surface films formed on Alloy 22 at 100 mVSCE and600 mVSCE, as well as for the air-exposed Alloy 22 control specimen, are presented inFigure 3-11, where atomic concentrations of all elements detected are provided. The relationbetween sputtering time and equivalent depth is calculated using a sputtering rate of2.27 nm/min [8.94 × 10!5 mil/min] obtained from an SiO2 film under the same sputteringconditions. A significant change in film composition can be seen as the applied regionincreases from a passive region to a potential of transpassive dissolution.

For Alloy 22 treated in the passive region, three areas can be defined in the concentration depthprofiles shown in Figure 3-11(a): the outer contamination layer (I), the inner passive layer (II),and the base material (III). Area II represents the passive film, where a mixture of nickel andchromium oxides is present. The passive film was observed to have a high ratio of chromium tonickel at the outer side, and a decreased ratio at the inner side of the film. Nevertheless, theconcentration depth profiles show an increased chromium concentration, but low nickel andmolybdenum concentrations, in the passive film compared to the bulk alloy, as denoted inFigure 3-11. The thickness of the passive film was determined by the depth at which the nickelconcentration reached a constant value. From this criterion, the passive film formed on theelectrochemically treated Alloy 22 surface at 100 mVSCE is approximately 5.4 nm[2.12 × 10!4 mils]. The three-area structure was also evident for air grown film on the Alloy 22control specimen in Figure 3-11(c). Unlike the Alloy 22 specimen at 100 mVSCE, the ratio ofchromium to nickel remains low across the passive layer. The air-grown-oxide film rich in nickelis also different from the Cr2O3 scale formed after high-temperature dry-air oxidation, asdiscussed in Chapter 2.

Anodic treatment at 600 mVSCE under transpassive conditions results in a thick, chromium-richoxide surface layer, as shown in Figure 3-11(b). The thickness of the film is greater than23.0 nm [9.05 × 10!4 mils]. The concentration depth profiles show that the measuredconcentration of chromium to nickel in the surface layer is much higher than that in the bulkalloy. In addition, high iron concentrations were noted across the surface layer. Dunn, etal. (2001) analyzed the layer deposited on the Alloy 22 specimen after transpassive dissolution. The measured iron concentration in the deposit was found to be much higher than in the bulkAlloy 22 (21.3 versus 3.8 wt%). The morphology of the oxide film before and after depthprofiling analysis is shown in Figure 3-12, in which it is seen that elongated patchespredominate. The discontinuous surface layer is an indication of breakdown of the passive filmand would not provide corrosion protection. It is important that the transformation of Cr2O3 to

3-12

0

20

40

60

80

0 20 40 60 80 100 120 140 160 180Sputtering Time, seconds

Con

cent

ratio

n, a

t. %

0

20

40

60

800 22.7 45.4 68.1

Equivalent Depth, angstroms

OCCrMoNi

(a) Alloy 22, 0.028M NaCl, 100 mVSCE

I II IIINi (bulk)

Cr (bulk)

Mo (bulk)

0

20

40

60

80

0 60 120 180 240 300 360 420 480 540 600 660Sputtering Time, seconds

Con

cent

ratio

n, a

t. %

0

20

40

60

800 22.7 45.4 68.1 90.8 113.5 136.2 158.9 181.6 204.3 227 249.7


OCCrMoFeNi

(b) Alloy 22, 0.028M NaCl, 600 mVSCE

I II

Fe (bulk)

Ni (bulk)

Cr (bulk)

Mo (bulk)

Figure 3-11. Concentration Depth Profiles of Alloy 22Electrochemically Treated in Deaerated 0.028 M NaCl at 95 °C

[203 °F] with an Applied Potential of (a) 100 mVSCE and (b) 600 mVSCEand (c) Profiles of Alloy 22 Control Specimen Air Exposed at RoomTemperature. Equivalent Depth Is Relative to the Sputtering Rate of

SiO2 Under the Same Conditions.

3-13

0

20

40

60

80

0 10 20 30 40 50 60 70 80Sputtering Time, seconds

Con

cent

ratio

n, a

t. %

0

20

40

60

800 11.35 22.7


OCCrMoNi

(c) Alloy 22, ControlI II III

Ni (bulk)

Cr (bulk)

Mo (bulk)

Figure 3-11. Concentration Depth Profiles of Alloy 22Electrochemically Treated in Deaerated 0.028 M NaCl at 95 °C

[203 °F] with an Applied Potential of (a) 100 mVSCE and(b) 600 mVSCE and (c) Profiles of Alloy 22 Control Specimen AirExposed at Room Temperature. Equivalent Depth Is Relativeto the Sputtering Rate of SiO2 Under the Same Conditions.

(continued)

Cr(VI) species in solution is expected to lead to transpassive dissolution in Alloy 22, accordingto the Pourbaix diagram of the nickel-chromium-molybdenum-water system (Pensado, et al.,2002). Nonetheless, there was no evidence of Cr(VI) species within the oxide film.

The contributions of each of the species were extracted by deconvoluting the curve of theNi 2p3/2, Cr 2p3/2, and O 1s core-level signals. Figure 3-13 shows the high-resolution spectrumcurve fits of the Ni 2p3/2, Cr 2p3/2, and O 1s regions of Alloy 22 passivated at 100 mVSCE beforeand after sputtering for 36 and 171 seconds. Each of the core level signals was fitted by alinear combination of characteristic peaks of corresponding species. One characteristic peak inthe Ni 2p signal that is not identified is the result of multiple splitting of the Ni 2p3/2 core-level,whereas the one in the Cr 2p signal is from the Cr 2p1/2 core level. The Ni 2p3/2 core-level signalwas fitted with three peaks corresponding to the bivalent (NiO), trivalent (Ni2O3), and metallicspecies, and the chromium spectrum was fitted with a combination of a trivalent peak (Cr2O3),consisting of triplet peaks, and a metallic peak. In addition, the O 1s signal was fitted with twopeaks corresponding to hydroxide and oxide species. Therefore, separate contributions andconcentrations of each species could be calculated. The Mo 3d5/2 spectra, also fitted with a setof oxidation states of molybdenum corresponding to the Mo(IV), Mo(V), Mo(VI), and metallicspecies, are not shown here. Because of the thin passive film, the metallic speciescontributions would be attributed to the photoelectrons emitted from the base alloy. Thecalculated oxide film compositions for Alloy 22 at 100 mVSCE at selected sputtering times aregiven in Table 3-1. The Ni-oxide is the total of Ni2O3 and NiO species, Cr-oxide as Cr2O3, and

3-14

(a)

(b)

Figure 3-12. Scanning Electron Microscope Images of Alloy 22 ElectrochemicallyTreated in Deaerated 0.028 M NaCl at 95 °C [203 °F] with an Applied Potential of

600 mVSCE for (a) Before and (b) After Depth-Profiling Analysis

3-15

Figure 3-13. High-Resolution Spectrum Curve Fits of the Cr 2p3/2, Ni 2p3/2, and O 1sRegions of Alloy 22 Passivated in Deaerated 0.028 M NaCl at 95 °C [203 °F] with anApplied Potential of 100 mVSCE for the As-Received (Top Row) and After Sputtering

for 36 Seconds (Middle Row) and for 171 Seconds (Bottom Row)

3-16

Mo-oxide is the total of all molybdenum oxides. As seen in Table 3-1, the passive film iscomposed of a mixture of chromium, nickel, and molybdenum oxides, and the outer side of thefilm is rich in chromium oxides. The high-resolution O 1s spectra in Figure 3-13 indicate that thechromium oxide in the passive film was hydrated at the outer side of the film, consistent with theobservation in nickel-chromium-iron alloys (Marcus and Grimal, 1992).

Considering the extent of the three surface areas discussed previously, the three sets of curvefits in Figure 3-13 are examples within the outer contamination layer, the inner passive layer,and the base material of the Alloy 22 specimen after passivation at 100 mVSCE. As anticipatedfrom the characteristics of the as-received surface, the contributions of chromium and nickeloxides were found to predominate in the outer contamination layer. It also is significant thatcontributions of chromium and nickel oxides after sputtering for 171 seconds persisted down tothe base material. The total oxide concentration, however, is low because the oxygenconcentration in the base material region was measured to be 8 atomic percent. The mostprofound feature of the species contributions on Alloy 22 is the dominance of Cr2O3 in thepassive film. The stable Cr2O3 passive film provides excellent corrosion resistance to Alloy 22,particularly the general corrosion resistance.

Figure 3-14 shows the species concentration depth profiles obtained for Alloy 22 at 100 mVSCEand for the control specimen, where the atomic concentrations of metallic and oxide states areprovided. A distinct Cr2O3-rich passive film was evident on the electrochemically treatedAlloy 22 surface in the passive region at 100 mVSCE. In contrast to the oxide film observed at100 mVSCE, metallic nickel species were found to predominate in the air grown surface layer onthe Alloy 22 control specimen.

Table 3-1. Calculated Oxide Film Composition of Alloy 22 Passivated at 100 mVSCE inDeaerated 0.028 M NaCl at 95 °C [203 °F]

SputteringTime (sec)

EquivalentDepth (D)

Oxide Film Composition (%)

Cr-Oxide Ni-Oxide Mo-Oxide

0 0 79 14 7

9 3.4 76 17 7

18 6.8 76 16 8

36 13.6 79 13 8

54 20.4 84 9 7

72 27.2 82 12 6

90 34.1 65 29 6

108 40.9 56 37 7

126 47.7 43 49 8

171 64.7 30 63 7

3-17

0

10

20

30

40

50

60

70

0 20 40 60 80 100 120 140 160 180Sputtering Time, seconds

Con

cent

ratio

n, a

t. %

0

10

20

30

40

50

60

700 22.7 45.4 68.1


Cr(ox)Cr(met)Ni(ox)Ni(met)O-2OH-

(a) Alloy 22, 0.028M NaCl, 100 mVSCE Ni (bulk)

Cr (bulk)

0

10

20

30

40

50

60

70

0 10 20 30 40 50 60Sputtering Time, seconds

Con

cent

ratio

n, a

t. %

0

10

20

30

40

50

60

700 11.35 22.7


Cr(ox)Cr(met)Ni(ox)Ni(met)O-2OH-

(b) Alloy 22, ControlNi (bulk)

Cr (bulk)

Figure 3-14. Species Concentration Depth Profiles of (a) Alloy 22Electrochemically Treated in Deaerated 0.028 M NaCl at 95 °C

[203 °F] with an Applied Potential of 100 mVSCE and (b) Alloy 22Control Specimen Air Exposed at Room Temperature. Equivalent

Depth Is Relative to the Sputtering Rate of SiO2 Under theSame Conditions.

3-18

Analyses of the passive films formed on Alloy 22 have shown that the composition and structureof the passive films depend largely on the pH of the solution and the applied potential (BechtelSAIC Company, LLC, 2004; Lloyd, et al., 2003). Lloyd, et al. (2003) analyzed the passive filmson the Alloy 22 specimens potentiostatically treated in a solution of 1 M NaCl + 0.1 M H2SO4(pH 1) using x-ray photoelectron spectroscopy and time-of-flight secondary ion massspectroscopy. They reported that the passive oxide films are composed of a molybdenum-richouter layer and a chromium- and nickel-rich inner layer. The oxide film thickness increasedslightly with increasing applied potential from 2.0 nm [7.87 × 10!5 mils] at 200 mV to 2.8 nm[1.10 × 10!4 mils] at 700 mV with respect to the Ag/AgCl electrode. The U.S. Department ofEnergy (DOE) conducted a series of investigations to characterize the oxide films formed onAlloy 22 as a function of pH and applied potential (Bechtel SAIC Company, LLC, 2004). A thinoxide film ranging from 2 to 5 nm [7.87 × 10!5 to 1.97 × 10-4 mils] was formed in the passiveregion at pH levels of 3 and 8. The oxide films are composed of chromium oxides, primarilyCr2O3. At or near transpassive conditions at pH 8, the surface films were observed to have twodistinct layers; a conformal, chromium oxide-rich inner layer of about 4 nm [1.57 × 10!4 mils] anda porous, nickel oxide-rich outer layer in thickness of about 30 to 40 nm [1.18 to1.57 × 10!3 mils].

The observation of the molybdenum-rich outer layer of the passive film that formed on Alloy 22in 1 M NaCl + 0.1 M H2SO4 solution (pH 1) by Lloyd, et al. (2003) could be the result ofpreferential dissolution of chromium under the aggressive, acidic conditions. Evidence from thepresent study on the observed composition and structure of the passive films formed onAlloy 22 in 0.028 M NaCl solution (pH 5.5) under both passive and transpassive conditionsagrees with the results reported by DOE (Bechtel SAIC Company, LLC, 2004), except for theduplex oxide structure formed under transpassive conditions. Transpassive conditions,however, are not expected to occur under the expected repository environments.

3.3.2 Long-Term Tests in 4 M NaCl Simulated Groundwater

Alloy 22 specimens were also characterized using x-ray photoelectron spectroscopy afterlong-term (about 2 years) immersion exposure in simulated groundwater of pH 7.5 underopen-circuit conditions and at 250 mVSCE, both in the passive region. Test specimens fromsections of two crevice specimens, outside and inside the crevice sites, were analyzed. Figure 3-15 shows the concentration depth profiles of both passivated Alloy 22 specimensoutside the crevice sites. The equivalent depth is calculated using a sputtering rate of1.49 nm/min [5.87 × 10!5 mil/min] for SiO2 film. It is apparent that the surface films formed onAlloy 22 had two surface areas: an outer deposition layer that contained silica as a result ofglass test cell dissolution (Area I), and an inner layer rich in all elements in the alloy (Area II). The species concentration depth profiles obtained from the high-resolution curve fits of theNi 2p3/2, Cr 2p3/2, and O 1s regions with time are shown in Figure 3-16. In both cases, the innerlayers were determined to be rich in the metallic species of nickel and chromium with lowconcentrations of the oxide species. Additionally, a high contribution, corresponding to waterspecies, was measured in the middle of the outer layer on both specimens, suggesting that thesilica deposit is porous. DOE also observed a coating of porous silica scale on Alloy 22treated at 250 mV in basic saturated water (Bechtel SAIC Company, LLC, 2003a) due toglassware dissolution (Bechtel SAIC Company, LLC, 2004). The film underlying the silica,however, was observed to be predominantly the metallic species of nickel with some chromiumoxide. The loss of the predominant chromium oxide underlayer in the presence of silica deposit

3-19

0

20

40

60

80

0 2000 4000 6000 8000 10000Sputtering Time, seconds

Con

cent

ratio

n, a

t. %

0

20

40

60

800 49.6 99.2 148.8 198.4 248

Equivalent Depth, nm

OCSiCrMoNi

(a) Alloy 22, 4M NaCl, Open-CircuitI II

Ni (bulk)

Cr (bulk)

Mo (bulk)

0

20

40

60

80

0 2000 4000 6000 8000 10000 12000 14000Sputtering Time, seconds

Con

cent

ratio

n, a

t. %

0

20

40

60

800 49.6 99.2 148.8 198.4 248 297.6 347.2


OCSiCrMoNi

(b) Alloy 22, 4M NaCl, 250 mVSCE

I IINi (bulk)

Cr (bulk)

Mo (bulk)

Figure 3-15. Concentration Depth Profiles of Alloy 22Electrochemically Treated in Deaerated 4 M NaCl Multi-Ionic Solution

at 95 °C [203 °F] (a) Under Open-Circuit Condition and (b) with anApplied Potential of 250 mVSCE. Equivalent Depth Is Relative to the

Sputtering Rate of SiO2 Under the Same Conditions.

