ONDRAF/NIRAS Repository Concept for Category C Wastes First full draft report

Transcription

1 Belgian agency for radioactive waste and enriched fissile materials Evolution of the Near-Field of the ONDRAF/NIRAS Repository Concept for Category C Wastes First full draft report Editor and principal author: Stephen Wickham, Galson Sciences Ltd NIROND-TR E April 2008

2

3 ONDRAF/NIRAS NIROND-TR report E GEOLOGICAL DISPOSAL PROGRAMME Evolution of the Near-Field of the ONDRAF/NIRAS Repository Concept for category B and C wastes First full draft report Editor and principal author: Stephen Wickham, Galson Sciences Ltd April 2008

4

5 This report was written on behalf of ONDRAF/NIRAS by: Editor and principal author: Stephen Wickham, Galson Sciences Ltd. It was reviewed by: Bennett, D., Galson, D., Hooker, P. (Galson Sciences Ltd) It was approved by: Dierckx A. (ONDRAF/NIRAS) Contact person and address: Stephen Wickham, Galson Sciences Ltd., Contact person at NIRAS/ONDRAF : [email protected] Credits : The content of this report has been discussed with many ONDRAF/NIRAS and SCK CEN experts, who have been asked to verify whether the content properly reflects their original contribution, and to confirm, in each field of expertise, that the complete set of relevant studies has been properly taken into account. The following specialists have been involved in this exercise: Eef Weetjens, Xavier Sillen and Ann Dierckx: end-user perspective Eef Weetjens, Xavier Sillen and Xiangling Li: THM behaviour Maarten Van Geet: chemical perturbations in the EDZ Lian Wang: chemical processes in concrete Frank Druyts and Bruno Kursten: corrosion aspects Hughes Van Humbeeck: general design issues Robert Gens: overall consistency. Robert Gens is also an editor of the report. The storyboard diagrams in Section 3 were drawn by Liz Harvey (Galson Sciences Ltd) ONDRAF/NIRAS Avenue des Arts BRUSSELS BELGIUM The data, results, conclusions and recommendations contained in this report are the property of ONDRAF/NIRAS. The present report may be quoted provided acknowledgement of the source. It is made available on the basis that it will not be used for commercial purposes. All commercial uses, including copying and re-publication, require prior written authorization of ONDRAF/NIRAS.

6 ONDRAF/NIRAS Document data sheet Document type Technical Report Publication date April 2008 Status of the document PD + IM IM MM First full draft report Series GEOLOGICAL DISPOSAL PROGRAMME Title Evolution of the Near-Field of the ONDRAF/NIRAS Repository Concept for Category C Wastes Author(s) of the primary document Stephan Wickham, Galson Sciences Ltd. Reference number N/A ONDRAF/NIRAS number NIROND TR E Author(s) of the integration module Order number Contract title IM identification number IM total no. of pages 0542d-1 N/A To be filled in for the management module only Author Distribution list Subcontractor identification GALSON N/A Date N/A MM identification number N/A MM total no. of pages N/A Analytical code 253-B61 FIDES code GCE11

7 Executive Summary This report provides a synthesis of the expected evolution of the engineered barrier system (EBS) in the ONDRAF/NIRAS concept for deep disposal of Category C waste (i.e., high-level radioactive waste (HLW) and spent fuel). It presents an up-to-date scientific view of the key processes involved and serves the following functions: It provides the scientific basis on which a reference model for the evaluation of repository safety can be constructed. It provides the necessary basis on which to construct scenarios for evaluation and demonstration of the safety case for deep disposal. It will be used in a formal exercise for checking completeness against a Features, Events and Processes list (FEP-list) specific to the Belgian disposal system. The reference model, development of scenarios for safety calculations and completeness checking will be reported in separate future projects. The report summarises the detailed phenomenological studies (currently available Level 5 Reports) within the ONDRAF/NIRAS Safety and Feasibility Case 1 (SFC-1) programme that are relevant to the EBS, and discusses the thermal, hydrological, mechanical, chemical, biological and radiation processes that may occur within the near-field. The report describes the key processes that may occur during repository construction and operation, and through the post-closure period. The report focuses on the disposal concept for HLW and spent fuel only. The expected evolution of the EBS concept for Long-Lived Low- and Intermediate-Level Waste (LILW-LL - Category B waste), and of the repository sealing system, will be presented in separate, additional reports. The report does not address spent fuel and HLW glass dissolution, radionuclide release and transport, or processes occurring within the damaged or disturbed zone of the Boom Clay, such as alkaline plume migration, because these processes will be reported in other SFC-1 documents. The report will be fully updated in future as further Level 5 reports and other literature become available. ONDRAF/NIRAS is considering the feasibility of disposing of LILW-LL, HLW and spent fuel in a deep geological repository excavated in the Boom Clay formation. The disposal concept for HLW and spent fuel is based on ensuring complete containment of the radioactivity during the thermal phase when temperatures in the repository will be significantly higher than those of the surrounding host rocks because of radioactive decay of the waste. The thermal phase corresponds to the period over which the performance of safety-critical disposal system components, primarily the retardation capability of the Boom Clay, could be compromised by the thermal transient, and will last for at least hundreds of years after emplacement of vitrified HLW, and possibly up to a few thousand years for spent fuel. The concept also seeks to avoid dissolution of spent fuel under radiolytically-induced oxidising conditions. A key component of the EBS design for HLW and spent fuel is the supercontainer. The supercontainer has been designed on the basis of the Contained Environment Concept, the intention of which is to establish and maintain a chemical environment around the overpack that is favourable to achieving the desired performance (containment of the radioactivity NIROND-TR E, April 2008 i

8 during the thermal phase). In the supercontainer, containment is achieved by placing the waste in a carbon steel overpack and surrounding the overpack with a Portland Cement (PC) concrete buffer. An outer stainless steel envelope may also be placed around the buffer and, if present, the envelope may be sealed or perforated to allow free entry and exit of water and gas. The sealed overpack encloses the canisters of HLW or spent fuel assemblies, and must contain and prevent the release of the radioactive waste for the duration of the thermal phase. Once fabricated, the supercontainer will be emplaced horizontally in tunnels excavated in the Boom Clay. The tunnels will be lined and supported with concrete wedge blocks (the tunnel liner). The void space between the supercontainer and the tunnel liner will be backfilled with a cementitious material (the backfill) before the tunnels are sealed with concrete or clay plugs. Immediately adjacent to the tunnel liner lies an excavation damaged zone (EDZ) of repository host rock that may have become damaged by excavation of the tunnels, and within which the changes to some of the host-rock parameter values are significant enough potentially to affect the performance of the disposal system. Beyond this lies an excavation disturbed zone (EdZ), where there are changes to some of the host-rock parameters relative to their nominal values but no significant impacts on long-term safety or repository performance. During the thermal phase, radiogenic heating will raise temperatures significantly throughout the disposal system near-field. Preliminary estimates of the thermal evolution of the system have been made based on generic parameters. Assuming a 60-year cooling time before supercontainer assembly, after 5 years most of the interior of the supercontainer is likely to be above 60 C, whether the waste is vitrified HLW or spent fuel. The peak temperature within the supercontainer, backfill and tunnel liner will be experienced within 5 years of supercontainer assembly and emplacement for vitrified HLW, and within 20 years for spent fuel. For vitrified HLW, the temperature close to the overpack will be >~80 C for a period of about 15 years, For spent fuel, the temperature close to the overpack will be >~90 C for a period of about 30 years. The supercontainer design and waste cooling time have been chosen so that the temperature at the outer surface of the overpack will not exceed 100 C for either waste type. The temperature at the outside edge of the buffer will rise to a maximum of ~65 C (after 5 years for HLW) and ~80 C (after 15 years for spent fuel). The temperature close to the backfill/liner/boom Clay interfaces will rise to a maximum of ~55 C (after 10 years for HLW) and ~70 C (after 15 years for spent fuel). These modelling results may be revised in future as a more complete set of input parameters becomes available. After emplacement of the backfill and sealing of the tunnels, the backfill is likely to exert a large suction potential and its saturation state will therefore increase rapidly. However, this will initially occur at the expense of the saturation of the Boom Clay close to the tunnel lining. Preliminary calculations indicate that about 1 year after backfilling, the concrete backfill materials will be ~99% saturated, but the degree of hydraulic saturation in the first decimetres of the surrounding Boom Clay will be lower than this. After about 2 years, however, both are likely to become completely saturated. In the scoping calculations the timescale is mainly dependent on the unsaturated behaviour of the EBS materials and Boom Clay, the porosity of the backfill, the hydraulic conductivity of the Boom Clay and the hydraulic gradient. However, the complexity of the resaturation process increases further when the interaction of the EBS NIROND-TR E, April 2008 ii