3-20

0

10

20

30

40

50

60

70


Con

cent

ratio

n, a

t. %

0

10

20

30

40

50

60

7074.4 99.2 124 148.8 173.6 198.4 223.2 248


Cr(ox)Cr(met)Ni(ox)Ni(met)SiO2H2OO-2OH-

(a) Alloy 22, 4M NaCl, Open-CircuitNi (bulk)

Cr (bulk)

0

10

20

30

40

50

60

70


Con

cent

ratio

n, a

t. %

0

10

20

30

40

50

60

70173.6 198.4 223.2 248 272.8 297.6 322.4 347.2


Cr(ox)Cr(met)Ni(ox)Ni(met)SiO2H2OO-2OH-

(b) Alloy 22, 4M NaCl, 250 mVSCE Ni (bulk)

Cr (bulk)

Figure 3-16. Species Concentration Depth Profiles of Alloy 22Electrochemically Treated in Deaerated 4 M NaCl Multi-Ionic Solution

at 95 °C [203 °F] (a) Under Open-Circuit Condition and (b) with anApplied Potential of 250 mVSCE. Equivalent Depth Is Relative to the

Sputtering Rate of SiO2 Under the Same Conditions.

3-21

is a concern. It is necessary to determine the stability of the passive oxide film in a range ofrelevant silica-containing environments.

Passive films formed on crevice sites were characterized using x-ray photoelectronspectroscopy and are shown in Figures 3-17 and 3-18. As was observed outside the crevicesites, the crevice sites had an outer deposition layer that contained silica as a result of glasstest cell dissolution. Figure 3-18 shows that a distinct chromium oxide (in Cr2O3) passive film,different from that formed outside the crevice sites, was identified on the crevice sites. Theseresults on passive film characterization of Alloy 22 using x-ray photoelectron spectroscopysuggest that the oxide films formed on Alloy 22 in the passive region consist primarily of theCr2O3 species, and these Cr2O3 surface films are responsible for the passivity of the alloy.

3-22

0

20

40

60

80


Con

cent

ratio

n, a

t. %

0

20

40

60

800 16.5 33 49.5 66 82.5


OCSiCrMoNi

(b) Alloy 22 Crevice Area 4M NaCl, 250 mVSCE

Ni (bulk)

Cr (bulk)

Mo (bulk)

Figure 3-17. Concentration Depth Profiles in the Crevice Area ofAlloy 22 Electrochemically Treated in Deaerated 4 M NaCl Multi-Ionic

Solution at 95 °C [203 °F] (a) Under Open-Circuit Condition and(b) with an Applied Potential of 250 mVSCE. Equivalent Depth Is

Relative to the Sputtering Rate of SiO2 Under the Same Conditions.

0

20

40

60

80


Con

cent

ratio

n, a

t. %

0

20

40

60

800 16.5 33 49.5 66 82.5


OCSiCrMoNi

(a) Alloy 22 Crevice Area 4M NaCl, Open-Circuit

Ni (bulk)

Cr (bulk)

Mo (bulk)

3-23

0

10

20

30

40

50

60

70


Con

cent

ratio

n, a

t. %

0

10

20

30

40

50

60

700 8.25 16.5 24.75 33 41.25 49.5


Cr(ox)Cr(met)Ni(ox)Ni(met)SiO2O-2OH-

(a) Alloy 22 Crevice Area4M NaCl, Open-Circuit

Ni (bulk)

Cr (bulk)

0

10

20

30

40

50

60

70


Con

cent

ratio

n, a

t. %

0

10

20

30

40

50

60

700 8.25 16.5 24.75 33 41.25 49.5


Cr(ox)Cr(met)Ni(ox)Ni(met)SiO2O-2OH-

(b) Alloy 22 Crevice Area4M NaCl, 250 mVSCE

Ni (bulk)

Cr (bulk)

Figure 3-18. Species Concentration Depth Profiles in the CreviceArea of Alloy 22 Electrochemically Treated in Deaerated 4 M NaCl

Multi-Ionic Solution at 95 °C [203 °F] (a) Under Open-CircuitCondition and (b) with an Applied Potential of 250 mVSCE. Equivalent

Depth Is Relative to the Sputtering Rate of SiO2 Under theSame Conditions.

4-1

4 LOCALIZED CORROSION

The abstraction of localized corrosion susceptibility of the waste package outer container isbased on a critical potential model. Key inputs to the localized corrosion abstraction are thecorrosion potential (Ecorr) and the critical potential for localized corrosion initiation (Ecrit). Basedon results obtained for passive chromium-containing alloys, the repassivation potential forcrevice corrosion (Ercrev) was selected as the Ecrit (Dunn, et al., 2000; Cragnolino, et al., 2004a). Results of electrochemical tests using mill-annealed, thermally treated, and welded specimensto determine the effect of fabrication processes on the long-term performance of wastepackages have been reported by Dunn, et al. (2003a). More recently, Dunn, et al. (2004) havereported the effects of environmental conditions, including the effects of oxyanions as inhibitorsfor localized corrosion. Information presented in this chapter summarizes previous reports andupdates the results of tests conducted at the Center for Nuclear Waste Regulatory Analyses.

4.1 Test Methods

The chemical compositions of the alloys used in this study are provided in Table 2-1. Alloys 825 and 625 were tested in mill-annealed condition. Although these alloys are no longerconsidered candidate container materials, results of tests with these alloys are reported andcompared with the results for Alloy 22 which was tested in both mill-annealed and thermallyaged {870 °C [1,598 °F] for 5 minutes} conditions. Short-term thermal aging was used toreplicate the enhanced localized corrosion susceptibility observed in the heat affected zones ofwelded Alloy 22.

The Ecorr for Alloy 22 was measured in standard test cells with a smooth cylindrical specimen(Figure 3-1) as the working electrode, a platinum counter electrode, and a saturated calomelelectrode as a reference that was maintained at room temperature. Temperature, chlorideconcentration, pH, and temperature were varied to determine the effects of environmentalvariations. At the completion of each test, the specimens were rinsed in deionized water, dried,and examined with an optical microscope.

Potentiodynamic tests were used to determine the Ercrev for Alloy 22. Limited tests were alsoconducted for Alloys 625 and 825. The tests were conducted in standard test cells identical tothose used in the Ecorr measurements, with a crevice specimen as the working electrode(Figure 3-1). Potentiodynamic scans were initiated at !200 mVSCE, and the potential of thespecimens was increased to a preset value in the range of 400 to 700 mVSCE at a rate of0.1 mV s!1. On reaching the preset maximum potential, the specimen was heldpotentiostatically to allow localized corrosion propagation. After a period of 5 to 8 hours, thepotential was decreased at a rate of 0.0167 mV s!1 to a potential of !500 mVSCE, where the testwas terminated. The crevice corrosion repassivation potential Ercrev is defined as the potentialwhere the current density remains below 2 × 10!6 A cm!2 [1.9 × 10!3 A ft!2] (Dunn, et al., 2003b). Test solutions used for the potentiodynamic tests included chloride only solutions to obtainbaseline values for the Ercrev. Tests were also conducted in solutions containing predeterminedmolar concentration ratios of oxyanions to chloride. The oxyanions included nitrate, sulfate,bicarbonate, carbonate, and fluoride.

Potentiostatic and open circuit tests were conducted to verify Ercrev values measured in thepotentiodynamic tests. Potentiostatic tests were conducted for a minimum of 30 days under astatic applied potential selected based on results of potentiodynamic tests. Open-circuit crevice

4-2

Figure 4-1. Glass Cell Used for Alloy 22 Localized Corrosion Propagation RateMeasurements. The Polytetrafluoroethylene (PTFE) Bolt Is Pressed Against Alloy 22Cylindrical Specimen to Form an Artificial Crevice. A Glass Wall Is Fitted Between

Two PTFE Lids to Form a Cell.

corrosion tests were conducted using an Alloy 22 crevice specimen galvanically coupled to alarge Alloy 22 plate through a potentiostat functioning as a zero resistance ammeter. Thepurpose of this test was to determine if localized corrosion could be initiated under open circuitconditions. The galvanic couple used in these tests simulated the situation typical for crevicecorrosion of a passive alloy. Crevice corrosion can occur under the occluded region created bythe crevice, and the crevice free surface acts as a site for the cathodic reaction. The redoxconditions were controlled by adding cupric chloride (CuCl2) to the 5 M sodium chloride (NaCl)test solution. The potential of the galvanic couple and the galvanic current density weremonitored throughout the tests. The potential and current transients were analyzed todetermine the initiation points and propagation time for localized corrosion. In addition, the testspecimens were examined after the completion of the tests for signs of localized corrosion.

The localized corrosion propagation rate of Alloy 22 was measured using a single crevice withdefined geometry (“pencil electrodes”) shown in Figure 4-1. A PTFE bolt with a cone-shapedfront end was pressed against an Alloy 22 cylindrical specimen measuring 6.3 mm [0.25 in] indiameter and 48 mm [1.89 in] to form an artificial crevice with a contact surface area of0.106 cm2 [0.0164 in2]. In some tests the PTFE bolt was modified to fit a spring whichcompressed the PTFE bolt against the Alloy 22 specimen. Tests were conducted in 300 mL[0.080 gal] test cells made of glass and PTFE at 110 °C [230 °F] in 4 M MgCl2 [8 M Cl!]solutions. Test cells were fitted with a water-cooled condenser to minimize solution loss atelevated temperatures. A saturated calomel electrode was used as a reference electrode in all

4-3

I Icorr aex= / ε (4-1)

h FAmax / ( )= Iaex EW ρ (4-2)

experiments. The reference electrode was connected to the solution through a water-cooledLuggin probe with a porous silica tip to maintain the reference electrode at room temperature. The Alloy 22 specimen was galvanically coupled to a platinum flag. The galvanic couplingcurrent and the potential were monitored using a computer controlled potentiostat. Thelocalized corrosion propagation rate was determined by performing the crevice corrosion testsfor various durations after the initiation of localized corrosion and subsequently examining thelocalized corrosion area and measuring the maximum penetration depths using anoptical microscope.

Localized corrosion propagation rates were also measured using a coupled multielectrode arraysensor (Yang, et al., 2002a,b; Yang and Sridhar, 2003a,b). The advantage of the multielectrodearray sensor is the ability to measure dissolution current associated with localized corrosionwhich is isolated to small areas. Traditional electrochemical methods that measure theresponse over a large surface area, such as the widely used electrical resistance and linearpolarization methods (Agarwala and Ahmad, 2000; Sridhar, et al., 2000; Mansfeld, 1987, 1976),are not sufficiently sensitive to determine propagation rates.

The coupled multielectrode array sensor uses miniature electrodes coupled to a common jointthrough independent small resistors (Figure 4-2) or zero-resistance ammeters. By properlyselecting the resistor values, the voltage drops across the resistors are small but sufficient formeasurement of the current flowing through each electrode. Under conditions where the sensorelements are susceptible to localized corrosion, the anodic current flows into the more corrodingelectrodes, and the cathodic current flows out of the less or noncorroding electrodes. Thecurrent from the most anodic electrode is the current from the most active anodic site or thedeepest pit for the case of pitting corrosion. If the small surface area of the electrode iscorroded uniformly, the penetration rate may be calculated using Eq. (4-1)

where

Icorr — total anodic corrosion current on the most corroding electrode (A)Ia

ex — the measured most anodic current (A)g (0 < g # 1) — current distribution factor that represents the fraction of electrons resulting

from corrosion that flows through the external circuit (Yang, et al., 2005,2002a)

The value of g is less than unity when the most corroding electrode is not significantly morecorroded than the other electrodes and still has cathodic sites that can receive the electronsfrom an anodic site on the same electrode. If the number of the electrodes in a multielectrodearray sensor probe is large, and the surface area of each electrode is sufficiently small or closeto the size of one single pit, the most corroding electrode is very likely to be completelycorroded and there is no cathodic site available on it. In this case, g = 1, and the penetrationrate of localized corrosion may be calculated using Eq. (4-2) (Yang, et al., 2002a):

4-4

EW i i i= ∑1 / ( / )m z W (4-3)

where

hmax — estimated maximum penetration rate (cm/s)F — Faraday constant (96485 C/mol)A — surface area of the electrode (cm2)D — density of the alloy or electrode (g/cm3)EW — equivalent weight

and Eq. (4-3).

where

mi, zi and Wi — mass fraction of each alloying element in the alloy, oxidation state of thealloying element in solution, and atomic weight of alloying element, i, inthe alloy, respectively

As the current from the most corroding electrode depends on only one of the many measuredcurrents from the electrodes on a probe and therefore has a greater uncertainty, statisticalparameters may be used to derive the most corroding current or the equivalent most corrodingcurrent based on the currents from all or a group of the electrodes. Therefore, three times thestandard deviation of the currents was often used as a better choice compared to the value ofthe most anodic current (Dorsey, et al., 2004; Yang, et al., 2004).

Epoxy

Joinedtogether

To Multi-Channel Voltmeter

R

RR

R

R

Electrode (insulated)

Sealingtube

(a) (b)

Figure 4-2. Schematic Diagram (a) and Different Configurations(b) of Multielectrode Array Sensor Probes (Yang and Sridhar, 2003a).

Reproduced with Permission from NACE International.

4-5

4.2 Corrosion Potential Measurements

The Ecorr of Alloy 22 was measured using replicated specimens exposed to a range of conditionsto determine the significance of environmental variations. Figure 4-3 shows an example of theevolution of the Ecorr in a solution containing 4 M NaCl adjusted to pH 8 with the addition ofNa2CO3. The temperature was varied from 95 to 25 °C [203 to 77 °F]. Similar tests wereconducted in solutions with predetermined compositions to systematically vary solution pH. Thedata shown in Figure 4-3 were reduced to an average and a standard deviation foreach specimen.