9 with the EDZ and host rock is taken into account, as demonstrated by the BACCHUS2 and RESEAL in-situ experiments. Calculations based on assumed material properties and the likely volume of unsaturated buffer materials, and assuming no stainless steel envelope, suggest that after hydrostatic pressures are attained in the backfill, the time required for the buffer to become saturated may be approximately 2 years. As for the backfill, a buffer material exerting a large suction can lead locally to temporary de-saturation of the Boom Clay close to the repository tunnels. The heat emanating from the waste may influence the distribution of water in the near-field. Close to the overpack, free water initially contained within the concrete buffer will be heated through contact with the overpack assembly. However, modelling suggests that the temperature increase in the buffer will induce little dehydration of hydrous cementitious phases, cause little porosity increase, generate little vapour, and is not likely to generate a dry zone adjacent to the overpack. Radiogenic heating and associated processes may potentially lead to various mechanical effects within the near-field, including expansion and contraction, solid volume changes, porosity changes, and cracking of the buffer. Scoping studies indicate that the tensile stresses expected as a result of the heat of concrete buffer hydration will be well below the tensile strength of the concrete. However, the main importance of the buffer is to condition the chemical composition of the supercontainer pore fluid, and this primary function will continue to be fulfilled regardless of any mechanical effects. After hydration, the solid phase assemblage within the buffer concrete will mostly comprise portlandite (Ca(OH) 2 ) and Calcium-Silicate-Hydrate gel. Initial buffer pore fluid ph values at 25 C may be well in excess of 12.4 due to the presence of alkali metal hydroxides. The initial redox conditions within the buffer concrete will be oxidising, but poorly poised. Similar initial conditions are likely within the backfill. Subsequently, a wide range of chemical effects may occur to alter the initial chemical conditions. The initially very high ph of the pore water in the cement barriers may decrease to slightly lower values (e.g., ph ~ 12.4 at 25 C, ph ~11 at 100 C) if the readily soluble alkali metal hydroxides are leached or diffuse away, and heating of the concrete buffer close to the overpack leads to the formation of high-temperature cement solid phases such as afwillite. However, the large mass of portlandite in the buffer will continue to buffer ph at levels >11 throughout the thermal phase and for a long time thereafter. At the irradiation rates expected within the buffer, radiolysis is not expected to have a significant impact on corrosion, but it may serve to prolong the duration of the aerobic phase. An experimental programme to investigate the possible impact of gamma radiolysis on corrosion of the carbon steel overpack at the expected irradiation rate within the supercontainer is currently underway. After the effects of radiolysis have ceased, anaerobic corrosion processes are likely to prevail at the overpack surface. Gas generation calculations indicate that hydrogen production due to gamma radiolysis of water will be small compared with hydrogen production due to anaerobic corrosion of steel components. Scoping calculations suggest that following buffer saturation, corrosion within the supercontainer will generate gas at a rate that is too fast for it to be removed by molecular NIROND-TR E, April 2008 iii

10 diffusion of dissolved gas, implying that a free gas phase will develop. Such gas would tend to migrate out of the supercontainer. An experimental programme to investigate further the generation of gas by corrosion of the overpack, and its potential consequences, is currently underway. If present, the external surface of the envelope may be exposed to various solutes, including chloride, carbonate, bicarbonate and various sulphur species in Boom Clay pore waters. Reaction with these solutes could lead to perforation of an initially intact envelope within a few years if aggressive ions (e.g., Cl -, S 2 O 2-3 ) reach the stainless steel/backfill interface before the oxygen in this region is consumed. However, it is likely that conditions at the surface of the envelope will rapidly become anaerobic, such that the corrosion potential falls below the pitting potential and, therefore, that perforation of an initially intact envelope will take much longer (> 100 years). Dark, hot, high-ph conditions, combined with low porosity will tend to suppress microbial activity, but the presence and possible persistence of microbes within the near-field cannot be fully discounted. Microbes could reduce oxidised sulphur present in the repository and the resulting reduced sulphur is capable of promoting pitting corrosion. Therefore, the potential for microbial activity within the buffer and the potential impact of reduced sulphur species on corrosion under high-ph conditions are being investigated further. The range of processes that will affect the chemistry of the near-field is extremely complex, and further more detailed modelling and experimental work is required in order to constrain the magnitude of the uncertainty associated with the various processes. This applies, in particular, to the likely concentration of aggressive species at the surface of the overpack. During the thermal phase, the interplay of processes within the supercontainer will be particularly complicated when the envelope is perforated or absent, with different fluids (water and gas) and different chemical species moving in opposite directions in response to hydraulic, temperature and chemical potential gradients. NIROND-TR E, April 2008 iv

11 Executive Summary i. Table of contents v. 1 Introduction Background Safety Functions and Functional Requirements of the EBS Design Concept Scope of the Report Structure of the Report 5 2 System Description Category C Disposal System Overview The Waste Form Vitrified HLW Spent Fuel The Supercontainer Overpack Buffer Envelope The Backfill The Tunnel Liner The Boom Clay The Excavation Damaged Zone (EDZ) and Excavation disturbed Zone (EdZ) 24 3 Overview of Expected Evolution Summary of Expected Evolution Expected Evolution with Envelope Absent or Initially Perforated Expected Evolution with Envelope Absent or Initially Perforated Expected Evolution with an Initially Intact Envelope Present Expected Evolution at Key Near-Field Locations Thermal Effects Hydraulic and Hydrothermal Effects 49 NIROND-TR E, April 2008 v

12 3.2.3 Mechanical Effects Chemical Effects Biological Effects Radiation Effects 53 4 Thermal Evolution Thermal Processes in the EBS Expected Thermal Evolution Vitrified HLW Spent Fuel Main Uncertainties 60 5 Hydraulic Evolution Hydraulic Processes Within the EBS Expected Hydraulic Evolution Saturation of the Backfill and Buffer Thermal Dehydration Effects Gas Effects Main Uncertainties 73 6 Mechanical Evolution Mechanical Processes Within the EBS Expected Mechanical Evolution Thermal Stresses Corrosion Effects Main Uncertainties 77 7 Chemical Evolution Chemical Processes in the EBS Initial Chemical Conditions Initial Chemical Conditions in the Boom Clay Initial Chemical Conditions in the Buffer Expected Evolution During Repository Construction and Operation Oxidation Effects Introduction of Microbes 86 NIROND-TR E, April 2008 vi

13 7.3.3 Effect of Radiolysis Effect of Elevated Temperatures Expected Chemical Evolution and Effects after Repository Closure The Return to Anaerobic Chemical Conditions Effect of Elevated Temperatures Interactions with Envelope Absent or after Envelope Perforation Main Uncertainties 93 8 Electrochemical Evolution of Metallic Barriers Electrochemical Processes Within the EBS Corrosion of the Envelope Corrosion of the Overpack Effect of Radiation on Anaerobic Corrosion of the Overpack Main Uncertainties Conclusions References 110 Input Data and Boundary Conditions to Support Modelling Studies of the Belgian EBS Design for HLW Disposal 117 A1 Introduction 117 A1.1 Background 117 A1.2 Structure of Appendix A 117 A2 Supercontainer Dimensions 117 A3 HLW Inventory and Radiolysis Calculations 121 A4 Concrete Buffer, Backfill and Wedge Blocks 125 A4.1 Cement 125 A4.2 Aggregate 125 A4.3 Concrete Proportioning 125 A4.4 Assumed Concrete Mix 125 A4.5 Superplasticiser 126 A4.6 Nominal physical properties of concrete buffer 126 A4.7 Water content at 60 C 127 A4.8 Physical properties of filler 128 NIROND-TR E, April 2008 vii

14 A4.9 Backfill 129 A4.10 Tunnel Liner 129 A5 Boundary Conditions for T-H Modelling 130 A5.1 Thermal conductivity of surrounding media 130 A5.2 Thermal power of vitrifed HLW 130 A6 Carbon Steel Overpack and Stainless Steel Envelope 132 A7 Boom Clay 135 A8 References 139 NIROND-TR E, April 2008 viii

15 1 Introduction 1.1 Background The Belgian radioactive waste management organisation, ONDRAF/NIRAS, is responsible for developing a deep disposal facility for Category B waste (i.e., low-level and intermediate-level long-lived radioactive waste (LILW-LL)) and Category C waste (i.e., vitrified high-level radioactive waste (HLW) and spent fuel). A primary aim of the ONDRAF/NIRAS programme for Category B and C wastes is to establish the feasibility of a deep disposal facility, without any presumption about precise repository location. Boom Clay, a poorly indurated argillaceous formation, is the reference medium for hosting such a disposal facility, but no official siting decision has yet been taken by the Belgian authorities. Most of the information related to the Boom Clay comes from the underground research laboratory (HADES) located beneath the Mol-Dessel region in north-east Belgium. This region also serves as a reference site for methodological research associated with the Category B and C waste disposal programme. There is an international consensus that the principal objective of any facility for the disposal of radioactive waste is to provide long-term safety by protecting humans and the environment from harm. This objective was stated by the International Commission on Radiological Protection (ICRP) as follows [35]: The principal objective of disposal of solid radioactive waste is the protection of current and future generations from the radiological consequences of waste produced by the current generation. The same objective was stated by the International Atomic Energy Agency [34] as follows: The objective of radioactive waste management is to deal with radioactive waste in a manner that protects human health and the environment, now and in the future, without imposing undue burdens on future generations. The commonly adopted management strategy to achieve this objective is to concentrate and contain the waste and to isolate it from the biosphere, and this strategy forms the starting point for the design of the Belgian deep disposal concept. ONDRAF/NIRAS has adopted a step-wise approach to disposal facility design. Various types of information and information sources, as identified in the ONDRAF/NIRAS safety strategy [54], inform the requirements of the disposal facility design, including: International principles of radioactive waste management and international recommendations and guidance. Boundary conditions, both external, as specified in national laws and regulations, and internal, as specified by ONDRAF/NIRAS early in the programme and unlikely to change. Disposal system development is then guided by various strategic choices, consistent with the principles and boundary conditions, and is carried out on two levels: NIROND-TR E, April

16 At a higher, more general level, the reference disposal concept is developed. The concept is developed to be robust to any reasonably foreseeable changes in principles, boundary conditions or knowledge. At a lower, more detailed level, specific design choices and decisions regarding repository implementation are made. The detailed design may be modified as the programme progresses through successive stages in order to adapt the repository to meet relevant requirements more successfully, take advantage of advances in science and engineering, and enable more robust safety and feasibility statements to be made. The general safety functions in the Belgian radioactive waste disposal concept are given in Table 1.1 [53]. Isolation (I-function) The objective of the I-function is to isolate the waste from man and the surface environment, in order to prevent direct access to the waste and to protect the disposal facility from the effects of surface processes. The I-function is divided into two sub-functions: I1 Reduce the likelihood and possible consequences of inadvertent human intrusion. I2 Create stable conditions for the waste and disposal system components in order to shield the repository from changes and disturbances at the surface or in the sub-surface. Engineered Containment (C-function) The objective of the C-function is to prevent for as long as possible any dispersion of contaminants outside the waste form and its primary containment. Delay and Attenuate Releases (R-function) The objective of the R-function is to retard and limit the eventual release of contaminants from the disposal system, after such time as the C-function is no longer present. Three sub-functions have been defined that contribute to both the delay and the attenuation of releases: R1 Limit the release from the waste form. Various physical and chemical mechanisms contribute to the resistance to leaching, such as slow dissolution and low solubility limits. R2 Limit water flow through the system. This function results in limited quantities of water infiltrating or moving through the system due to the presence of low permeability barriers (the Boom Clay and the tunnel and shaft seals). R3 Retard contaminant migration. Processes such as contaminant precipitation and sorption contribute to this sub-function, whereby contaminants released from the waste are mostly retained for an extended period. 2 NIROND-TR E, April 2008

17 Table General safety functions of the Belgian radioactive waste disposal concept, the barriers or components of the disposal system likely to provide or contribute to these functions, and the time frame over which they are expected to operate [53]. Safety function Sub-function Likely contributing component Time frame (y) Isolation (I) Reduce likelihood and consequences Institutional controls 10 2 of human intrusion (I1) Geological barrier 10 6 Create stable conditions for the disposal system (I2) Stable geological setting 10 6 Engineered containment (C) Prevent releases for as long as possible Engineered barriers Delay and attenuate the releases (R) Limit release from waste form (R1) Limit water flow through system (R2) Conditioned waste form Geological barrier, seals 10 4 (vitrified waste) (spent fuel) 10 6 Retard contaminant migration (R3) Geological barrier 10 6 Engineered containment must be provided for the duration of the thermal phase and the precise time frame may be modified based on future studies. The thermal phase corresponds to the period over which the performance of safety-critical disposal system components, primarily the retardation capability of the geological barrier, could be compromised by the thermal transient. Table 1.1 demonstrates that the Engineered Barrier System (EBS) is primarily responsible for providing one function, the engineered containment function (C). The main contributor to the R2 function will be the geological barrier, the Boom Clay, although the tunnel and shaft seals will probably also contribute to this function. ONDRAF/NIRAS is preparing safety cases to demonstrate the long-term safety of its deep disposal concept for Category B and C wastes. The safety cases will include the assessment basis, an overall description of the disposal system, and a description of the scientific and technical data and understanding relevant to the assessment of system safety and feasibility. This report supports the safety case for disposal of Category C wastes by describing the expected evolution of the near-field of the Category C disposal system, focusing in particular on the EBS. The EBS evolution report is not yet complete and will be further developed over the coming months and years as further information (e.g., from research studies) becomes available. A companion report focuses on the expected evolution of the EBS for Category B waste. NIROND-TR E, April

18 1.2 Safety Functions and Functional Requirements of the EBS Design Concept Based on the disposal system concept, a set of functional requirements has been established for each of the various sub-systems and components. The design of the EBS has been elaborated based on these functional requirements [30]. Feedback from focused research studies, scoping calculations and safety assessments is being used to further refine and improve the EBS design. The assignment of specific safety functions to the various sub-systems and components of the Category C disposal concept is a stepwise process as described in the ONDRAF/NIRAS safety strategy [54] [19]: Based on the strategic choices driving the disposal concept, the safety functions that are expected to be fulfilled by the system and its sub-systems and components are defined. Within the SFC-1 programme this was done in the period The assignment of safety functions is justified in the light of ongoing phenomenological assessment of the disposal system, sub-systems and components. Within the SFC-1 programme this will be carried out in the period 2006 to Finally, the assignment of safety functions will be confirmed based on confidence in the phenomenological understanding of the disposal system. 1.3 Scope of the Report This report presents an up-to-date scientific view of the processes involved in the expected evolution of the EBS design for Category C wastes, and serves the following functions: It provides the scientific basis on which a reference model for the evaluation of repository safety can be constructed. It provides the necessary basis on which to construct scenarios for evaluation and demonstration of the safety case for deep disposal. It will be used in a formal exercise for checking completeness against a FEP-list specific to the Belgian disposal system. The reference model, development of scenarios for safety calculations and completeness checking will be reported in separate future projects. The report focuses on the expected evolution of the EBS in the disposal concept for Category C waste involving the current reference design, namely a BSC-1 supercontainer comprising a carbon steel overpack and a concrete buffer, with or without an outer stainless steel envelope. The Boom Clay is the reference host formation for the disposal concept. The following subject areas are excluded: The report does not address the expected evolution of the EBS for Category B waste, and the expected evolution of the tunnel and shaft seals. These are described in companion reports. The report does not address spent fuel dissolution or HLW glass dissolution because these issues are addressed elsewhere in ONDRAF/NIRAS s safety case documentation 4 NIROND-TR E, April 2008

19 programme. However, other processes operating within the overpack, such as corrosion of primary waste containers, are included. The report does not address radionuclide release and transport processes. The report does not address processes mainly operating within the Boom Clay surrounding the repository, such as alkaline plume migration. The report refers to various scoping calculations that explore the behaviour of the supercontainer EBS design. Scoping calculations are a form of quantitative analysis, based on simplified assumptions, that is intended to capture quickly a general understanding of the behaviour of an engineered or natural system. These calculations provide a first-order examination of the sensitivity of the system to critical parameters and may involve numerical modelling or simple mathematical or algebraic reasoning. It is important to note that many of these calculations were performed early in the development of the disposal concept. Although the results of these calculations are indicative about the timeframe and magnitude of certain processes, the reader should be aware that they do not represent a systematic phenomenological analysis. 1.4 Structure of the Report The report is structured as follows: Section 2 provides a component-by-component description of the near-field and EBS in the ONDRAF/NIRAS disposal concept for Category C wastes. Section 3 provides an overview of the expected evolution of the near-field at the system level and for key locations within the near-field. Section 4 summarises the expected thermal evolution of the near-field and identifies the main uncertainties. Section 5 summarises the expected hydraulic evolution of the near-field and identifies the main uncertainties. Section 6 summarises the expected mechanical evolution of the near-field and identifies the main uncertainties. Section 7 summarises the expected chemical evolution of the near-field, including the evolution with and without an envelope, and identifies the main uncertainties. Radiological and biological EBS processes are also summarised in this section. Section 8 describes the electrochemical processes that are expected to affect the various metallic components within the near-field, and identifies the main uncertainties. Section 9 presents conclusions and summarises the principal remaining uncertainties. Appendix A provides a compilation of input data and boundary conditions relevant to the disposal concept for Category C waste. NIROND-TR E, April