The corrosion potentials measured as a function of pH in chloride solutions at 95 °C [203 °F] areshown in Figure 4-4. The points are average values and the error bars indicate one standarddeviation of the measured corrosion potential for each specimen. In alkaline solutions, thecorrosion potential does not appear to be a function of solution pH or chloride concentration. Inaddition, with the exception of the larger variation in the Ecorr observed in 0.028 M Cl at pH 8.2,the Ecorr was not dependent on the surface condition of the Alloy 22 specimens. These solutionscontained chloride, with added sodium bicarbonate (NaHCO3) and sodium carbonate (Na2CO3)to alter solution pH. Carbonate and bicarbonate salts are common in groundwater and do notact as oxidizing or reducing agents. For polished specimens in alkaline 0.028 M NaCl solution,the Ecorr was as low as !340 mVSCE at the start of the test and increased and stabilized at valuesin the range of !200 to 0 mVSCE. Oxidized specimens demonstrated much greater variability inEcorr compared to the polished specimens. Initial values of the Ecorr were as high as 65 mVSCEand, unlike the polished specimens, decreased with time. Initial variations and a generaldecrease in the oxide film thickness may contribute to the specimen-to-specimen variation andthe general decrease in the Ecorr. The Ecorr in acidic solutions was more than 300 mV greater

0 20 40 60 80 100Time, days

-0.2

-0.1

0

0.1

0.2

Pot

entia

l, V

SC

E

20

40

60

80

100

Tem

pera

ture

, o C

Temperature

Platinum

Alloy 22(2 specimens)

Figure 4-3. Alloy 22 Corrosion Potential Measured atTemperatures from 25 to 95 °C [77 to 203 °F]. Test

Performed Under Open-Circuit Conditions.

4-6

than the Ecorr in alkaline solutions, and was also independent of the chloride concentration. Under acidic conditions, the Ecorr of the oxidized specimens increased from initial values in therange of !40 to 60 mVSCE and stabilized near 250 mVSCE. The specimen-to-specimen variationin acidic solutions was typically less than 50 mV. Similar values of Ecorr were obtained forpolished specimens in acidic solutions with 4 M NaCl.

The effect of temperature on the Ecorr is shown in Figure 4-5. Polished Alloy 22 specimens wereexposed to 4 M NaCl at temperatures of 25 to 95 °C [77 to 203 °F]. The pH of the test solutionwas varied from 8 to 12 by the addition NaHCO3 and Na2CO3. For all test conditions, the Ecorrdecreased with increasing temperature with the Ecorr values at 25 °C [77 °F] approximately 150to 200 mV greater than the values at 95 °C [203 °F]. Additional tests are being conducted todetermine the effect of temperature in neutral and acidic solutions.

4.3 Repassivation Potential Measurements

Dunn, et al. (2004, 2003a,b) and Cragnolino, et al. (2004a) contain the results of Ercrevmeasurements for Alloy 22. Fabrication processes such as welding, postweld heat treatments,and short-term exposures to temperatures where topologically close packed phases are stable,can reduce the the Ercrev. Figure 4-6 shows the results of repassivation tests with mill-annealed,thermally aged {870 °C [1,598 °F]}, welded, and welded plus solution annealed Alloy 22. Datafrom welded specimens was obtained using both 12.7- and 38.1-mm [0.5- and 1.5-in]-thick gastungsten arc welded plates (Dunn, et al., 2004, 2003a). In general, as-welded materials exhibit

2 4 6 8 10 12Solution pH

-300

-200

-100

0

100

200

300

400

500

600

Cor

rosi

on p

oten

tial,

mV

SCE 0.028 M NaCl Polished

0.028 M NaCl Oxidized

4.0 M NaCl Polished

Alloy 22 corrosion potential Air-saturated solutions, 95 oC [203 oF]

0.014 M MgCl2 Polished0.014 M MgCl2 Oxidized

Figure 4-4. Corrosion Potentials of Alloy 22 As a Function ofSolution pH. Polished Specimens Were Polished Prior to

Immersion. Oxidized Specimens Were Thermally Oxidized at200 °C [392 °F] for 30 Days. Test Performed Under

Open-Circuit Conditions.

4-7

E A Arcrev T T01 2( ) ( )= + (4-6)

greater variability than thermally aged materials. With the exception of the thermally agedmaterial, the Ercrev values of the as-welded and welded plus solution annealed materials aresimilar at chloride concentrations above 0.5 M. The welded-plus-solution annealed material hasa lower Ercrev value in chloride concentrations less than 0.5 M. The thermally aged material hadthe lowest Ercrev value in, however, solutions with at least 0.01 M Cl!, and the value was similarto that of the welded plus solution annealed material at lower chloride concentrations.

Regression equations for Alloy 22 in the mill-annealed condition have been reported based onthe results of Ercrev measurements in chloride solutions (Dunn, et al., 2004, 2003a; Cragnolino,et al., 2004a). Similarly, regression equations for thermally aged Alloy 22 have also beenreported. The regression equation for Ercrev is shown in Eq. (4-5)

E E Brcrev rcrev T T Cl= + −0 ( ) ( )log[ ] (4-5)

where

andB B B( ) ( )T T= +1 2 (4-7)


-600

-400

-200

0

200

Cor

rosi

on p

oten

tial,

mV

SCE

pH 8pH 10pH 12

Alloy 22 Air-saturated 4M NaCl

Figure 4-5. Corrosion Potential of Alloy 22 As a Function ofTemperature. Solution pH Adjusted with the Addition ofEither Sodium Carbonate or Sodium Bicarbonate. Test

Performed Under Open-Circuit Conditions.

4-8

10-4 10-3 10-2 10-1 100 101 102

Chloride concentration, molar

-400

-200

0

200

400

600R

epas

siva

tion

pote

ntia

l, m

VS

CE

Mill Annealed

Crevice Corrosion

No CreviceCorrosion

95 oC [203 oF]

Time = 5 min/870 oC [1,598 oF]

As-weldedWelded + Solution Annealed

Time up to 4 hr/870 oC [1,598 oF]

Figure 4-6. Crevice Corrosion RepassivationPotentials for Mill-Annealed, Thermally Aged,

As-Welded and Welded-Plus-Solution AnnealedAlloy 22 in Chloride Solutions at 95 °C [203 °F]

4-9

Table 4-1. Parameters for the Localized Corrosion Regression Equations forMill-Annealed and Thermally Aged Alloy 22

MetallurgicalCondition

T (°C)[°F]

[Cl!]crit(M)

A1(mVSCE)

A2(mV/°C)

B1(mV)

B2(mV/°C)

Mill-annealed(as-received)

80 to 125 °C[176 to 257 °F]

0.5 1,300 !13.1 !362.7 2.3

Thermally aged 5 minutes at870 °C [1,598 °F]water quench

60 to 95 °C[140 to 203 °F]

0.25 to0.01

800 !10.0 !584.2 3.7

Equation parameters for the mill-annealed and thermally aged Alloy 22 are shown in Table 4-1and are considered valid when the inhibitor to chloride ratio is insufficient to affect the Ercrev. From Figure 4-6, the results for the thermally aged material are a suitable to bound the effectsof fabrication processes on the Ercrev of the Alloy 22 waste package outer container. Theregression equations for the repassivation potential do not consider the beneficial effects ofinhibiting species. Several studies have shown that nitrate, sulfate, carbonate, and bicarbonatecan inhibit localized corrosion of stainless steels and nickel-chromium-molybdenum alloys(Dunn, et al., 2004; Szklarska-Smialowska, 1986; Kehler, et al., 2001; Jallerat, et al., 1984). Additional tests were performed to evaluate the inhibiting effects of anionic species which arelikely to be present in groundwater or solutions produced by the deliquescence of dust. Theeffect of nitrate on the localized corrosion susceptibility of mill-annealed and thermally agedAlloy 22 in chloride solutions is shown in Figure 4-7. Tests were performed in magnesiumchloride solutions to evaluate the inhibiting effects of nitrate in concentrated chloride solutions attemperatures above 100 °C [212 °F]. In 4 M MgCl2 at 110 °C [230 °F], localized corrosion wasobserved in solutions with nitrate to chloride molar concentration ratios up to 0.1. At highernitrate to chloride concentration ratios, no localized corrosion was observed. The nitrate tochloride molar concentration ratio necessary to inhibit localized corrosion at 80 °C [176 °F] wasslightly lower. Similar results were obtained for thermally aged Alloy 22, however, higher nitrateto chloride ratios were necessary to inhibit localized corrosion. At 110 °C [230 °F], localizedcorrosion was inhibited when the nitrate to chloride molar concentration ratio was 0.3. Intergranular attack in the crevice regions was observed in solutions with low nitrate to chloridemolar concentration ratios. At 80 °C [176 °F], a lower nitrate to chloride ratio (0.1) is necessaryto inhibit localized corrosion.

Inhibiting effects of other anionic species were evaluated in 0.5 M NaCl using thermally agedAlloy 22 specimens. Test conditions were selected for the more susceptible thermally agedmaterial a range of anion to chloride concentration ratios. Figure 4-8 shows the results of testsconducted in 0.5 M NaCl with additions of either sulfate or fluoride at 95 °C [203 °F]. Localizedcorrosion was observed when the sulfate to chloride molar concentration ratio was 0.5;however, the repassivation potential was high compared with the results in pure 0.5 M NaCl. Athigher and lower concentration ratios, localized corrosion was inhibited, but the specimens hada heavy, dark gold colored film and showed evidence of transpassive dissolution. Results fortests conducted in fluoride and chloride solutions; do not show the same trend observed withthe other anions. For test conditions where localized corrosion was not observed, thespecimens exhibited significant uniform corrosion. With the exception of the results obtained for

4-10

0 0.1 0.2 0.3 0.4Nitrate to Chloride Molar Con centration Ratio

-400

-200

0

200

400

600

Rep

assiv

atio

n P

ote

ntia

l, m

VSC

E

No Crevice Corrosion

Crevice Corrosion

Alloy 22, 4 M MgCl2

MA 110 oCMA 80 oC

TA 80 oCTA 110 oC

Figure 4-7. Crevice Corrosion RepassivationPotentials for Mill-Annealed and Thermally Aged

Alloy 22

0 0.1 0.2 0.3 0.4Nitrate to Chloride Molar Con centration Ratio

-400

-200

0

200

400

600

Rep

assiv

atio

n P

ote

ntia

l, m

VSC

E


Crevice Corrosion

Alloy 22, 4 M MgCl2

MA 110 oCMA 80 oC

TA 80 oCTA 110 oC

Figure 4-7. Crevice Corrosion RepassivationPotentials for Mill-Annealed and Thermally Aged

Alloy 22

0 0.2 0.4 0.6 0.8 1Anion to Chloride Molar Concentrat ion Ratio

-200

-100

0

100

200

300

400

500

Rep

assi

vatio

n P

oten

tial,

mV

SC

E

No Crevice CorrosionCrevice Corrosion

Thermally aged Alloy 22, 0.5 M NaCl 95 oC870 oC for 5 min/water quench

SO42-F-

Figure 4-8. Crevice Corrosion RepassivationPotential Measurements Showing the Effect of

Fluoride and Sulfate on the Localized CorrosionSusceptibility of Alloy 22

4-11

fluoride to chloride molar concentration ratios of 0.3 and 0.6, the repassivation potentialdecreased slightly with increasing additions of fluoride.

Inhibition of localized corrosion was observed when the molar concentration ratio of carbonateto chloride was greater than 0.05 (Figure 4-9). A higher concentration ratio of 0.2 was requiredto inhibit localized corrosion with bicarbonate. When the test solutions contained sufficientcarbonate or bicarbonate concentrations to inhibit localized corrosion, no intergranular attackwas observed, but light etching of specimens, consistent with the onset of transpassivedissolution, was observed.

Results of speciation calculations for the bicarbonate- and carbonate-containing test solutionsare shown in Figures 4-10 and 4-11. These calculations indicate that the carbonateconcentration is much greater than the bicarbonate concentration when sodium carbonate isadded to sodium chloride. Conversely, the bicarbonate concentration is much greater than thecarbonate concentration when sodium bicarbonate is added to the sodium chloride solution. Asmall amount of sodium carbonate complex is calculated to be stable for both solutions. Theresults of the speciation calculations indicate that both carbonate and bicarbonate can beeffective inhibitors when present in sufficient concentrations relative to chloride.

The evolution of the chemistry of water contacting the waste package will affect the localizedcorrosion susceptibility of Alloy 22. The maximum anion concentrations as a function ofchloride concentration at 95 °C [203 °F] are shown in Figure 4-12. To inhibit localized corrosion

0 0.2 0.4 0.6 0.8 1Anion to Chloride Molar Concentrat ion Ratio

-200

0

200

400

600

Rep

assi

vatio

n P

oten

tial,

mV S

CE


Crevice Corrosion

Thermally aged Alloy 22870 oC for 5 min/water quench

0.5 M NaCl 95 oCHCO3

-CO32-

Figure 4-9. Crevice Corrosion RepassivationPotential Measurements Showing the Effect ofCarbonate and Bicarbonate on the Localized

Corrosion Susceptibility of Alloy 22

4-12

0.01 0.1 1Concentration of NaHCO3, molar

0

0.1

0.2

0.3

Spe

cies

con

cent

ratio

n, m

olar

7

8

9

10

11

pH

CO32-

HCO3-

NaCO3-

pH0.5 M NaCl95 oC

Figure 4-10. Results of Speciation Calculations ofSolutions Containing Sodium Bicarbonate and

Sodium Chloride at 95 °C [203 °C]

0.01 0.1 1Concentration of Na2CO3, molar

0

0.1

0.2

0.3

Spec

ies

conc

entra

tion,

mol

ar

7

8

9

10

11pH

CO32-

HCO3-

NaCO3-

pH0.5 M NaCl95 oC

Figure 4-11. Results of Speciation Calculations forSolutions Containing Sodium Carbonate and Sodium

Chloride at 95 °C [203 °F]

4-13

in NaCl solutions, the anion to chloride concentration ratio must be at least 0.1 for carbonateand nitrate and 0.2 for bicarbonate. These ratios are consistent with previous investigationsconducted with several stainless steels and nickel base alloys (Jallerat, et al., 1984;Szklarska-Smialowska,1986). From the data shown in Figure 4-12, it is apparent thatcarbonate, bicarbonate, and sulfate are sufficiently soluble in solutions containing up to 1 M Cl!. At higher chloride concentrations, the reduced solubility of sulfate and bicarbonate, combinedwith the higher concentration ratios required to be effective inhibitors, may prevent sufficientconcentrations of these anions to inhibit localized corrosion. Nitrate is not shown in Figure 4-12;however, nitrate salts are known to have high solubility.