20 6 NIROND-TR E, April 2008

21 2 System Description This section describes key features of the disposal system. Given the scope of this EBS evolution report, the focus is on the concept for disposal of Category C wastes. However, Category B wastes will be disposed of in the same repository as shown in Figures 2.1 and 2.2. The regions of the repository for the Category B and C wastes will be spatially separated (Figures 2.1 and 2.2). Figure A diagram providing a schematic view of the ONDRAF/NIRAS disposal concept for Category B and C wastes [52]. In the repository the regions for disposal of Category B and C wastes will be spatially separated NIROND-TR E, April

22 Figure A simplified view of the ONDRAF/NIRAS disposal concept for Category B and C wastes illustrating separation between the regions for the two waste types It is important to recognise that several aspects of the Category C disposal system are not yet finalised, in particular the presence or absence of a supercontainer envelope, the composition of the filler at the contact between buffer and overpack, and the exact position and placement of the supercontainer within the disposal tunnels. Dimensions of supercontainer components and compositions of EBS materials are also not yet finalised or optimised. The design and dataset provided is based on the latest information and assumptions. 2.1 Category C Disposal System Overview The principal elements of the EBS for the deep disposal of vitrified HLW and spent fuel are shown in Figures 2.3 and NIROND-TR E, April 2008

23 Figure Cross sections through a disposal tunnel containing the BSC-1 supercontainer with vitrified HLW. Dimensions are approximate and the position of the supercontainer in relation to the tunnel is shown schematically. In the disposal tunnel the supercontainer will rest on supports. The gap between the supercontainer and the tunnel liner will vary from ~20 cm near the roof to ~50 cm at the sides NIROND-TR E, April

24 Figure Schematic diagram showing a cross section through a disposal tunnel (gallery) containing the BSC-1 supercontainer with spent fuel. Dimensions are approximate and the position of the supercontainer in relation to the tunnel liner is shown schematically The principal component of the EBS shown in Figures 2.3 and 2.4 is the BSC-1 supercontainer (Figure 2.5). The supercontainer has been designed on the basis of the Contained Environment Concept, the intention of which is to establish and maintain a chemical environment around the overpack that is favourable to achieving containment of the radioactivity during the thermal phase [50] [51]. 10 NIROND-TR E, April 2008

25 6mm stainless steel envelope Figure Schematic diagram illustrating the BSC-1 supercontainer for vitrified HLW emplaced horizontally within tunnels excavated in the Boom Clay. The position of the supercontainer in relation to the tunnel liner is shown schematically The supercontainer comprises a carbon steel overpack and a Portland Cement (PC) concrete buffer (protective barrier around the overpack), with or without an outer stainless steel envelope. The overpack encloses the canisters of HLW or the spent fuel assemblies, and is designed to contain and prevent the release of the radioactive waste during the thermal phase. The duration of the thermal phase primarily represents the time over which the host rock experiences a significant thermal transient that could compromise its fulfilment of the R- function to delay and attenuate releases of radioactivity to the environment. Carbon steel has been chosen for the overpack because its corrosion behaviour in the highly alkaline environment that will be conditioned and buffered by the surrounding cement is well known, and because carbon steel is much less prone to localised corrosion processes than other steels. A PC concrete has been chosen for the buffer because this will provide a high-ph environment around the overpack that will be present throughout the thermal phase and for a considerable period thereafter. In this highly-alkaline chemical environment, the external surface of the overpack will be fully passivated and corrosion will be inhibited. The buffer also provides radiological shielding and thereby simplifies waste handling requirements. The buffer may also help to provide a low-hydraulic conductivity environment to slow the infiltration of externally-derived aggressive species to the overpack surface. It is assumed that the outer envelope of the supercontainer, if present, will be made of a low-carbon stainless steel with an enhanced Mo content (AISI 316L hmo) as suggested by the ONDRAF/NIRAS Corrosion Panel [50]. The high Mo content increases the resistance of the steel to localised corrosion. Once fabricated, the supercontainer will be emplaced horizontally in tunnels excavated in the Boom Clay (Figures 2.1, 2.2 and 2.5). In order to provide support and stabilise the open tunnels during the operational phase of the repository, the tunnels in the clay are lined with NIROND-TR E, April

26 concrete wedge blocks (the tunnel liner). Any fall-outs from the roof of the tunnels will be filled with cementitious grout to ensure that all void space is filled. The void space between the supercontainer and the tunnel liner will be backfilled with further cementitious materials (the backfill) before the tunnels are sealed. Two tunnel sealing options are currently considered by ONDRAF/NIRAS involving either concrete or clay seals, while the shaft seals will be clay. The design for the supports on which the supercontainer will be placed is yet to be chosen. Likewise, the exact composition and mechanism for emplacement of the different concrete phases, and the design of the envelope and supercontainer closure mechanism, are yet to be finalised. Immediately adjacent to the tunnel liner lies an excavation damaged zone (EDZ) of repository host rock that may have become damaged by excavation of the tunnels, and within which the changes to some of the host-rock parameter values are significant enough potentially to affect the performance of the disposal system. Beyond this lies an excavation disturbed zone (EdZ), where there are changes to some of the host-rock parameters relative to their nominal values but no significant impacts on long-term safety or repository performance. The EDZ is considered to lie within the repository near-field. The Boom Clay will act as a very effective barrier to the migration of radionuclides [49]. However, there are only limited data on the behaviour of radionuclides in the Boom Clay at elevated temperatures, and it can be difficult to quantify accurately radionuclide migration in regions with high and variable thermal gradients. The disposal concept for Category C wastes, therefore, envisages containment of the radioactivity during the thermal phase, so that radionuclides do not enter the Boom Clay when temperatures are elevated by the radioactive decay of the waste. The thermal phase corresponds to the period over which the performance of safety-critical disposal system components could be compromised by the thermal transient, and will be primarily constrained by the temperature range over which the Boom Clay is unable to fulfil reliably the R-function to delay and attenuate releases of radioactivity to the environment. The thermal phase will last for at least hundreds of years for vitrified waste, and possibly up to a few thousand years for spent fuel, after emplacement of wastes in the repository. It is also desirable for the overpack to remain intact throughout the period when dissolution of spent fuel could be enhanced by the existence of oxidising conditions, caused by alpharadiolysis. In this way the safety assessment can be based on an assumption that spent fuel dissolution takes place under reducing conditions. In this case the overpack should, ideally, remain intact for at least a few thousand years [56]. In corrosion studies to date, as a conservative measure, it has been assumed that the overpack for spent fuel must remain intact for 10,000 years. Table 2.1 summarises the dimensions and weight of the BSC-1 supercontainer [4]. For the purposes of post-closure safety assessment, the overpack, which surrounds the canisters of HLW or the spent fuel assemblies is assumed to provide the required radionuclide containment function (the C-function) during the thermal phase [49]. Table Dimensions of the BSC-1 supercontainer for HLW and the UOX and MOX spent fuel types ([4] Table 5). Dimensions are given in metres and correspond to designs with a 30 mm-thick 12 NIROND-TR E, April 2008

27 overpack and a normal density concrete buffer. The data for UOX correspond to the conservative reference design, but this design relates to the overpack construction technique and has no influence on the dimensions Vitrified HLW UOX MOX Outer diameter (m) Outer length (m) Overpack diameter (m) Weight (tonnes) (max) 31 (max) Number of canisters/fuel assemblies per supercontainer Recently, a further study has been carried out in which there has been a 12 mm reduction in the length given in Table 2.1 [68] because the study considered that the envelope was only present at the sides of the supercontainer, not at the ends. In future it is intended to standardise the supercontainer dimensions as much as possible in order to facilitate handling on the surface and to allow a uniform cross sectional geometry to be used for all Category C disposal tunnels. In this way the design and construction of the tunnel floor, the design of surface and underground handling equipment, and the emplacement of the backfill can all be standardised with important benefits for operational safety, quality control, efficient use of space and cost savings. 2.2 The Waste Form The Belgian deep disposal concept involves the disposal of three waste types, LILW-LL, vitrified HLW and spent fuel, each with its own characteristic forms and packaging. In the case of HLW and spent fuel (Category C), the vitrified waste canisters and spent fuel assemblies will be placed within the supercontainer overpack, with residual void space probably filled with inert media according to waste type. This report does not directly address spent fuel dissolution or HLW glass dissolution because these issues are to be addressed elsewhere in ONDRAF/NIRAS s safety case documentation programme. Therefore, only a short summary of the characteristics of Category C waste is provided Vitrified HLW Vitrified HLW is generated at the AREVA plant at Cap de la Hague in France. The waste is encapsulated in AISI 309 stainless steel canisters for transport and storage. After at least fifty years of storage, the HLW canisters will be emplaced within an overpack as the first stage of supercontainer assembly. Each overpack will have an external diameter of m (Table 2.1 and Figure 2.3) and will contain two vitrified waste canisters. The canisters will be positioned NIROND-TR E, April

28 symmetrically within the overpack using guides, and will be in end-to-end contact [4]. The space between canisters and overpack will probably be filled with glass frit, or an equivalent inert granular material or powder, and all air will be removed from the residual void space before the overpack is sealed by welding. The use of silica glass frit is considered advantageous because it will promote silica saturation and will tend to reduce the rate of dissolution of vitrified HLW Spent Fuel Category C waste includes two types of spent fuel, UOX and MOX. The reference fuels are the 14 UOX assembly from the Doel 4 and Tihange 3 reactors, and the 12 MOX assembly from the Doel 3 and Tihange 2 reactors. After reactor de-fuelling, spent fuel will undergo a period of storage before supercontainer assembly. The fuel assemblies will be placed in carbon steel boxes for ease of handling, and these may be filled with sand as a criticality control. The sand is intended to limit the amount of water that could potentially enter the boxes and surround the fuel. Alternatively, the boxes may be filled with an inert gas to protect against corrosion. The boxes will be supported within the overpack by a cast iron basket. All air will be removed from the overpack interior before it is sealed by welding. There are two overpack designs according to spent fuel type [4]: The overpack for UOX spent fuel will have an external diameter of m and will contain four separate boxes positioned in parallel (Table 2.1 and Figure 2.4). The overpack for MOX spent fuel will have an external diameter of m and will contain one box only (Table 2.1). 2.3 The Supercontainer The supercontainer will be fabricated in four stages (Figure 2.6): Stage 1 Stage 2 Stage 3 Stage 4 Emplacement of Phase 1 of the concrete buffer in the steel envelope (if present). Emplacement of the waste overpack. Emplacement of Phase 2 of the buffer in the small annulus remaining around the overpack. Formation of closure (Phase 3 concrete) and sealing of the supercontainer. Stages 2 to 4 will be carried out in a hot cell to ensure protection of workers from radiation. The possible composition of the buffer concrete is discussed by Belgatom [6]. The designs of the stainless steel envelope (if present) and lid have yet to be finalised. 14 NIROND-TR E, April 2008

29 Phase 1 concrete In hot cell Phase 3 concrete Phase 2 filler Stage 1 Stage 2 Stage 3 Stage 4 Figure The four stages of supercontainer fabrication. In the remainder of this section, we describe in more detail the composition, function, and outstanding design issues related to the components of the supercontainer. Further details of the assumed reference compositions, dimensions and physical properties of the supercontainer materials are provided in Appendix A. The supercontainer design is under development, and the report addresses the latest design at the time of writing Overpack The overpack will surround the vitrified HLW canisters or spent fuel assemblies and provide the main barrier to radionuclide release during the thermal phase. Based on recommendations of the ONDRAF/NIRAS corrosion panel [50], the overpack will be manufactured from carbon steel [4]. In the BSC-1 design illustrated in Figures 2.3, 2.4 and 2.5, the sides of the overpack have a nominal thickness of 0.03 m, while the end-pieces have a thickness of 0.06 m and are welded in place. This thickness is well in excess of the corrosion margin assumed to be required to ensure that the overpack remains intact during the thermal phase [4]. This corrosion margin is m for vitrified HLW and m for spent fuel see Table 2.2. Table Recommended thickness of the carbon steel overpack (after Belgatom [4]). The figures in brackets represent the thickness required for mechanical stability plus the corrosion margin Waste category Corrosion margin (mm) Total thickness (mm) Vitrified HLW 4 20 (16 + 4) Spent fuel (UOX) 6 26 (20 + 6) Spent fuel (MOX) 6 22 (16 + 6) NIROND-TR E, April

30 The reference design overpack is made from P-235 carbon steel (see Appendix A for physical and chemical data). The primary function of the overpack is to provide total containment of the radionuclides in the HLW and spent fuel throughout the thermal period. A low-alloy steel (2¼Cr1Mo-type) was originally recommended for the overpack by the Belgian expert review panel [50]. However, low-alloy steels are not necessarily adequately resistant to sulphide attack and are more difficult to weld, and carbon steel is now preferred on the basis of its more predictable behaviour. Initially, the overpack will be surrounded by oxygenated, highly-alkaline, PC-based concrete. This environment is particularly favourable for controlling and limiting corrosion because it will lead to passivation of the steel. Oxygen will be removed by reaction with the overpack steel, leading to anaerobic conditions and negative corrosion potentials close to the overpack that will probably be sustained over the long-term [41]. Under such anaerobic, high-ph conditions, the uniform corrosion potential will be lower than the pitting potential, and the long-term uniform corrosion rate is expected to be sufficiently low (~<0.1 µm per year - [56] [44] that the overpack will not fail by corrosion during the thermal period. It is possible that radiolysis of the buffer in the region close to the overpack may delay the onset of anaerobic conditions at the overpack surface. Corrosion processes are described in more detail in Section Buffer Phase 1 Concrete The Phase 1 concrete buffer will surround the overpack and may also be encased by the stainless steel envelope (Figures 2.3, 2.4, 2.5 and 2.6). The primary function of the Phase 1 concrete is to provide a high-ph environment at the surface of the overpack, at least during the thermal phase. High-pH conditions will fully passivate the carbon steel and keep corrosion rates low, thereby ensuring that the overpack will completely contain the waste during the thermal phase. The buffer also functions as a radiological shield so that dose rates at the outer surface of the supercontainer are low (maximum dose of 25 µsv per hour at a distance of 1 metre from the surface) and the containers can be handled without using additional shielding. This will be a significant benefit during the operational phase. A Phase 1 concrete composition has been defined as follows (see [28]), Table 2.3 and Appendix A for further details). Table 2.3 provides compositions for both a normal and a self compacting concrete, and ONDRAF/NIRAS is currently evaluating the performance of both types with a view to identifying a customised specification. The cement in the concrete is CEM I to European Standard BSENV 197-1:1992, with the additional restriction that the cement has not been interground with materials other than gypsum, contains no slag, and has a low SO 3 content, preferably not exceeding 2% SO 3. A further restriction is that the content of Ca 3 Al 2 O 6, as calculated from the Bogue formula [9], should not exceed 5%. The organic content of interground gypsum should be low, and the cement should have been ground without the use of organic grinding aids. The specific surface area (fineness) of the cement should not exceed cm 2 g -1. Coarse and fine limestone aggregates are specified to make up the supercontainer 16 NIROND-TR E, April 2008

31 concrete, the aggregates containing not more than 2% of magnesium, silicon and aluminium (as oxides). This particular cement specification avoids a number of potential problems: The potential for significant lowering of ph values is reduced by using cement only, i.e., by not permitting blending agents such as fly ash or silica fume to be used. Delayed ettringite formation and the consequent risk of expansion is eliminated by limiting the SO 3 content of cement. Formation of dense hydrogarnet and consequent increases in porosity and permeability are limited by imposing the chemical limits on the alumina (as represented by Ca3Al2O6) content. Alkali-aggregate reaction, arising from use of siliceous aggregates and resulting in expansion and cracking, is eliminated by using sand-grade high-purity limestone aggregate. The potential for organic complexing agents is reduced by eliminating organic grinding aids and limiting the organic content of the interground gypsum. Cement compositions under consideration by ONDRAF/NIRAS are based on the Belgian specification for CEM I high sulphate-resistant (HSR) and low-alkali (LA) cement. CEM I HSR (specification NBN B12-108) is characterised by: C3A (3CaO. Al2O3) 3.0% (in weight) Al2O3 5.0% (in weight) SO3 3.5% (in weight) CEM I LA (specification NBN B12-109) is characterised by: Na2Oeq (% in weight Na2O equivalent) 0.6% in weight (Na2Oeq = % Na2O % K2O) NIROND-TR E, April