4.4 Crevice Corrosion Initiation Tests

Potentiostatic tests were conducted to evaluate the use of the Ercrev as a parameter for thecritical potential for the initiation of localized corrosion. To be a valid parameter, the initiation oflocalized corrosion should not occur at potentials below the Ercrev (considering the uncertainty inthe measurement). Results of the potentiostatic tests for Alloys 825, 625, and 22 in 5.5 M NaClat 100 °C [212 °F] are shown in Figure 4-13 as a function of the potential difference ()E)between the applied potential (Eapplied) and the Ercrev. A large positive value of )E indicated thepotentiostatic test was conducted at applied potential well above the Ercrev. For reference,values of the crevice corrosion initiation (Ecrev) and Ercrev are shown for the alloys. The crevicecorrosion initiation time is a strong function of potential. For all materials, the crevice corrosioninitiation time decreases as the applied potential increases. At applied potentials near the Ercrev,the crevice corrosion initiation time increases. At applied potentials below the Ercrev, initiation oflocalized corrosion was not observed for Alloy 22. Although the initiation of crevice corrosionwas observed for Alloy 825 at an applied potential approximately 40 mV below the measured

0.001 0.01 0.1 1 10Chloride Concentration, molar

0.01

0.1

1

10

Max

imum

ani

on c

once

ntra

tion,

mol

ar

Maximum concentration of anions in NaCl at 95 oC

CO32-

HCO3-

SO42-

F-

Figure 4-12. Speciation Calculations Showingthe Maximum Anion Solubilities in Sodium

Chloride Solutions

4-14

Ercrev, the uncertainty in the determination of the Ercrev can be 50 mV (Dunn, et al., 2000). TheErcrev value represented in Figure 4-13 is based on a single measurement for each alloy.

Galvanic coupling potential and current data from an open circuit crevice corrosion test areshown in Figure 4-14. By galvanically coupling the crevice corrosion specimen to a largeAlloy 22 plate, both the current and the potential of the specimen can be monitored under opencircuit conditions. For this test, the area ratio of the Alloy 22 plate to the crevice specimen wasapproximately 5:1. Because only a fraction of the crevice specimen was covered by the creviceformer, the actual Alloy 22 plate crevice area ratio was approximately 40:1. The addition ofCuCl2 to the concentrated sodium chloride solution increased the Ecorr of the galvanic couple tovalues in the range of 260 mVSCE, and the galvanic couple current density was below10!7 A/cm2. Initiation of crevice corrosion is apparent from the sharp increase in current densityand the corresponding drop in potential of the galvanic couple. The potential of the galvaniccouple decreased and approached !100 mVSCE, or within 50 mV of the measured Ercrev value. After several hours, the specimen repassivated and the potential of the galvanic coupleincreased to values near 260 mVSCE. Several additional crevice corrosion initiation events wereobserved throughout the test, and each event was marked by a noticeable increase in currentdensity and a potential decrease that was typically more than 100 mV. Results fromFigure 4-13 suggest that the Ercrev is a valid parameter for the critical potential for localizedcorrosion initiation of nickel-chromium-molybdenum alloys in chloride solutions. The resultsfrom Figure 4-14, however, suggest that the decrease in Ecorr as a result of crevice corrosioninitiation can be significant. The rapid decrease in the potential of the galvanic couple is likelythe result of two factors; a rapid increase in the active area immediately following initiation, and

0.0001 0.001 0.01 0.1 1 10 100 1000Localized Corrosion Initiation TIme (days)

-200

-100

0

100

200

300

400

500

600

700

800

ΔE

= E

appl

ied -E

rcre

v (m

V)

Ercrev

Creivce Corrosion Initiation Time5.5 M NaCl, 100 oCClosed symbols = crevice corrosionOpen symbols = no crevice corrosion

Alloy 22Alloy 625Alloy 825

CPP Ecrev Alloy 22

CPP Ecrev Alloy 625

CPP Ecrev Alloy 825

Potentials in mVSCE Ecrev Ercrev

640 -145154 -260

-160 -320

Figure 4-13. Initiation Time for Crevice Corrosion As aFunction of Potential for Alloys 825, 625, and 22. Initiation Times Measured Under Potentiostatic

Conditions. Note: CPP—Cyclic potentiodynamicpolarization

4-15

a potential drop within the occluded region. With a sufficient reduction in the Ecorr, the potentialdriving force for active crevice corrosion cannot be sustained, and the crevice site repassivates.

Figure 4-15 shows results of an additional tests where the concentration of CuCl2 was increasedto raise the corrosion potential of the galvanic couple. At the start of the test, the potential of thecouple was 411 mVSCE. Localized corrosion was initiated within a few minutes and propagatedcontinuously for more than 3 days. Although the initiation of localized corrosion resulted in alarge potential drop of the couple, the potential remained above the Ercrev. After 3 days, a fewbrief periods of passivity were observed where the corrosion potential increased sharply, andthe current abruptly decreased. After 5 days, crevice corrosion repassivated, and the potentialof the couple was above 360 mVSCE. Although the potential remained high, no crevice corrosionwas reinitiated after 5 days.

Observation of the test specimen after test completion revealed crevice corrosion on most of the24 crevice sites created by the serrated PTFE crevice washer. The average and standarddeviation of the crevice attack was 297 ± 77 :m [0.012 ± 0.003 in]. Because multiple crevicesites were initiated, the propagation rate cannot be accurately determined. The results inFigure 4-15 suggest that the depth of crevice attack may be limited even in stronglyoxidizing conditions.

4.5 Localized Corrosion Propagation Rate Measurements

Galvanic coupling currents and potentials of the Alloy 22 specimens measured as functions oftime are shown in Figure 4-16. Only test 6 was conducted using the modified PTFE bolt. Aftera short induction period, (several seconds to 60 hours), the current increased as the potential

0 50000 100000 150000 200000 250000Time, seconds

-0.1

0

0.1

0.2

0.3

0.4

0.5

0.6

Gal

vani

c co

uple

pot

entia

l, V S

CE

1x10-9

1x10-8

1x10-7

1x10-6

Gal

vani

c co

uple

cur

rent

den

sity

, A/c

m2

Potential

Current density

Figure 4-14. Open-Circuit Corrosion Test for Alloy 22 in5 M NaCl with the Addition of CuCl2 at 100 °C [212 °F]. InitialPotential of the Couple Was 260 mVSCE. Test Performed byGalvanically Coupling an Alloy 22 Crevice Specimen to an

Alloy 22 Plate.

4-16

fell, indicating the initiation of localized corrosion. The increasing current was caused byactivation of crevice corrosion on the increasing surface area of active corrosion sites, whichwas driven by galvanic coupling to a proton reduction process on the platinum electrode. Withtime, the current and potential reached steady-state, suggesting stable propagation.

Figure 4-17 shows the optical micrographs of the corroded Alloy 22 specimens. Except forTest 6, the corrosion is localized at one side of the creviced area. The difference between twosides of the originally creviced area may have been caused by an uneven crevice tightness. InTest 6 where a spring was used, the corrosion appeared as pitting corrosion and the corrodedsites are more evenly distributed across the creviced area.

Table 4-2 summarizes the propagation durations and the corresponding maximum penetrationdepths from this series of tests. The plot of penetration depth as a function of propagation timein Figure 4-18a shows that at shorter times, the rate of penetration is rapid. At longer times thepenetration rate decreases. For a mass-transport controlled pit growth process, the penetrationfits Eq. (4-8) (Alkire and Wong, 1988; Beck and Alkire, 1979; Hunkeler and Böhni, 1981;Mughabghab and Sullivan, 1989; Shoesmith, et al., 1995).

d k t= 0 5. (4-8)where

d — maximum penetration deptht — timek — constant depending on the system

0 200,000 400,000 600,000 800,000

Time, seconds

-0.2

0

0.2

0.4

0.6

0.8

1

Gal

vani

c co

uple

pot

entia

l, V S

CE

10-10

10-9

10-8

10-7

10-6

10-5

10-4

10-3

Gal

vani

c co

uple

cur

rent

den

sity

, A/c

m2

Current density

Potential

Figure 4-15. Open Circuit Crevice Corrosion Test forAlloy 22 in 5 M NaCl with the Addition of CuCl2 at 100 °C[212 °F]. Initial Potential of the Couple Was 410 mVSCE. Test Performed by Galvanically Coupling an Alloy 22

Crevice Specimen to an Alloy 22 Plate.

4-17

Figure 4-16. Measured Current Density and Potential for Alloy 22 SpecimensGalvanically Coupled to Platinum at Various Durations in 4 M MgCl2 Solution at 110 °C

[230 °F]

-0.22

-0.12

-0.02

0.08

0.18

0.28

0.38

0.48

0 100 200 300 400 500

Time (hours)

Pote

ntia

l, V

SCE

Test 1 Test 2

Test 3 Test 4

Test 5 Test 6

1.0E-08

1.0E-07

1.0E-06

1.0E-05

1.0E-04

1.0E-03

0 100 200 300 400 500Time (hours)

Cur

rent

den

sity

, A/c

m2

Test 1 Test 2

Test 3 Test 4

Test 5 Test 6

4-18

Figure 4-17. Photographs Showing Localized Corrosion Features of Alloy 22Single-Crevice Specimens After Corrosion in 4 M MgCl2 Solution at 110 °C [230 °F]

Test 6. 12.9 daysTest 5. 20.0 daysTest 4. 13.1days

Test 3. 6.94 daysTest 2. 5.44 daysTest 1. 3.70 days

4-19

Table 4-2. Maximum Penetration Depths of Alloy 22 Crevice Specimens Propagated atVarious Durations in 4 M MgCl2 Solution at 110 °C [230 °F]*

Test NumberPropagation Durations

(Days)Maximum Penetration Depths

(:m)

1 3.70 46

2 5.44 116

3 6.94 56

4 13.1 130

5 20.0 128

6* 12.9 108

*In this test, the polytetrafluoroethylene (PTFE) bolt in Figure 4-1 in the in the test cell was modified to fit a springinside to compress the PTFE bolt against Alloy 22 specimen.

Maximum penetration depths were plotted against square root of propagation duration inFigure 4-18b. A preliminary fit of the experimental data to the equation yielded Eq. (4-9)

d tmax..= 312 0 5 (4-9)

The fit of the experimental data to Eq. (4-9) suggests that the propagation is a diffusioncontrolled process.

4.6 Coupled Multielectrode Array Sensors

The localized corrosion of Alloy 22 was measured in a 0.1 N FeCl3 solution at temperaturesfrom 18 to 90 °C [64 to 194 °F] using a multielectrode array sensor probe with 8 electrodesmade from 1-mm [0.04-in] diameter Alloy 22 wire (Yang, et al., 2002b). The apparent surfacearea of the electrode is 0.785 mm2 [1.22 × 10!3 in2]. The test was conducted together with threeother iron-nickel-chromium alloy multielectrode array sensor probes (Alloy 600, Type 316 and304 SS) for comparison. The composition of the alloys used in the experiments is shown inTable 4-3. Figure 4-19 shows the standard deviation of the currents from each of the fourmultielectrode array sensor probes as a function of time and the responses of the standarddeviations of the four probes to the change in temperature. For all the four multielectrode arraysensor probes, there were clear increases each time the temperature was increased.

Figure 4-20 presents the standard deviations that were averaged over time for the fourmultielectrode array sensor probes as a function of temperature. As shown in Figures 4-19 and4-20, the resistance of the metals to localized corrosion in the 0.1 M ferric chloride solution attemperatures up to 90 °C [194 °F] increases in the following order:

Alloy 600 < 304 stainless steel < 316 stainless steel < Alloy 22

4-20

10

30

50

70

90

110

130

150

0 5 10 15 20

Propagation duration, days

Max

imum

pen

etra

tion

dept

h, µ

m

(a)

Figure 4-18. (a) Maximum Penetration Depths Measured on Alloy 22 Specimens CreviceCorroded in 4 M MgCl2 for Various Propagation Durations at 110 °C [230 °F]; (b) Plot of

Maximum Penetration Depths Against Square Root of Propagation Durations. TheStraight Line Is the Linear Fit of the Experimental Data to Equation d = k t 0.5.

10

30

50

70

90

110

130

150

0 1 2 3 4 5 6

(Propagation duration)0.5, days0.5

Max

imum

pen

etra

tion

dept

h, µ

m

d = 31.2 t0.5

(b)

4-21

Table 4-3. Chemical Compositions (%wt) of the Metal Wires Used in the Multielectrode ArraySensor Probes

MetalsUNS*No. Ni† Cr Fe Mo Mn W Co

Si(ppm)

C(ppm)

Type 304 S30400 9.5 18.5 Bal.‡ NA§ < 2 NA NA NA < 800

Type 316 S31600 11 17.7 Bal. 3 < 2 NA NA NA < 1,200

Alloy 22 N06022 Bal. 21.2 4.5 13.3 0.25 3.0 1.9 400 30

Alloy 276 N10276 Bal. 15.4 6.1 15.48 0.4 3.6 0.73 540 60

Alloy 600 N06600 Bal. 15.5 8 NA NA NA NA NA < 1,500

*UNS—Unified numbering system.†Ni–nickel, Cr–chromium, Fe–iron, Mo–molybdenum, Mn–manganese, W–tungsten, Co–cobalt, Si–silicon,C–carbon‡Bal—Balance§—Not available

4 8 12 16 20Time (hour)

1E-010

1E-009

1E-008

1E-007

1E-006

1E-005

Sta

ndar

d D

evia

tion

of C

urre

nts

(A)

Sat. KCl

DI water

0.25 M FeCl3

1M NaNO3 Added

DI Water

0.0025 M FeCl3

DI Water

Figure 4-19. Typical Standard Deviation Signals from a25-Electrode Coupled Multielectrode Array Sensor Probe

Made of Type 304 SS in Different Environments (Yang andSridhar, 2003a). Reproduced with Permission from

NACE International.

4-22

1.0E-12

1.0E-11

1.0E-10

1.0E-9

1.0E-8

1.0E-7

1.0E-6

1.0E-5

1.0E-4

1.0E-3

7/25/200121:00

7/26/20013:00

7/26/20019:00

7/26/200115:00

7/26/200121:00

7/27/20013:00

Time

Sta

ndar

d D

evia

tion

of C

urre

nt (

A)

Alloy 22Alloy 600

316 SS304 SS

18oC41oC

60oC 90oC

316 SS

304 SS

Alloy 600

Alloy 22

Figure 4-20. Standard Deviation Signals of the Alloy 22and Three Other Probes in 0.1 M Ferric Chloride Solution atDifferent Temperatures (Yang, et al., 2002b). Reproduced

with Permission from NACE International.