32 Table Potential concrete buffer options currently under consideration by ONDRAF/NIRAS. The two options involve a normal PC concrete or a self-compacting concrete. The cement is an HSR and LA cement as specified above Component Content (kg m -3 ) Normal PC concrete Self-compacting concrete CEM I 42.5 N LA HSR LH Calcitec 2001 MS (ground CaCO 3 ) Limestone sand (0-4 mm) Calcareous aggregates (2-6 mm) Calcareous aggregates (6-14 mm) Calcareous aggregates (14-20 mm) Water : cement ratio Superplasticiser (Glenium ) 4.75% 14% At present, it is more likely that a normal PC concrete will be chosen for the buffer on account of its lower superplasticiser content. The superplasticiser Glenium will probably be used on account of the fact that it will not initiate or promote corrosion of steel. Should safety considerations require additional radiological shielding in excess of that provided by the concrete buffer in the current BSC-1 supercontainer design, consideration could be given to increasing the diameter of the supercontainer, or to replacing the limestone aggregate with a denser aggregate. Increasing the diameter of the supercontainer may not be practical as it could lead to increased fabrication and handling difficulties, and would require an increased tunnel diameter. It may be practical, however, to provide more shielding by replacing some of the limestone aggregate, probably the coarser fraction, with a denser aggregate such as hematite. The Phase 1 concrete buffer will be pre-cast and allowed to cure so that moisture is prevented from evaporating from it. Therefore the buffer will have a significant pore water content when the supercontainer is initially assembled. 18 NIROND-TR E, April 2008

33 Phase 2 Filler When the Phase 1 buffer is fabricated, its inner cylindrical cavity must be large enough to permit insertion of the overpack (see Figure 2.6). After overpack insertion, an annular gap will remain between buffer and overpack; the Phase 2 filler is the material used to fill this gap. The principal functions of the filler are to fill the void space at the overpack surface, thereby ensuring a good contact between the overpack, filler and buffer to contribute to minimising overpack corrosion and allowing heat transfer from the overpack. The composition of the Phase 2 filler has not yet been chosen. Two powder materials were initially proposed for the Phase 2 filler (see Appendix A): anhydrous lime (CaO), and slaked lime (portlandite - Ca(OH) 2 ). A cementitious grout is a further possibility that is also being considered. A cementitious grout would probably be easier to emplace than a powder within the narrow annulus between the overpack and Phase 1 concrete. At the time of emplacement, the overpack surface would be hot (~80-90 C) but there is experience in the oil industry of emplacement of grouts against similarly hot surfaces. On the other hand, portlandite and lime are both chemically compatible with the proposed composition of the Phase 1 concrete, and both would also provide a high-ph environment and minimise overpack corrosion, but these powders may be difficult to emplace, particularly in the hot-cell environment in which the supercontainer will be fabricated. Phase 3 Concrete End-Piece and Lid The main functions of the Phase 3 concrete end-piece are the same as for the Phase 1 concrete. The Phase 3 concrete end-piece will have the same composition as the Phase 1 concrete and it is important that there is a good bond between the Phase 1 and Phase 3 materials. The process for closure of the supercontainer has not been finalised, but various possible mechanisms have been examined. If the Phase 3 end-piece is pre-cast, sealing could involve bolting, or an additional layer of fresh cement that is cast directly onto the top of the end-piece. A possible design for a bolted end-piece is shown in Figure 2.7 [78]. NIROND-TR E, April

34 Figure Proposed design for a pre-cast end-piece that is bolted in place [68] Envelope It is as yet undecided whether the supercontainer will be encased in a cylindrical envelope. If it is present, the envelope will probably be made from 6 mm-thick stainless steel sheeting with a welded bottom and lid. The method of attaching the lid has not been finalised. The primary function of the envelope is to provide mechanical strength and thereby facilitate fabrication of the buffer and handling of the supercontainer. The envelope may, if sealed, also prevent water ingress from the Boom Clay for a time, and may facilitate monitoring during the operational period by allowing instrumentation to be attached to the external surface of the supercontainer. However, no reliance is placed on the envelope for ensuring long-term radiological safety, and it is possible that the envelope may be manufactured with vents, or not be present at all. It is assumed that if present, the outer supercontainer envelope will be manufactured from a low-carbon stainless steel with an enhanced Mo content (AISI 316L hmo) as suggested by the ONDRAF/NIRAS Corrosion Panel [50]. The physical and chemical properties for this material 20 NIROND-TR E, April 2008

35 are provided in Appendix A. The 316L high-mo stainless steel was recommended in preference to normal 316L stainless steel, because an enhanced content of 2.5 to 2.75 % Mo significantly reduces the propagation of localised corrosion, such as pitting or corrosion beneath deposits. The reduction in localised corrosion is attributed to the precipitation of molybdenum oxide (probably MoO 2 ). An important design issue concerns whether the envelope should be sealed, should have vents, or should be completely absent. A vented or absent envelope would allow the escape of gases generated by corrosion or radiolysis within the supercontainer, and would also allow external water to migrate into the concrete buffer pore space, leading to rapid saturation of both backfill and buffer after the tunnels are sealed. The expected evolution scenario described in this report considers both vented and unvented evolution. It is possible that a sealed envelope could remain intact for thousands of years if its electrochemical potential were to fall below the protection potential for chloride-induced pitting (~ -200mV she ) and reduced sulphur species were absent [41]. However, uncertainty over the timing of envelope perforation, and hence also of the time at which the buffer becomes saturated by externally-derived pore fluids, would be removed if it were initially vented. 2.4 The Backfill After emplacement of supercontainers within the repository disposal tunnels, void space will remain between the supercontainer and the tunnel wall. This void space will be filled with a cementitious backfill. The exact composition of the backfill and its emplacement mechanism has yet to be determined. It is likely to be based on a PC with a carbonate aggregate, similar to the buffer concrete, although other aggregates, such as quartz sand, have been tested. For example, an initial backfill development programme was carried out by Degussa (now BASF) involving the following components: The hydraulic binder involved a mixture of CEM I 52.5 N HSR LA and Carmeuse limestone powder (CaCO 3 ). The aggregate was graded river sand, 0-4 mm, washed and dried. The superplasticiser was Glenium - a polycarboxylate ether-based material. The backfill was sufficiently fluid to be pumpable at a water:cement ratio of between 1.34 and 1.45, and yielded a pore fluid having ph >13 and containing only trace chloride. The density was 2190 kg m -3 and the compressive strength 7.8 MPa after 28 day testing on mock-up samples at 12 MPa. 2.5 The Tunnel Liner In order to stabilise the open tunnels during the operational phase of the repository, and thereby minimise damage to the adjacent Boom Clay, the tunnels are lined with concrete wedge blocks. The composition of the wedge block concrete is given in Table 2.4. NIROND-TR E, April

36 Table The composition of the tunnel liner wedge block concrete currently under consideration by ONDRAF/NIRAS Component Content (kg m -3 ) PCCP CEM I 52.5N 430 PFA (Rugby Cement) mm graded crushed quartzite gravel 1055 Grade M washed quartzite grit sand 568 Content (litres m -3 ) Structuro *EMSAC 500 S 140 Water 30 (maximum) * EMSAC 500 S is an aqueous suspension of Elkem Microsilica manufactured by Elkem Materials. 2.6 The Boom Clay The Boom Clay is an Oligocene silty clay or argillaceous silt, with a high pyrite and glauconite content in the more silty layers. It is characterised by compositional banding on a scale of several tens of centimetres, reflecting mainly cyclical variations in grain size and in the carbonate and organic matter content. The more reactive mineralogical components of the Boom Clay are calcite, siderite and pyrite, and these control the redox potential and the concentrations of dissolved calcium and iron in the clay pore waters. The mineralogical composition of the Boom Clay is provided in Table 2.5 (see also Appendix A) and selected petrophysical and hydraulic parameters are given in Table 2.6. The Boom Clay forms a large component of the far-field of the disposal system and is, therefore, not the focus of this report. Nevertheless, by providing a reservoir of pore water that can migrate inward to the near-field, it provides an important boundary condition on the composition of near-field pore fluids. Boom Clay pore water is essentially a NaHCO 3 solution containing mg l -1 dissolved organic carbon [15]. The observed major cation concentrations can be explained by cation exchange and mineral dissolution and precipitation mechanisms. The maximum Eh is about -270 mv and is probably controlled by equilibrium between pyrite and siderite under the local geochemical conditions. Further information about Boom Clay pore fluid composition is provided in Section 5 and in Appendix A. 22 NIROND-TR E, April 2008