PRE wt Cr wt Mo wt W N′ = + + +% . (% % )3 3 30% (4-11)

PRE at Cr at Mo at W at N′′ = + + +% . (% % )3 3 30% (4-12)

Furthermore, the corrosion resistance of Alloy 22 is more than 3 orders of magnitude better thanthe other three metals in the 0.1 M FeCl3 solution in terms of the localized corrosion currentmeasured from the multielectrode array sensor probes. This order is consistent with the pittingresistance equivalent numbers (PRE) calculated according to the elemental composition of thealloys using the following formulas (Rebak and Crook, 1999) (Table 4-4):

PRE wtCr wt Mo wt N= + +% .3 3% 20% (4-10)

where

%wt — weight percent%at — atomic percent

Eq. (4-10) has been commonly used for iron base alloys and Eqs. (4-11) and (4-12) areproposed for nickel base alloys.

Figure 4-21 shows that the localized corrosion currents measured from all multielectrode arraysensor probes increased with the increase in temperature, especially for the Alloy 22 probe. The temperature dependence of the Alloy 22 localized corrosion currents changed significantlywhen the temperature changed from 60 to 90 °C [140 to 194 °F]. The critical crevice corrosion

4-23

Table 4-4. Pitting Resistance Equivalent (PRE) Numbers Calculated According to theElemental Composition of the Multielectrode Array Sensor Alloys

PRE Numbers N06022 N10276 S31600 S30400 N06600

PRE (Eq. 1) 65.09 66.5 27.6 18.5 15.5

PRE' (Eq. 2) 74.99 78.4 27.6 18.5 15.5

PRE" (Eq. 3) 56.32 55.8 24.9 19.7 17.1

1.0E-10

1.0E-9

1.0E-8

1.0E-7

1.0E-6

1.0E-5

1.0E-4

0 20 40 60 80 100

Temperature (oC)

Sta

ndar

d D

evia

tion

(A)

Alloy 22316 SS304 SSAlloy 600

Figure 4-21. Relationship Between the Standard DeviationSignals from an Alloy 22 and Three Other Probes in 0.1 M

Ferric Chloride Solution and Temperature (Yang, et al.,2002b). Reproduced with Permission from

NACE International.

4-24

Figure 4-22. Posttest Examination of the Surface of an Electrode in the Alloy 22 Probe(Yang, et al., 2002b). Reproduced with Permission from NACE International.

temperature in chloride solutions containing Fe3+ ions is approximately 50 to 67 °C [122 to153 °F] according to a review by Cragnolino, et al. (1999). From Figure 4-21, the localizedcorrosion current for Alloy 22 at 18 °C [64 °F] was low (close to the lower detection limit of theinstruments, 1 × 10!11 A). The localized corrosion current increased, however, when thetemperature was changed to 41°C [106 °F]. This increase suggests that the critical localizedcorrosion temperature for Alloy 22 in the 0.1 M FeCl3 solution is probably even lower. It may bebetween 18 and 41 °C [64 to 106 °F].

Posttest examination confirmed that the most corroded electrode of the Alloy 22 probe wasslightly pitted (Figure 4-22). However, the most corroded electrodes of the Types 316 and304 SS probes were severely pitted (Figures 4-23 and 4-24) after the exposure. Because themost corroded electrodes were severely corroded, and others were not, on the Alloy 600 (notshown) and the Types 304 and 316 SS probes, the corrosion rate of these metals may becalculated with (Eq.) 4-2 by assuming the current distribution factor in Eq. (4-1), g, is equal tounity. Since the most corroded electrode of the Alloy 22 probe was only slightly corroded(Figure 4-22), however, its g was probably much less than unity. Therefore, its corrosion ratecould not be reliably estimated with Eq. (4-2). To use the coupled multielectrode array sensorprobe to measure the localized corrosion rate or to measure the bounding value of the corrosionrate, an improved method such as that described by Yang, et al (2005) should be used. Withthe improved method, the potential of the coupling joint of the multielectrode array sensor probeis slightly raised so that none of the unreacted areas on the most corroding electrode would beable to receive the electrons from the neighboring anodic sites. Therefore, the g is equal to one.

The localized corrosion current of Alloy 22 was also measured with the same Alloy 22multielectrode array sensor probe. Figure 4-25 shows the responses of the standard deviationsof the currents from the Alloy 22 multielectrode array sensor probe to the changes in NaClconcentration at room temperature along with the standard deviation signals from other three

4-25

Figure 4-23. Posttest Examination of the Surface of an Electrode in the Type 304 SSProbe (Yang, et al., 2002b). Reproduced with Permission from NACE International.

Figure 4-24. Posttest Examination of the Surface of an Electrode in the Type 316 SSProbe (Yang, et al., 2002b). Reproduced with Permission from NACE International.

4-26

1.0E-11

1.0E-10

1.0E-9

1.0E-8

1.0E-7

10/14/200112:00

10/15/200112:00

10/16/200112:00

10/17/200112:00

10/18/200112:00

Time

Sta

ndar

d D

evia

tion

of C

urre

nt (A

)

Alloy 22Alloy 276316 SS304 SS

De-ionized water

0.01M NaCl 0.1M NaCl0.002 M NaCl

0.5M NaCl 4M NaCl

316 SS

304 SS

Alloy 276

Alloy 22

Figure 4-25. Responses of the Standard Deviations of theCurrents from the Alloy 22 and Three Other Probes (Types 304 and316 SS and Alloys 276) to the Changes in Chloride Concentration

at Room Temperature (Yang, et al., 2002b). Note: The DashedLines Indicate the Time at Which the Chemistry Concentration

Was Changed. Reproduced with Permission fromNACE International.

probes (Alloy 276, Types 304 and 316 SS). The standard deviation of the Alloy 22multielectrode array sensor probe was initially 3 × 10!9 A in deionized water immediately after itwas polished, but it decayed to 5 ×10!11 A over 3 days. After the probe became passivated inthe deionized water, it did not respond to the addition of chloride at concentration levels up to4 M. In contrast, the probe signals from the other probes (Alloy 276 and Types 316 and304 SS) increased with each step increase in chloride concentration. The low localizedcorrosion current shown in Figure 4-25 for Alloy 22 is consistent with the measured criticallocalized corrosion temperature, which is approximately 60 to 80 °C [140 to 176 °C] for Alloy 22crevice specimens in NaCl solutions at concentrations above 0.5 M (Dunn, et al., 2003b).

For Types 304 and 316 SS and Alloy 22, the order of the resistance to localized corrosion is thesame as that described for the ferric chloride tests. The resistance of Alloy 276 is between thatof the Type 316 SS and Alloy 22. This order of resistance is consistent with the industrialexperience for these alloys in chloride solutions. The quantitative corrosion rate of Alloy 22could not be reliably calculated because of uncertainty in g for the Alloy 22 probe, which doesnot experience significant localized corrosion in the NaCl solution.

5-1

i T i eC (T)

C (T )er r

refER T T O

bulk

Obulk

ref

Z FRT

ao

refr r

f s01 1

2

2

( )/

= −− −

⎛

⎝⎜

⎞

⎠⎟ −

βφ

(5-2)

5 MODELS FOR WASTE PACKAGE DEGRADATION

In this chapter a performance assessment model for the onset of localized corrosion isdiscussed, including equations for computing the corrosion potential, the repassivation potential,and values for equation parameters. Also a performance assessment model for estimatingcompositions of solutions arising from dynamic evaporation is presented. The localizedcorrosion and environmental models are used to estimate the probability for the waste packageto exhibit localized corrosion assuming it is contacted by seeping water.

5.1 Corrosion Potential Model

Equations programmed in the TPA code Version 5.0 predict an increasing corrosion potentialwith increasing temperature. This trend is opposite to that shown by experimental data(i.e., corrosion potential decreases with increasing temperature). On the other hand, TPA 5.0equations predict decreasing corrosion potential with increasing pH, consistent withexperimental trends. However, the corrosion potential computation implemented in the TPAcode decreases approximately 500 mV for every unit increase in pH, which is a steeperdependence than observed experimentally. In general, the TPA code Version 5.0overestimated the corrosion potential compared to experimental data. This overestimationresults in much higher frequency of localized corrosion on waste packages. To implement amore realistic model for the onset of localized corrosion, the equations for the computation ofcorrosion potential have been revised. This section documents the revised equations for thecomputation of corrosion potential and the mechanistic rationale for the changes.

The reduction of oxygen or water results in consumption of electrons. This “negative” current isreferred to as cathodic current density. On the other hand, electrons are released from metaldissolution. This “positive” current is referred to as anodic current density. The total current isthe addition of the negative (cathodic) and positive (anodic) contributions. The corrosionpotential is the potential at which the total current density is zero. The likely dominant cathodicreaction under oxidizing conditions and neutral to alkaline pH is

O H O 4 e 4 OH2 2+ + →− −2 (5-1)

Assuming a first order reaction with respect to oxygen, O2, and ignoring activity coefficientcorrections, the contribution to the cathodic current density from oxygen reduction, , as ai r

0

function of temperature and applied potential at the oxide-solution interface can be computed as(Sridhar, et al., 1993)

5-2

where (only international units are provided; English system requires a modified equation)

— reference current density, A/m2i rref

— activation energy for the reference current density, J/molEao

R — ideal gas constant, 8.314 J mol!1 K!1 T — temperature, KTref — reference temperature, equal to 298 K

— oxygen concentration in solution as a function of temperature, mol/kgC TObulk

2( )

— oxygen concentration in solution at the reference temperature,C TObulk

ref2( )

mol/kg ~ mol/LZr — number of electrons involved in the fundamental charge transfer reduction

reaction, 4$r — charge transfer coefficient for the reduction reaction, dimensionlessF — Faraday’s constant, 9.64867 × 104 Coul/molNf/s — potential drop at the film-solution interface, V

The concentration of oxygen in solution as a function of temperature can be estimated using theempirical relationship in Battino (1981) for a pure water system

C T pO eObulk T T

2

3 6 2

2

2

0 2984 5 59617 10 1 04967 10

( ). . .

=−

×+

×K K-1 -1mol kg atm (5-3)

where the temperature, T, is in units of K and the partial pressure of oxygen over the solution,pO2, in units of atm.

Consistent with the Point Defect Model (Macdonald, 1992), the potential drop at the film-solutioninterface, Nf/s, is assumed to be linearly related to the corrosion potential

φ α α α φf s E corr Ho

Oo

f soE RT

F cRTF

pOp/ /log [ ] log= −

⎛

⎝⎜

⎞

⎠⎟ −

⎛

⎝⎜

⎞

⎠⎟ +

+

10 102H (5-4)

where

"E, "H, "o — dimensionless constants[H+] — hydrogen ion concentration, mol/Lco — constant equal to 1 mol/Lpo — constant equal to 1 atm

— constant potential, V φ f so

/

The original equation for the potential drop at the film-solution interface, Nf/s, in Macdonald(1992) was generalized by including a term dependent on the partial pressure of oxygen, pO2,and an explicit dependence on temperature. After algebraic manipulation, the expression forthe partial cathodic current density due to oxygen reduction becomes

5-3

[ ]i T i

C T

C Te e

cpOpr

Oref O

bulk

Obulk

ref

ERT

Z FRT

E

o

n

o

naef

r ref

corrH

O

( )( )

( )= −

⎛

⎝

⎜⎜

⎞

⎠

⎟⎟

⎛

⎝⎜

⎞

⎠⎟

− −+

2

2

2β H

(5-5)

where

— positive constant, A/m2 i ref

— effective activation energy, J/molEaef

— effective charge transfer coefficient, dimensionlessβ ref

nH, nO — dimensionless constants

Equation (5-5) was implemented in a revised version of the TPA code (Version 5.0.1 currentlyunder development), with input parameters , , , and nH defined by the user. In thei r

ef Eaef β r

ef

TPA code it is assumed that the ambient partial pressure of oxygen is constant (equal to0.21 atm) for the computation of the corrosion potential. Both the partial pressure of oxygenand the constant nO are assigned fixed values that cannot be modified by the user to avoidprediction of corrosion potentials outside experimentally tested ranges of oxygenpartial pressure.

The constant is defined asβ ref

β α βref

E r= (5-6)

Since the charge transfer coefficient, $r, and the constant "E are both positive and less thanone, is also positive and less than one. Likewise, the constants , , nH, and nO canβ r

ef i ref Ea

ef

be written as functions of terms in Eqs. (5-2) and (5-4). Based on Center for Nuclear WasteRegulatory Analyses calculations, it can be concluded that the constants and arei r

ef Eaef

positive, and the constants nH and nO can be of any value. Values for the constants , ,i ref Ea

ef

nH, and were estimated by comparing computed values of corrosion potential toβ ref

experimental data. The parameter estimation approach is later described in this report.

Assuming that O2 is transported a distance δ before being reduced at the metal-film interfaceand assuming steady-state conditions, Sridhar, et al. (1993) derived the following expression forthe partial contribution to the cathodic current density by the oxygen reduction reaction

i T i Ti T C T

C T Z F D

rrO

rO

Obulk

ref

Obulk

r r

( ) ( )( ) ( )

( )

=

−1 2

2

δφ τ

(5-7)

where

— partial cathodic current density from Eq. (5-5), A/m2i TrO ( )

Dr — diffusion coefficient of oxidizing species, m2/sN — porosity, dimensionlessJ — tortuosity, dimensionless

5-4

Equations (5-5) and (5-7) are very similar to the equations used in the TPA Version 5.0 code(see Mohanty, et al., 1997, for a detailed discussion of the original TPA equations). The maindifferences are the effective charge transfer coefficient, , concept and the introduction of aβ r

ef

constant, nH, controlling the pH dependence of the partial cathodic current density.

Equation (5-7) is valid for neutral to alkaline solutions. Under acidic conditions, a reaction likelyto dominate the cathodic current density is

O 4 H 4 e 2 H O2 2+ + →+ − (5-8)

Following the derivation previously described, an expression of identical form to Eqs. (5-5) and(5-7) can be derived for the cathodic current density associated with Eq. (5-8). Thecorresponding constants , , nH, and nO are expected to have different values in thei r

ef Eaef

acidic regime. Different values for the acidic and alkaline pH range are reported later inthis report.