37 Table Mineralogical composition of Boom Clay. Values in % total dry wt. [49] [70] [15] [16] Clay minerals Illite Smectite + illite/smectite Kaolinite Chlorite Chlorite/smectite 30-60% 10-45% 10-30% 5-20% 0-5% 0-5% Quartz 15-60% K-Feldspars Albite Carbonates Calcite Siderite Dolomite Ankerite 1-10% 1-10% 1-5% 1-5% present present present Pyrite 1-5% Organic Carbon 1-5% Others: Glauconite, apatite, rutile, anatase, ilmenite, zircon, monazite, xenotime present present NIROND-TR E, April

38 Table Petrophysical and hydraulic parameters of Boom Clay (compiled from [2] [3] [31] [75] [49] Property Value Units Bulk density (sat) kg m -3 Average grain density 2650 kg m -3 Water content % dry wt Total porosity (from migration experiments) vol % In situ temperature 16 C Thermal conductivity 1.68 W m -1 K -1 Specific heat capacity 1400 J kg -1 K -1 Heat capacity 2.8 MJ m -2 K -1 Hydraulic conductivity Lab Field Vertical: m s -1 Horizontal: Vertical: m s -1 Horizontal: The Excavation Damaged Zone (EDZ) and Excavation disturbed Zone (EdZ) Adjacent to the disposal tunnels is a zone of the Boom Clay that has been disturbed or damaged by the process of tunnel construction. In order to distinguish between damage that is of concern to safety assessment because it affects the radionuclide transport properties of the rock, and other, less significant disturbance of the mechanical properties, the SELFRAC project proposed the following definitions [8]: The Excavation disturbed Zone (EdZ) was defined as a hydro-mechanically and geochemically modified zone without major changes in flow and transport properties and with no negative effects on long-term safety. The Excavation Damaged Zone (EDZ) was defined as a zone with hydro-mechanical and geochemical modifications that could induce significant changes in flow and transport properties, for example, a one or more orders of magnitude increase in flow permeability. Subsequently, in order to provide consistent definitions for use within SFC-1, ONDRAF/NIRAS has proposed the following definitions of the disturbed Zone (dz) and Damaged Zone (DZ) [73]. The dz and DZ refer to damage by all possible processes, not just those associated with excavation: 24 NIROND-TR E, April 2008

39 The DZ is a zone in the host rock where parameters relevant to the performance of the disposal system are modified to lie outside their nominal range. As the host rock properties change with time, the extent of the damaged zone may also change with time. The dz is a zone in the host rock where any parameter may be modified to lie outside its nominal range. For thermal and hydraulic perturbations the extent of the dz might be large and perhaps not limited by a specific spatial boundary. The EDZ and EdZ are defined in the same way as the DZ and dz but are related exclusively to the effects of excavation. A consequence of the new definitions is that the EDZ extends up to ~6 metres into the Boom Clay during the operational phase because the hydraulic conductivity is increased within this zone and this is a parameter that is relevant to safety [73]. Based on current knowledge, the original hydraulic conductivity is restored after resaturation of the EBS (see below). Note that the new definitions are based on observations at the HADES underground laboratory and may be further revised before they are officially adopted for use within the SFC-1 programme. Much of the knowledge of the expected properties of the EDZ and EdZ associated with a deep repository in the Boom Clay is based on excavation of the Connecting Gallery, an extension to the HADES underground research laboratory at Mol, and on in-situ experiments, performed therein, forming part of the SELFRAC project [8]. The initial magnitude of stresses in the Boom Clay adjacent to a newly constructed tunnel is high compared to the strength of the Boom Clay and it is therefore good practice to emplace the tunnel liner quickly in order to minimise the thickness of the EDZ. With the tunnel excavation techniques employed to create the Connecting Gallery, fracture openings were kept smaller than 1 mm at the excavation wall. The fractures form a conjugate set with a spacing of a few decimetres. The in-situ experiments have allowed quantification of the hydro-mechanical evolution of the EDZ and EdZ around the Connecting Gallery, including the effect of sealing processes [8]. The radial extent of the fracture zone around the tunnel is about 1 m, but there is a slight increase in hydraulic conductivity up to 6-8 m into the host-rock. The results confirm that the Boom Clay has important sealing properties. Two years after tunnel excavation, sealing had reduced the fracture zone from 1 m to less than 0.6 m around the tunnel. The hydraulic conductivity in the sealed zone and beyond in the host-rock remained lower than m s -1. It can therefore be concluded that the maximum hydraulic conductivity of the EDZ will be approximately one order of magnitude higher than that of intact Boom Clay, which is typically in the range 2 to m s -1 (Table 2.6). Some uncertainties remain concerning the evolution of the EDZ and EdZ during the postclosure phase, particularly the response of the EDZ to gas evolution and pressurisation within the EBS, temperature increase and chemical effects. Although the behaviour of the EDZ is not thought to pose a threat to long-term safety, further research will be undertaken by ONDRAF/NIRAS in order to reach a better phenomenological understanding of EDZ and EdZ processes. NIROND-TR E, April

40 26 NIROND-TR E, April 2008

41 3 Overview of Expected Evolution In this section, the information contained in the report is brought together to provide a brief overview of the expected evolution of the near-field of the disposal system for HLW and spent fuel. The evolution described focuses on the BSC-1 reference design with a reference inventory of vitrified HLW or spent fuel (Category C waste), characterised by: A supercontainer having the components illustrated in Figures 2.3 and 2.4 and dimensions listed in Table 2.1. The full dimensions of the supercontainer are provided in Appendix A. Vitrified HLW corresponding to two AREVA canisters per supercontainer. Spent fuel corresponding to four carbon steel boxes of UOX spent fuel, or to one carbon steel box of MOX spent fuel, per supercontainer. Details of the HLW radiological inventory are provided in Appendix A. Supercontainer and overpack dimensions relevant to HLW and spent fuel disposal are noted in Table 2.1. A carbon steel overpack having the composition specified in Appendix A. A PC concrete buffer that is allowed to cure such that moisture is prevented from evaporating. Nominal physical properties are provided in Appendix A. A filler of Portlandite powder or cementitious grout. A pre-cast concrete end-piece, sealed in place by bolting, or by a layer of fresh cement. A stainless steel envelope having the composition specified in Appendix A that is initially sealed and impermeable but may in time become perforated as a result of corrosion processes. ONDRAF/NIRAS has not yet decided if the envelope will be initially sealed, or will be constructed with openings. Where relevant, the alternative evolution in the case where the envelope is initially perforated or absent is also described. The following boundary conditions are assumed (see Appendix A): The vitrified HLW is emplaced in canisters at the time of vitrification. The canisters are stored in a surface storage facility for at least 60 years before overpack fabrication. Spent fuel assemblies are placed in carbon steel boxes and these would also be stored for ~60 years before overpack fabrication. Assembly of the supercontainer, involving emplacement of the HLW canisters or fuel assembly boxes inside the overpack, and insertion of the overpack within the concrete buffer, takes place after the storage period. Assembly will take place approximately years after vitrification or spent fuel retrieval. The time of assembly of the supercontainer is chosen such that the predicted maximum temperature of the overpack external surface or concrete buffer does not exceed 100 C. The supercontainer is emplaced in the repository immediately after fabrication and the tunnels are backfilled and sealed as soon as practicable after emplacement to minimise any chemical, mechanical or hydrological disturbance of the Boom Clay. During the operational phase, individual tunnels within the repository are likely to remain open for a NIROND-TR E, April

42 maximum of 10 years. The whole repository will probably remain operational for years. The operational period is considered by ONDRAF/NIRAS to include the period of construction, waste emplacement and closure of the repository. During the period of emplacement, it is assumed that each supercontainer is retrievable until the tunnel in which it is emplaced is backfilled and sealed. The description presented below addresses first the evolution of the entire system (Section 3.1), and second three different locations within the near-field (Section 3.2), namely within the buffer close to the overpack, within the buffer close to the surface of the supercontainer, and close to the backfill/liner/boom Clay interfaces (positions A, B and C in Figure 3.1). C B A Figure Schematic diagram showing a cross section through a disposal tunnel (gallery) containing the BSC-1 supercontainer with vitrified HLW and the positions A, B and C discussed in Section 3.2. Dimensions are in millimetres and are nominal. The position of the supercontainer in relation to the tunnel liner is shown schematically 28 NIROND-TR E, April 2008