The original equations in the TPA code, described in Mohanty, et al. (1997), considered waterreduction as another potential contributor to the cathodic current density

2 H O 2 e H 2 OH2 2+ → +− − (5-9)

The same equations in Mohanty, et al. (1997) were adopted to compute the partial cathodiccurrent density due to water reduction. In the revised version of the TPA code it is assumed,however, that the partial contribution to the cathodic current density by the water reductionreaction is negligible. Since the Yucca Mountain environment will be oxidizing due to thepotential repository location in the unsaturated zone, the water reduction contribution to thecathodic current density is likely to be negligible [Eq. (5-9) is important under reducingconditions]. The equation implemented in the TPA code to compute the current density due towater reaction is

i T i rw wref

ER T T

Z FRT

Eaw

refw w

corr( ) = −

− −⎛

⎝⎜

⎞

⎠⎟ −

1 1

eβ

(5-10)

where

— reference current density, A/m2iwref

— activation energy for the reference current density, J/molEaw

$w — charge transfer coefficientZw — number of electrons involved in the fundamental charge transfer reduction

reaction, 1

The total cathodic current density is computed as the sum of the two partial contributions

i T i T i Tc r w( ) ( ) ( )= + (5-11)

5-5

In the revised version of the TPA code, a low value of the reference current density, , isiwref

assumed, so thati T i Tw r( ) ( )<< (5-12)

Therefore, TPA code results are insensitive to the water reduction reaction parameters.

In the revised version of the TPA code, the anodic current density, ia(T), is computed accordingto the equation

i T i ea ao

ER T Taa

refa

( ) =− −

⎛

⎝

⎜⎜

⎞

⎠

⎟⎟

1 1(5-13)

where

— reference temperature, KTrefa

— value of the current density at the reference temperature, A/m2i ao

— activation energy for the anodic current density, J/molEaa

Equation (5-12) differs from the approach in TPA Version 5.0 code by the introduction of anactivation energy, , to model the temperature dependence of the anodic current density. AEa

a

value for this activation energy was estimated in Pensado, et al. (2002).

As mentioned previously, the corrosion potential is defined as the potential at which the totalcurrent density is zero, that is

i T i Ta c( ) ( )+ = 0 (5-14)

Anodic passive dissolution is very slow in the absence of environmental conditions leading tolocalized corrosion, with current densities in the order of 10!8 A/cm2 [9.29 × 10!6 A/ft2]. Thus,cathodic reactions most likely will be controlled by kinetic processes and not by diffusion ofoxidizing species (i.e., the diffusive flux is fast compared to kinetic rates). Therefore, for theAlloy 22-water system, it is reasonable to assume that the diffusive path length, *, is negligible. Assuming * = 0 and = 0, an analytical expression for the corrosion potential can be derivediw

ref

E E EZ F

EZ F

TT

RTZ F

Hc

pOp

ii

C TC Tcorr

aa

aef

r ref

aa

r ref

refa

r ref

o

n

o

n

ref

ao

Obulk

Obulk

ref

H O

=−

− +⎛

⎝⎜

⎞

⎠⎟

⎛

⎝⎜

⎞

⎠⎟

⎡

⎣⎢⎢

⎤

⎦⎥⎥

+

β β βln [ ] ( )

( )2 2

2

(5-15)

The parameters , nH, and were estimated, for acidic and alkaline solution ranges, byβ rer i r

ef

performing a least square fitting to measurements of the corrosion potential as a function oftemperature and pH. Values for these parameters, as well as assumed values for the otherparameters, are listed in Table 5-1.

In the TPA code a general approach is implemented to numerically find the root of Eq. (5-14). The general approach computes a value for the corrosion potential for cases

5-6

Table 5-1. Parameters to Simulate the Corrosion Potential As a Function ofTemperature and pH. Only Values Appropriate for Use with the Revised Version of theTPA Code Are Listed. Computed Corrosion Potentials Using Eq. (5-15) Are in Units of

Volts Versus Standard Hydrogen Electrode (VSHE).

ParameterValuepH <6

ValuepH >6 Justification

(kJ/mol)Eaa 44.7 44.7 Value estimated*

(kJ/mol)Eaef 40 40 Value for the activation energy used in the

TPA Version 4.0 code†

[Coul/(m2 yr)]i ao 3,200 3,200 Value derived from Eq. (4-3)* at =Tref

a

368.15 K

(K)Trefa 368.15 368.15 Selected reference temperature for the

anodic current density. This referencetemperature can be arbitrary, but the entry inthe above row in this table must beconsistent with this selected temperature.

Tref (K) 298.15 298.15 Selected reference temperature

β rer 0.01287 0.0248 Value derived by curve fitting

nH 0.0256 0.01897 Value derived by curve fitting

[Coul/(m2 yr)]i ref 7.57 × 109 5.51 × 109 Value derived by curve fitting

nO 0.01287 0.0248 Selected nO= to preserve the sameβ rer

functional dependence on the partialpressure of oxygen as in Mohanty, et al.‡

*Pensado, O., D.S. Dunn, G.A. Cragnolino, and V. Jain. “Passive Dissolution of Container Materials—Modeling and Experiments.” CNWRA 2003-01. San Antonio, Texas: Center for Nuclear Waste RegulatoryAnalyses. 2002.†Mohanty, S., T.J. McCartin, and D.W. Esh, coords. “Total-system Performance Assessment (TPA) Version 4.0Code: Module Descriptions and Users Guide (Revised).” San Antonio, Texas: Center for Nuclear WasteRegulatory Analyses. 2002.‡Mohanty, S., G.A. Cragnolino, T. Ahn, D.S. Dunn, P.C. Lichtner, R.D. Manteufel, and N. Sridhar. “EngineeredBarrier System Performance Assessment Code: EBSPAC Version 1.1, Technical Description and User’s Manual.” CNWRA 97-006. San Antonio, Texas: Center for Nuclear Waste Regulatory Analyses. 1997.

5-7

when accounting for diffusive control is needed and when the contribution by water reduction tothe total current density is significant.

In the revised TPA abstraction, variability in the corrosion rate is assumed to be due to thevariability in the Alloy 22 passive current density, . The parameter is sampled from ai a

o i ao

triangular distribution ranging from 1,600 C/(m2 yr) to 6,400 C/(m2 yr) with mode at3,200 C/(m2 yr), at a reference temperature, , equal to 368.15 K. The variation of theTref

a

corrosion potential as a function of pH and temperature is displayed in Figure 5-1. The upperbound, most likely estimate, and the lower bound in the corrosion potential were computed byassuming = 1,600 C/(m2 yr), 3,200 C/(m2 yr), and 6,400 C/(m2 yr), respectively. The TPAi a

o

code range of corrosion potentials appropriately encloses the range of experimental values. Therefore, assuming constant values for all parameters listed in Table 5-1, with the exception of

, is appropriate for estimating values of the corrosion potential for the performancei ao

assessment abstraction. The assumption nO = was selected to adopt an identical functionalβ rer

dependence of the corrosion potential on the oxygen partial pressure from previous versions ofthe TPA code (e.g., see Mohanty, et al., 1997). The capability of Eq. (5-15) to predict thecorrosion potential at partial pressures other than pO2 = 0.21 atm has not beenexperimentally verified.

5.2 Localized Corrosion Model

In the TPA code, localized corrosion is assumed to initiate if the corrosion potential exceeds acritical potential for localized corrosion, Ercrev, defined as

E T A A T B B Tc

Eo

rcrev rcrevCl( ) ( ) log [ ]

= + + + +−

1 2 1 2 10 Δ (5-16)

In the TPA code abstraction, the critical potential is computed only if an aqueous environmenton the waste package surface is established and if the chloride concentration, [Cl!], exceeds acritical value, [Cl!]crit. Values and distribution for the terms A1, A2, B1 and B2 used in the TPAcode are listed in Table 5-2 for mill-annealed Alloy 22 and for thermally aged material (intendedto simulate Alloy 22 welds). The term )Ercrev is a correcting term to account for the effect ofinhibitors, defined as a function of inhibiting oxyanion concentrations (nitrate, sulfate, andcarbonate) and chloride. Variability in the critical potential as a function of temperature,chloride, and oxyanion concentration is introduced in the TPA code abstraction from theassumed variability/uncertainty in the term A1. A symmetric envelope of 100 mV around mostlikely values of the repassivation potential as a function of the environmental variables isassumed. Extrapolation of Eq. (5-16) to temperatures higher than 95 °C [203 °F] for thermallyaged material results in underestimation of the critical potential, which is conservative. Experimentally it was demonstrated that, as temperature increases, the rate of decrease in thecritical potential is less than linear with respect to temperature variations. Thus, assuming thelinear relationship with respect to temperature in Eq. (5-16) would yield lower values of thecritical potential than were experimentally determined at temperatures higher than 95 °C[203 °F]. In this analysis, Eq. (5-15) was extrapolated to 110 °C [230 °F], which would yieldconservative estimates of the probability of localized corrosion on welded waste package areas.

5-8

-200

0

200

400

600

800

1000

1 3 5 7 9 11 13pH

E cor

r, m

V SH

E

TPA upper bound

TPA lower bound

TPA upper bound

TPA lower bound

TPA most likely estimate


Acidic range Neutral to alkaline range

95 °C [203 °F]

(e)

Figure 5-1. Comparison of Computed Corrosion Potentials to Experimental Data(Diamonds) As a Function of pH and Temperature (mVSHE = mVSCE + 241.2 mV); (a) 25 °C

[77 °F], (b) 40 °C [104 °F], (c) 60 °C [140 °F], (d) 80 °C [176 °F], and (e) 95 °C [203 °F]

-200

-50

100

250

400

550

700

850

1000

7.5 8.5 9.5 10.5 11.5 12.5pH

Eco

rr, m

VS

HE


TPA upper bound

TPA lower bound

25 °C [77 °F]

(a)

-200

-50

100

250

400

550

700

850

1000

7.5 8.5 9.5 10.5 11.5 12.5pH

Eco

rr, m

VS

HE


TPA upper bound

TPA lower bound

40 °C [104 °F]

(b)

-200

-50

100

250

400

550

700

850

1000

7.5 8.5 9.5 10.5 11.5 12.5pH

Eco

rr, m

VSH

E


TPA upper bound

TPA lower bound

60 °C [140 °F]

(c)

-200

-50

100

250

400

550

700

850

1000

7.5 8.5 9.5 10.5 11.5 12.5pH

Eco

rr, m

VS

HE


TPA upper bound

TPA lower bound

80 °C [176 °F]

(d)

5-9

Table 5-2. Parametersand Distributions Used to Define the Critical Potential As aFunction of the Temperature and Chloride Concentration in the TPA Code

MetallurgicalCondition

Temperature Rangeof Validity(°C) [ °F]

[Cl!]crit (mol/L)

A1(mVSHE)

A2(mV/°C)

B1(mV)

B2(mV/°C)

Mill-Annealed 80 to 125[176 to 257]

0.5 Triangulardistribution:1541.2,1591.2,1641.2

!13.1 !362.7 2.3

ThermallyAged

60 to 95[140 to 203]

Log-uniformdistribution:0.01, 0.25

Triangulardistribution:991.2,1041.2,1091.2

!10.0 !584.2 3.7

The term )Ercrev is used to account for the inhibiting effect of oxyanions on localized corrosion. When oxyanions such as nitrate, sulfate, or bicarbonate are present in the system, the criticalpotential sharply increases. The term )Ercrev is computed as

ΔE r rr

n

nrcrev mV= 800 min( , )

(5-17)

The term, r, is a linear combination of oxyanion to chloride concentration ratios defined as

r rr

rr

n

s

n

c

= + ++−

−

−

−

− −

−

[NO ][Cl ]

[SO ][Cl ]

[CO ] [HCO ][Cl ]

3 4 32

32

(5-18)

The reference ratios rn, rs, and rc are determined from single-inhibitor type solution (nitrate,sulfate, or carbonate) experiments. The repassivation potential is observed to increase steeplywith increasing oxyanion-to-chloride ratio, up to a maximum ratio (referred to as rn, rs, and rc fornitrate, sulfate, and carbonate-bicarbonate solutions, respectively) above which no additionalincrease in repassivation potential is observed. For the TPA code abstraction, to account formixed solutions, an additive effect to the repassivation potential from the various oxyanions wasassumed [Eq. (5-18)]. The reference ratios for mill-annealed and thermally aged material aredifferent and are listed in Table 5-3.

The selection of Ercrev as the critical potential for the initiation of localized corrosion is based onresults obtained with Alloy 22 and similar nickel-chromium-molybdenum alloys that aresusceptible to crevice corrosion in chloride solutions, but resistant to pitting corrosion. Formation of crevices on waste packages in a repository system may be limited to the outercontainer-pallet contact area. Additional crevice sites could result from direct contact betweenthe waste package and drip shield caused by drift degradation and rockfall. As described in

5-10

Table 5-3. Values of the Reference Ratios to Compute the Term )Ercrev in the TPACode Abstraction

Metallurgical condition rn rs rc

Mill-annealed 0.1 0.5 0.2

Thermally aged 0.3 0.5 0.2

Chapter 3, crevice corrosion specimens used PTFE crevice washers to form reproduciblecrevice sites. Actual crevices formed on the waste package outer container surfaces may besignificantly larger compared to crevice sites in test specimens. Actual crevices are likely to beformed by metal to metal or metal to rock rubble contacts. However, the results obtained usingAlloy 22 specimens with PTFE crevices are appropriate to represent crevices encounteredduring post-closure conditions.

5.3 General Corrosion Model

If conditions for the onset of localized corrosion (Ecorr > Ercrev) are not present, it is assumed inthe performance assessment model that Alloy 22 undergoes passive dissolution indefinitely inan aqueous environment. The general corrosion model TPA is described in Mohanty, et al.(2002). In the revised version of the TPA code, the model was modified to account for thetemperature dependence of the passive current density. The Arrhenius dependence ontemperature derived in Pensado, et al. (2002) is adopted (activation energy of 44.7 kJ/mol inTable 5-1). The passive current density at a reference temperature, 368.15 K, is defined asinput data for the revised performance assessment model. At any time, the passive currentdensity and the general corrosion rate are computed as functions of the waste packagetemperature in the revised version of the TPA code. The extent of the penetration of thegeneral corrosion front is computed according to the algorithm described inMohanty, et al. (2002). Waste package failure is assumed to occur when the corrosion fronthas penetrated a distance equal to the initial waste package thickness.

The dominant assumption in the model for general corrosion is that if environmental conditionsleading to localized corrosion or loss of passivity are not established, passive dissolutionprevails indefinitely in aqueous environments. This assumption holds true unless the passiveoxide evolves into a less protective composition or structure. The composition and structure ofthe passive film are functions of the alloy substrate. Therefore, the only reason for the oxidecomposition/structure to change over time is as a result of a change in the alloy substrate. Onemechanism that could induce changes at the alloy-oxide interface is preferential dissolution ofan alloy element. For example, if nickel dissolves faster than other alloy elements(e.g., chromium, molybdenum) and a planar interface is preserved during the dissolutionprocess, the alloy close to the metal-oxide interface would eventually develop a differentcomposition from the bulk alloy (the interface would be deficient in nickel, in this example). Ifthe oxide breaks for any reason, a newly formed oxide may not necessarily be of the samecomposition/structure as the original oxide, given the different nature of the exposed alloy. There is no evidence that a preferential dissolution process is established by the passivedissolution process; however, additional studies could produce evidence that long-term passivedissolution is stoichiometric. This would enhance the mechanistic basis for the assumption thatpassive dissolution can be sustained indefinitely in the absence of environmental conditions

5-11

promoting localized corrosion or loss of passivity. Studies performed on josephinite (a rockcontaining naturally occurring nickel-iron alloy) indicate stoichiometric composition of corrosionproducts in the rock with respect to the natural alloy composition (Cragnolino, et al., 2004b). Assuming similar solid-state diffusion of atoms for the main alloying elements, Pensado, etal. (2002) concluded that passive dissolution evolves toward stoichiometric dissolution, if themetal-oxide interface retains a planar shape.

Another mechanism that could induce changes to the alloy in the neighborhood of themetal-oxide interface is intergranular oxidation and establishment of fast diffusion paths alonggrain boundaries. There is no evidence of intergranular oxidation in 5-year experiments by theU.S. Department of Energy (DOE). It is likely that oxygen does not have enough thermalenergy to significantly diffuse into the alloy at repository relevant temperatures {less than 300 °C[572 °F]}. Future experiments could be estimate the extent of intergranular oxidation, fromhigh-temperature experiments in air extrapolated to lower temperatures, to demonstrate thatoxygen diffusion inside the alloy is likely negligible at repository relevant temperatures.

In summary, processes with the potential to affect passive dissolution over a long time period (inthe absence of environmental conditions leading to localized corrosion or loss of passivity) arerelated to changes in the substrate alloy in the region close to the metal-oxide interface. Anindirect way to determine that such processes may be occurring is by measuring the extent ofcongruency of the passive dissolution process. Nonstoichiometric dissolution, sustained overextended periods, would suggest a changing alloy composition over time. Currently, there is noevidence of long-term nonstoichiometric dissolution. To enhance the mechanistic justificationfor the assumption of long-term passive dissolution, experiments could measure the extent ofcongruency of the passive dissolution process and estimate the extent of oxygen penetrationinto the alloy over time at repository relevant temperatures. If the alloy composition in theregion close to the metal-oxide interface does not change, disrupting the oxide film has noimplication to waste containment because a fresh passive oxide with similar properties to theoriginal oxide would reform. The only processes currently envisioned that could compromisethe capability of the waste package to contain radioactive waste (other than mechanicalinteractions) depend on establishing environmental conditions leading to localized corrosion orloss of passivity. The next section discusses the probability of the establishment ofsuch conditions.

5.4 Environmental Model

The lifetime of the waste packages will depend on the quantity and chemistry of water thatcontacts them. There are three potential sources of this water—seepage into the drifts fromoverlying rocks, condensation due to in-drift cold trap processes, and deliquescence of salts industs that have deposited on the waste package surface (Browning, et al., 2004). The initialcomposition of seepage and condensed waters will likely differ significantly and will evolve overtime as a result of evaporation, mineral precipitation, interaction with rocks and engineeredmaterials, and mixing. Deliquescence of salts sequesters moisture from the atmosphere andgenerate brines that could initiate waste package corrosion. Deliquescence occurs underspecific temperature and relative humidity conditions and is strongly dependent onsalt compositions.

Spatial and temporal variations in temperature and relative humidity within the drifts stronglyaffect the quantity and chemistry of in-drift waters (Fedors, et al., 2004; Manepally and

5-12

Fedors, 2003). It is appropriate, but difficult, to evaluate waste package lifetime in the context ofthese complexly evolving in-drift environmental conditions. To simplify the identification andevaluation of plausible scenarios for aqueous corrosion of waste package materials, threeenvironments—Environments 1, 2, and 3—that are distinguished by the temperature at the driftwall and in-drift relative humidity were defined for the abstraction of near-field environments(Mohanty, et al., 2004; Browning, et al., 2004). In Environment 1, seepage and condensatewaters do not contact the waste package because the drift wall temperature is high (greaterthan a seepage threshold temperature, nominally set to a value around 105 °C [221 °F] for theperformance assessment abstraction), such that water from overlying rocks does not seep intothe drift, and condensate does not form on drift walls. In this environment, the only feasiblesource of water to initiate waste package corrosion is the deliquescence of salts on the wastepackage surface. In Environment 2, the drift wall temperature is below the seepage thresholdtemperature, such that water can seep from overlying rocks. Significant evaporation of seepagewater can occur in Environment 2 because the waste package temperature is higher than thedrift wall temperature, and the in-drift relative humidity is relatively low (< 90 percent). InEnvironment 3, the drift wall temperature is lower than in the other two environments and in-driftrelative humidity is high (> 90 percent). Under these conditions, evaporation of seepage wateris significantly reduced, and the in-drift waters are likely to be dilute and not cause wastepackage corrosion. In the TPA code abstraction, the waste packages will not necessarily beexposed to all three environments. For example, if the drip shields remain intact for anextended period of time, the waste packages will experience only Environments 1 and 3.

Available data on the composition of dusts sampled inside the Exploratory Studies Facility(Peterman, et al., 2003) and of atmospheric dusts sampled in the Yucca Mountain vicinity[National Atmospheric Deposition Program (NRSP–3)/National Trends Network, 2004] showthat nitrate, a corrosion-inhibitor, is an abundant component. If the limited data on dustcompositions are representative of dusts that may deposit on waste package surfaces in thepotential Yucca Mountain repository, localized corrosion of the waste package material inEnvironment 1 is unlikely because of sufficient concentrations of nitrate. As noted previouslyparagraph, waste package corrosion also is unlikely in Environment 3. Thus, Environment 2has the most potential for aqueous corrosion of the waste package because seepage watersmay deliver significant amounts of chloride ions onto the waste package surface. Subsequentevaporation of these waters can generate brines with high chloride concentrations. On theother hand, the same evaporation process could also increase the concentration of corrosion-inhibiting oxyanions such as nitrate, sulfate, and carbonate on the waste package surface.

As discussed in the previous section, in the TPA code localized corrosion is assumed to initiateif the corrosion potential, Ecorr, exceeds a critical potential, Ercrev, for localized corrosion. Ecorr is afunction of temperature and pH, whereas Ercrev depends on temperature and on theconcentrations of chloride, nitrate, sulfate, carbonate, and bicarbonate anions. To providevalues of pH and anion concentrations ([Cl!], [NO3

!], [CO32!] + [HCO3

!], [SO42!]) for the TPA

abstraction, thermodynamic simulations of the chemical evolution of in-drift waters resultingfrom evaporation of seepage waters were conducted. The simulations allowed determination ofthe types of brines that may form and the ranges of brine chemistry that may contact the wastepackage during Environment 2 (if the drip shield is not intact). The thermodynamic calculationswere supplemented by an alternative approach based on the concept of chemical dividesdeveloped by Hardie and Euster (1970). A similar approach was used in previous studies anddiscussed in Pabalan, et al. (2002a,b). In the chemical divide concept, the chemical types ofbrines and salt minerals that form upon evaporation of natural waters are determined by early

5-13

Figure 5-2. Ternary (Ca-SO4-HCO3) Phase Diagram Plotting Yucca Mountain UnsaturatedZone Porewater Compositions (+) Reported by Yang, et al. (2003, 1998, 1996) “ChemicalDivide” Lines Separate the Compositions into Three Brine Types: (i) Calcium-chloride,

(ii) Neutral, and (iii) Alkaline (Hardie and Eugster, 1970). The Compositions Representedby the Diamond Symbols Were Used as Input to the Thermodynamic Simulations.

precipitation of insoluble minerals (e.g., calcite and gypsum), and natural waters are consideredto evolve into three types of brines upon evaporation: (i) calcium-chloride, (ii) neutral, and(iii) alkaline. The calcium-chloride, neutral, and alkaline brine types are equivalent, respectively,to the calcium-chloride, sulfate, and carbonate brine types referred to in DOE documents(e.g., Bechtel SAIC Company, LLC, 2003b).

The evaporation simulations were conducted using LabAnalyzer and StreamAnalyzerVersion 1.3 (OLI Systems, Inc., 2004). These codes are more user-friendly versions of theEnvironmental Simulation Program (ESP) Version 6, which was used in the previous studies. The three codes, all developed by OLI Systems, Inc. (Morris Plains, New Jersey), use the samethermodynamic database. The codes allow simulation of aqueous chemical systems fortemperatures to 300 °C [573 °F], pressures to 1,500 bar [21,800 psi], and ionic strengths up to30 molal. Twenty-nine water compositions were used as input to the evaporation simulation. These compositions were selected to represent the broad composition range of the more than150 samples of Yucca Mountain unsaturated zone porewater reported by Yang, et al. (2003,1998, 1996). An implicit assumption in the use of these data is seepage water in Environment 2will have chemical characteristics similar to Yucca Mountain unsaturated zone porewaters. The unsaturated zone porewater compositions reported by Yang, et al. (2003, 1998, 1996), and thecompositions selected for the evaporation simulations are shown in Figure 5-2.

5-14

Table 5-4. Rank Correlations to Be Used in a Performance Assessment Abstraction

Correlated Quantities Spearman Rank Correlation Coefficient

pH, [Cl!] !0.8

pH, [CO32!] + [HCO3

!] 0.9

[Cl!], [CO32!] + [HCO3

!] !0.8

The evaporation simulations were done at 110 °C [230 °F], a temperature that can be sustainedat the waste package surface for several hundred years, and at a constant pressure of 0.85 bar[12.3 psi], the approximate atmospheric pressure at Yucca Mountain, or at 1 bar [14.5 psi]. Thecalculated pH, [Cl!], [NO3

!], [CO32!] + [HCO3

!], and [SO42!], grouped using the three brine types

of Hardie and Eugster (1970), are shown as box plots in Figure 5-3. The figure shows theranges in pH and anion concentration that result from the evaporation of seepage waters withcompositions shown in Figure 5-2. As shown in Figure 5-3, the three water types exhibit distinctchemical characteristics upon evaporation.

The thermodynamic simulation of seepage water evaporation used only a subset of theunsaturated zone porewater chemistry data reported by Yang, et al. (2003, 1998,1996)because the simulations are computationally intensive. To provide a basis for estimating thefrequency of occurrence of the three brine types and their respective chemical characteristics,the full set of Yang, et al. (2003, 1998, 1996) data on unsaturated zone porewater compositionswas used with the chemical divide concept of Hardie and Eugster (1970). A previous study(Pabalan, et al., 2002a) indicated there is generally good agreement between thethermodynamic simulations and the chemical divide approach, and the latter could providequalitative information on the chemistry of seepage waters that have undergone evaporativeconcentration. In the Yang, et al. (2003, 1998, 1996) reports, chemical compositions for156 porewater samples from the Yucca Mountain unsaturated zone were listed. Of these156 compositions, 13 are calcium-chloride type, 37 are neutral type, and 106 are alkaline type. Based on these numbers and for the purpose of TPA abstraction, it is assumed thatEnvironment 2 seepage waters will have chemical characteristics of calcium-chloride-, neutral-and alkaline-type brines with a frequency of 8, 24, and 68 percent, respectively, and withassociated ranges in pH, [Cl!], [NO3

!], [CO32!] + [HCO3

!], and [SO42!] illustrated in Figure 5-3.

For the TPA abstraction, empirical cumulative distribution functions were constructed for pH,[Cl!], [NO3

!], [CO32!] + [HCO3

!], and [SO42!] from the results of the thermodynamic simulations.

These cumulative distributions were used to construct stochastic samples of possible solutioncompositions by linear interpolation. Vectors {pH, [Cl!], [NO3

!], [CO32!] + [HCO3

!], [SO42!]} were

constructed for each brine type, and a single population of vectors was assembled preservingthe frequency of occurrence of each brine type. For example, a population of 10,000 vectorswas constructed by combining 800 calcium-chloride brine vectors, 2,400 neutral brine vectors,and 6,800 alkaline brine vectors. Figure 5-4 displays the distribution functions for pH and theconcentrations of anionic species derived from the total 10,000-vector population. TheSpearman rank correlation of the 10,000-vector set was computed [see, for example, Heltonand Davis (2002), Chapter 5, for a definition of the Spearman rank correlation matrix]. Relevantrank correlations are reported in Table 5-4. Rank correlation coefficients of absolute value lessthan 0.3 were ignored because their consideration is irrelevant to the assessment of theprobability of localized corrosion.

5-15

Figure 5-3. Box Plots Showing the Range in pH and Concentrations of Chloride, Nitrate,Total Carbonate, and Sulfate of the Three Brine Types That Could Form by Evaporation of

the Yucca Mountain Unsaturated Zone Porewaters Plotted in Figure 5-2. The Left andRight Edges of the Box Are the 25th and 75th Percentile Values, Respectively, and the Line

Inside the Box Is the Median Value. The 10th and 90th Percentile Values Are Shown asWhiskers and Outliers Are Represented by Black Dots.

5-16

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

5 6 7 8 9 10 11 12pH

Cum

ulat

ive

Dis

tribu

tion

Func

tion (a)

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 1 2 3 4 5[NO 3

-], mol/L

Cum

ulat

ive

Dis

tribu

tion

Func

tion (c)

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

1E-09 0.0000001 0.00001 0.001 0.1 10[CO 3

2-]+[HCO 3-], mol/L

Cum

ulat

ive

Dis

tribu

tion

Func

tion (d)

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

1E-06 0.00001 0.0001 0.001 0.01 0.1 1[SO42-], mol/L

Cum

ulat

ive

Dis

tribu

tion

Func

tion (e)

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

3 5 7 9 11[Cl-], mol/L

Cum

ulat

ive

Dis

tribu

tion

Func

tion (b)

Figure 5-4. Estimated Concentrations at 110 °C [230 °F] (a) pH, (b) Chloride,(c) Nitrate, (d) Carbonate-Bicarbonate, and (e) Sulfate

5-17

The dominant correlation in Table 5-4 is between pH and the amount of carbonate-bicarbonatein the system. This correlation is an indication that the highest carbonate-bicarbonateconcentrations are exhibited in alkaline brines. The negative correlation between pH and thechloride concentration arises because acidic calcium-chloride brines are highly concentrated inchloride. The negative correlation between chloride and carbonate-bicarbonate concentrationsis a consequence of the first two correlations.

Using interpolated Latin-Hypercube sampling, the Spearman rank correlation coefficients inTable 5-4, and the Iman-Conover procedure (Helton and Davis, 2002) to induce a desired rankcorrelation on a random sample, vectors {pH, [Cl!], [NO3

!], [CO32!] + [HCO3

!], [SO42!]} were

drawn to construct a stochastic population of 10,000. These vectors were used to compute thecorrosion potential, Ecorr, and the critical potential for localized corrosion, Ercrev, with theequations previously discussed. To consider variability in the corrosion potential, theparameters and A1 were sampled from triangular distributions defined in Table 5-2 and in thei a

o

paragraph following Table 5-1. The potential Ercrev was computed for mill-annealed and forthermally aged materials to assess the susceptibility of different parts of the waste package tolocalized corrosion. Resulting distributions of potentials are summarized in Figure 5-5.

The point at which complementary cumulative distribution curves, Figure 5-5 (b), cross thevertical axis can be used to estimate the frequency of localized corrosion. According to thisfigure, in 3 percent of the samples Ecorr is above Ercrev for mill-annealed material, and for up to26 percent of the samples Ecorr is above Ercrev for thermally aged material. This could mean that3 percent of the waste packages contacted by seepage at temperatures around 110 °C [230 °F]would be affected by localized corrosion on the mill-annealed body, and 26 percent on thewelded area. Waste packages will be protected from seepage by titanium alloy drip shields;thus, only a fraction of the waste packages might be directly exposed to seepage water. Bothfrequencies, 3 and 26 percent, are higher than reported from similar analysis in Mohanty, et al.(2004) (0.1 and 19 percent) because those frequencies are based on DOE distributions ofconcentrations of chemical species, which considered less frequent occurrence of neutral andcalcium-chloride brines.

The present analysis also differs from the analysis in Mohanty, et al. (2004) in the considerationof inhibiting oxyanions carbonate-bicarbonate and sulfate. The available sulfate concentrationis very low, Figure 5-4(e), to be an effective inhibitor. Therefore, this inhibitor can bedisregarded in a performance assessment model. On the other hand, carbonate andbicarbonate can be important inhibitors. For example, Figure 5-6 shows that if the presence ofcarbonate-bicarbonate in the system is disregarded, Ecorr will exceed Ercrev in 4 percent of thesamples in mill-annealed material and 53 percent of the samples in thermally aged material. Carbonate and bicarbonate are important inhibitors in alkaline brines (unimportant in the othertwo brine types because of its low concentration) that should be accounted for in a performanceassessment model.

5-18

Figure 5-5. (a) Distribution Functions for the Corrosion Potential and Critical Potentialfor Localized Corrosion Resulting from the Adopted Stochastic Sampling Approach

(b) Complementary Cumulative Distribution Function for the Difference Ecorr ! Ercrev. It isNoted that in 3 Percent of the Samples, Ecorr Exceeds Ercrev in Mill-Annealed Material and

26 Percent in Thermally Aged Material.

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

-400 -200 0 200 400 600 800 1000Potential, mVSHE

Cum

ulat

ive

Dis

tribu

tion

Func

tion

E corr

E rcrev (mill annealed)

E rcrev (thermally aged)

(a)

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

-1500 -1000 -500 0 500 1000 1500E corr - E rcrev , mV

Com

plem

enta

ry C

DF

Thermally aged

Mill annealed

3%

26%

(b)

5-19

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

-400 -200 0 200 400 600 800 1000Potential, mVSHE

Cum

ulat

ive

Dis

tribu

tion

Func

tion

E corr

E rcrev (mill annealed)

E rcrev (thermally aged)

(a)

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

-1500 -1000 -500 0 500 1000 1500E corr - E rcrev , mV

Com

plem

enta

ry C

DF

Thermally aged

Mill annealed

4%

53%

(b)

Figure 5-6. (a) Distribution Functions for the Corrosion Potential and Critical Potential forLocalized Corrosion Resulting from the Adopted Stochastic Sampling Approach

Disregarding the Presence of Carbonate-Bicarbonate in the System; (b) ComplementaryCumulative Distribution Function for the Difference Ecorr ! Ercrev. It Is Noted that in

4 Percent of the Samples, Ecorr Exceeds Ercrev in Mill-Annealed Material and 54 Percent inThermally Aged Material.

5-20

5.5 Conservative Assumptions in the Localized Corrosion Model

This section highlights relevant conservative assumptions in the localized corrosion model forperformance assessment of the waste package. In order for concentrated solutions to develop,seepage must directly impinge on waste packages during the thermal pulse. If drip shields arecapable of diverting seeping water during the thermal pulse, concentrated brines (seeFigure 5-3) will not form on the C22 outer container, and, consequently, localized corrosion willnot occur. The probability of localized corrosion displayed in Figure 5-5 is conditional on theprobability for the drip shield not to perform its diverting function against water seepage. Therefore, the probability for localized corrosion to initiate in the repository system is muchsmaller than indicated in Figure 5-5. Further performance assessments are necessary toevaluate the risk implications for the overall system.

In the performance assessment model, it is assumed that once initiated, localized corrosionprevails unless the environment changes. However, experimental data presented in Chapter 4(e.g., Figure 4-15) suggests that localized corrosion could repassivate, even in stronglyoxidizing conditions. This stifling mechanism may limit the extent of crevice attack. Localizedcorrosion experiments are performed in bulk solutions. In the repository, if concentrated brinesdevelop, localized corrosion may be activated under a limited supply of water. The stiflingmechanism is likely to be more effective in solutions of limited volume due to local saturationwith corrosion products.

The kind of localized corrosion that could affect Alloy 22 is crevice corrosion. Formation ofcrevices on the waste packages in the repository system may be limited to the outercontainer-pallet contact area. Additional crevice sites could form from direct contact betweenthe waste package and drip shield caused by drift degradation and rockfall. Therefore, theprobability for waste packages to be affected by localized corrosion is conditional on theprobability of formation of crevice sites. If crevice corrosion affects a waste package, thecorroded area will be limited and the degraded waste package could still offer protection againstwater seepage and radionuclide release. In the current version of the NRC-CNWRA TotalSystem Performance Assessment code, there is no consideration of the area affected bycrevice corrosion. Instead, if the localized corrosion front penetrates the waste packagethickness, seepage is assumed to contact waste forms and activate radionuclide releases.

The proposed model for localized corrosion allows evaluation of its importance to wastepackage degradation and radionuclide release in the context of a total system performanceassessment model. The localized corrosion model is conservative because it (i) ignores thepossibility of localized corrosion repassivation, (ii) ignores crevice site requirements, and(iii) disregards the geometry of the affected area by localized corrosion (i.e., the waste form isassumed exposed to water seepage if the waste package is penetrated by localized corrosion). Model refinement may be required if radionuclide releases due to localized corrosion are largelyoverestimated in the total system performance assessment model.

6-1

6 SUMMARY AND CONCLUSIONS

The proposed waste package design for the disposal of high level radioactive waste consists ofan outer container made of Alloy 22, a corrosion resistant nickel-chromium-molybdenum-tungsten alloy, surrounding an inner container made of Type 316 nuclear grade stainless steel. Degradation processes that are potentially important to the performance of the waste packageinclude dry-air oxidation, uniform corrosion, localized corrosion, microbially influenced corrosion,stress corrosion cracking, and mechanical disruption. The Center for Nuclear Waste RegulatoryAnalyses has conducted independent investigations and developed model abstractions for theTPA code to evaluate the effect of the waste package degradation processes on the overallperformance of the potential repository.

6.1 Dry-Air Oxidation

The consequence of dry-air oxidation over extended periods on the passive behavior of Alloy 22is analyzed via extrapolation of short term experiments at high temperatures. Dry-air oxidationkinetics and oxide morphology studies were conducted at 850 °C [1,562 °F] and 1,100 °C[2,012 °F] for up to 120 hours. At 1,100 °C [2,012 °F], the oxidation mechanism involvesexternal Cr2O3 scale formation and evaporation, and internal oxidation. At 850 °C [1,562 °F], theoxidation mechanism is controlled predominantly by outward diffusion of chromium ions to formexternal Cr2O3 scale. Internal oxidation, however, was also observed in localized areas. Short-term, high temperature tests indicate the oxidation rate is a function of temperature andthe thickness of the oxide layer.

6.2 Passive Dissolution and Oxide Film Analyses

Under environmental conditions where passivity is maintained, the uniform corrosion rate ofAlloy 22 is slow, and long waste package lifetimes are projected. The passive dissolution ratewas determined to be a function of temperature and environmental conditions. Based onpassive corrosion rate measurements, the activation energy for passive dissolution wasdetermined to be in the range of 33.5 to 49.6 kJ/mol [8.0 to 11.9 kcal/mol]. The passivedissolution rate decreased with time under constant conditions. The decrease in the passivedissolution rate may be attributed to improved corrosion resistance of the oxide film underpassive conditions. Passive corrosion rates in long-term exposures, were measured to be3.5 × 10!5 mm/yr [1.4 × 10!3 mpy] at 95 °C [203 °F]. The oxide films that formed on Alloy 22under passive and transpassive conditions were characterized by x-ray photoelectronspectroscopy combined with indepth profiling. Both the short- and long-term electrochemicallytreated specimens were analyzed to determine changes in the thickness and chemicalcomposition of the surface films as a function of environmental conditions and time. For short-term tests in 0.028 M NaCl, it was found that a thin oxide film of approximately 5.4 nm[2.12 × 10!4 mils] consisting primarily of Cr2O3 was formed in the passive region. The surfacefilm formed under transpassive conditions however, is thick, in the form of elongated patches. The surface layers formed on Alloy 22, both inside and outside the crevice sites, after long-termimmersion in 4 M NaCl simulated groundwater, had an outer silica deposit layer as a result ofglassware dissolution. As observed in the short-term passive specimen, a distinct Cr2O3passive film was also identified on the crevice sites. The presence of the chromium oxidepassive film is responsible for the passive behavior of the alloy.

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6.3 Localized Corrosion

Initiation of localized corrosion processes promoted by chloride ions, such as pitting or crevicecorrosion, may significantly decrease waste package lifetimes. Crevice corrosion of Alloy 22 ispossible in a limited range of conditions when the corrosion potential is greater than the crevicecorrosion repassivation potential. The corrosion potential of Alloy 22 was measured in a rangeof solution compositions and temperatures. Solution pH and temperature were the mostimportant environmental conditions that affect the corrosion potential. Increased temperaturesand pH values decreased the corrosion potential. The chloride concentration of the solutionsmay be an important parameter for the dissolved oxygen concentration, however, the corrosionpotential of Alloy 22 was found to be independent of chloride concentration in the range from0.028 to 4 M chloride. In acidic solutions the corrosion potential was near 300 mVSCE, whereasin near neutral and alkaline solutions, the corrosion potential was typically below 0 mVSCE at95 °C [203 °F].

The crevice corrosion repassivation potential was measured as a function of environmental andmetallurgical conditions. Fabrication processes such as welding increase the localizedcorrosion susceptibility of Alloy 22. Oxyanions such as nitrate, carbonate, bicarbonate, andsulfate can inhibit localized corrosion. The concentration ratio of these oxyanions to chloridenecessary to inhibit localized corrosion was determined as a function of metallurgical conditionand temperature. Although all the oxyanions are inhibitors of localized corrosion, it is notednitrate has greater solubility than to carbonate, bicarbonate, or sulfate. The increased solubilityof nitrate is significant, especially in concentrated chloride solutions.

A limited number of localized corrosion propagation tests was conducted. The propagation oflocalized corrosion can occur only when the corrosion potential remains above the repassivationpotential. Large drops in the corrosion potential, on the order of 200 to 300 mV, were measured as a result of localized corrosion initiation. The reduction in potential occurring as a result oflocalized corrosion initiation may promote repassivation and significantly limit both propagationrates and penetration depths.

A coupled multielectrode array sensor consisting of arrays of miniature electrodes that simulatedifferent areas on a single piece of metal was used to measure the localized corrosion ofAlloy 22 in a 0.1 M FeCl3 solution at temperatures between 18 and 90 °C [64 to 194 °F], and inNaCl solutions at room temperature. Results for Alloy 22 were compared to measurementsobtained using multielectrode array sensors with stainless steel and nickel base alloy sensingelements. The measured localized corrosion currents from the Alloy 22 multielectrode sensorsconfirmed that Alloy 22 is not subject to localized corrosion in NaCl solutions (up to 4 M inconcentration) or 0.1 M FeCl3 solutions at room temperature. The localized corrosion current ofAlloy 22 in the 0.1 M FeCl3 solution increased slightly when the temperature was raised from18 to 41 °C [64 to 106 °F], and increased further when the temperature was raised to 90 °C[194 °F], indicating the initiation of localized corrosion.

6.4 Passive Dissolution and Localized Corrosion Model

The equation for calculating corrosion potential in the TPA code was revised. The revisedequation and parameters yield values of corrosion potential as functions of temperature and pHthat are consistent with experimental data. In the performance assessment model it is assumedthat variability and uncertainty in the corrosion potential are the result of variability and

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uncertainty in the anodic current density. The revised equation is being implemented in arevised version of the TPA code.

A model is proposed to account for the combined effect of inhibitors, assuming a linearsuperposition effect. This model is being implemented in a revised version of the TPA code.

Distribution functions for pH, chloride, nitrate, carbonate-bicarbonate, and sulfate concentrationsfor a performance assessment abstraction are proposed. Spearman rank correlationcoefficients for correlated sampling in a performance assessment abstraction are proposed toaccount for the features exhibited by calcium-chloride, neutral, and alkaline brines and theirdifferent frequencies of occurrence. These distribution functions can be used to assess theprobability of localized corrosion if waste packages are contacted by seepage at hightemperatures {around 110 °C [230 °F] or less}. The TPA code already has the structure fordirect implementation of these distribution functions and suggested correlation coefficients.

The probability of localized corrosion if waste packages are directly contacted by seepage waterat high temperatures {around 110 °C [230 °F]} was estimated. It was estimated that 3 percentof the waste packages contacted by seepage could exhibit localized corrosion on themill-annealed surface and 26 percent could exhibit localized corrosion on welded areas.

Although 3 and 26 percent probabilities may appear high, the total probability of occurrence oflocalized corrosion depends on the probability of the waste packages being directly contactedby seepage. Drip shields are intended to reduce this probability. Also, it has been argued thatextrapolation to temperatures beyond 95 °C [203 °F] to compute the critical potential forlocalized corrosion of thermally aged material is conservative; thus, the 26-percent estimate isconservative. Total system performance assessment analyses can assess consequences oflocalized corrosion, accounting for drip shield protection and the limited welded area on thewaste package surface.

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