43 3.1 Summary of Expected Evolution The figures and tables provided in Section 3.1 summarise the sequence of processes and events expected within the near-field; cross references are provided to the subsequent report sections giving more detailed information on the processes and events. The timings given in the figures and tables are indicative and are based on scoping calculations that have been carried out using generic parameters and EBS design assumptions that are not necessarily completely consistent with the design presented in Section 2. The expected evolution described has not yet been linked to safety assessment calculations but will ultimately be translated into a reference evolution scenario at which point the timeframes will become fixed. Figures 3.2 and 3.3 illustrate the main components of the EBS and near-field of the disposal system for vitrified HLW. In Sections and 3.1.2, significant stages in the evolution of the EBS are illustrated in a series of diagrams, which use the schematics in Figures 3.2 and 3.3 as a baseline. The key events and processes expected to occur at related time steps are tabulated on the facing pages (Tables 3.1 and 3.2). Section summarises the expected sequence of events and processes for a supercontainer where the stainless steel envelope is absent or is initially perforated. Section summarises the key events and processes as they are expected to develop for a supercontainer with an initially sealed envelope. Figure Schematic diagram showing a transverse cross section through a disposal tunnel containing a supercontainer with vitrified HLW. The main components of the EBS and near-field are labelled NIROND-TR E, April

44 Figure Schematic diagram showing a longitudinal cross section through a disposal tunnel containing a supercontainer with vitrified HLW. The main constituents of the EBS and near-field are labelled The timescale employed in this Section is normalised to closure of the tunnel at T0, following a ten year operating period (for a given tunnel). By T+10,000 years, there is expected to be negligible difference in the state of the EBS arising as a result of the initial state of the stainless steel envelope (sealed, perforated or absent). However, there are expected to be notable differences in the extent of certain processes, depending on whether the supercontainer contains HLW or spent fuel. Therefore, at this time step only, a supercontainer with HLW is illustrated in Section and a supercontainer with spent fuel is illustrated in Section These diagrams (Figures 3.10 and 3.11 in Section and Figures 3.18 and 3.19 in Section 3.1.2) are interchangeable in terms of the influence of initial state of the envelope. At all other time steps, a supercontainer with HLW is illustrated, but the events and processes described are expected to apply to both the HLW and spent fuel disposal concepts unless explicitly stated. The thermal power of vitrified HLW has been calculated for each time step shown in the figures below using the data reported by Put and Henrion [60] (see Appendix A, Section A5). The reported values assume a 60 year period of storage of HLW canisters in a surface storage facility prior to supercontainer fabrication, and immediate supercontainer emplacement at T-10 in the repository tunnels following fabrication. This approach is assumed to give the maximum possible thermal power of canisters emplaced in the repository. 30 NIROND-TR E, April 2008

45 This page intentionally left blank NIROND-TR E, April

46 3.1.1 Expected Evolution with Envelope Absent or Initially Perforated Figure Processes occurring in the supercontainer between fabrication and emplacement Figure Processes occurring in the repository tunnels during the operational phase prior to supercontainer emplacement and backfilling 32 NIROND-TR E, April 2008

47 Table The expected sequence of near-field processes and events focusing on a single disposal tunnel, for a supercontainer where the envelope is initially perforated or absent. The timescale is normalised to closure of the tunnel at T0 after a ten year operating period. All timings are indicative and do not correspond to any formalised safety assessment timeframe Time (approximate time in years relative to the time of supercontainer emplacement and backfilling) T-10, i.e., during repository excavation / construction T-10 to T0, i.e., during the operational phase up to supercontainer emplacement and backfilling. The operational period is considered by ONDRAF/NIRAS to include the period of construction, waste emplacement and closure of the repository. During the period of emplacement, it is assumed that each supercontainer is retrievable until the tunnel in which it is emplaced is backfilled and sealed. T0, supercontainer emplacement and backfilling (see also Figures 3.6 and 3.7) Key processes / events Fracturing of the Boom Clay and formation of a ~1 m-thick EDZ containing conjugate fracture zones. The EDZ has enhanced hydraulic conductivity due to decompression and variable stress (Section 2.6.1). Emplacement of concrete tunnel liner. Introduction of oxygen to the repository (Section 7.3.1). Oxidation of pyrite and organic matter in the EDZ and formation of sulphate and thiosulphate ions (Section 7.3.1). Precipitation of sulphate salts on the tunnel liner (Section 7.3.1). Possible reduction of sulphate and generation of reduced sulphur species by sulphate reducing bacteria (SRB). Suction-generated desaturation of the Boom Clay in the region close to the tunnel walls. Sealing of fractures in the EDZ (Section 7.3.1). Introduction of waste. Heating begins. Minor redistribution of sulphur species in the EDZ by diffusion (Section 7.3.1). Re-saturation of initially partially-saturated backfill begins (Section 5.2.1). Re-saturation of the concrete buffer will also begin at this time. Corrosion of the surface of the perforated envelope commences (if present) (Section 8.2). Partially-saturated conditions may develop in the Boom Clay near the repository (Section 5.2.1). NIROND-TR E, April

48 Figure Transverse cross section through the disposal tunnel showing key processes and events occurring during the peak thermal phase Figure Longitudinal cross section through the disposal tunnel showing key processes and events occurring during the peak thermal phase 34 NIROND-TR E, April 2008

49 Time (approximate time in years relative to the time of supercontainer emplacement and backfilling) T0 to T+5 Key processes / events Backfill re-saturation probably complete (Section 5.2.1). Re-saturation of the buffer will probably also complete within this period and corrosion of the overpack outer surface will commence (Section 8.3). Trapped air diffuses out of buffer as a dissolved phase. Radiolysis of buffer pore fluid is most intense. Temperature peaks at close to 100 C near to overpack surface. Generation of water vapour in response to heating buffer concrete. If the Boom Clay has become undersaturated near to the repository, saturated conditions will begin to return (Section 5.2.1). T+5 to T+20 Oxygen trapped within the EDZ, the repository tunnels and the supercontainer is consumed by corrosion, oxidation of pyrite and organic matter, and to a lesser extent by microbial activity. Passive layers on the outer surface of the overpack and on both surfaces of the envelope continue to evolve. The overpack uniform corrosion rate is typically < 0.1 µm per year. Localised corrosion of the perforated envelope may also occur (if present). Within the buffer, the radiation field will decay but the oxidising effect of radiolysis may still persist over this period. Chemical conditions within the EBS will become anaerobic but will not be well poised (Section 7.4.1). Temperatures within the buffer reach a maximum of 80 C to 100 C. Temperatures within the rest of the EBS reach at least 50 C to 60 C (Section 4.2). Thermal expansion of EBS components occurs. Re-mobilisation of sulphur species may occur through dissolution of precipitated sulphate salts (Section 7.4.2). The Boom Clay near the repository may reach temperatures as high as C, potentially liberating carbon dioxide and hydrogen sulphide gases (Sections 4.2 and 7.4.2). Hydration of the buffer concrete continues, but the development of crystalline cement phases is hindered by slow kinetics (Section 5.2.4). Complex coupled THM processes may occur. Thermal expansion may lead to cracking of EBS components and changes in porosity and permeability that may, in turn, influence heat transfer. Creep of the backfill may occur. NIROND-TR E, April

50 Figure Transverse cross section through the disposal tunnel showing key processes occurring approximately 300 years after emplacement Figure Longitudinal cross section through the disposal tunnel showing key processes occurring approximately 300 years after emplacement 36 NIROND-TR E, April 2008

51 Time (approximate time in years relative to the time of supercontainer emplacement and backfilling) Key processes / events T+20 to T+100 Hydration of the buffer concrete continues (Section 7.3.4). By about 30 years, all of the EBS is beginning to cool slowly (Section 4.2). T+100 to T+200 Various chemical species in the Boom Clay waters may react with the buffer and its pore waters, and start to migrate towards the overpack (Section 7.4.3). Calcite is expected to precipitate in the outer part of the backfill, and possibly the buffer, as a result of the chemical reaction between the HCO - 3 waters from the Boom Clay and the more alkaline waters of the concrete buffer. This may lead to porosity reduction (Section 7.4.3). T+100 to T+3,000 Overpack and uniform corrosion continues at a very slow rate, leading to the production of hydrogen gas (Section 8.2 and 8.3). Localised corrosion of the envelope may continue. Further reaction between chemical species in pore waters and buffer phases occurs leading to additional calcite precipitation (Section 7.4.3). NIROND-TR E, April

52 Figure Transverse cross section through the disposal tunnel showing key processes occurring approximately 10,000 years after emplacement Figure Longitudinal cross section through the disposal tunnel showing key processes occurring approximately 10,000 years after emplacement 38 NIROND-TR E, April 2008

53 Time (approximate time in years relative to the time of supercontainer emplacement and backfilling) T+3,000 T+3,000 to T+100,000 T+100,000 or longer Key processes / events The earliest possible time of overpack failure is uncertain and is being addressed through ongoing research. The time suggested here (T+3,000) is pessimistic and likely to be revised. If a uniform corrosion rate of <0.1 µm/year is assumed, the overpack lifetime could be as long as 40,000 years (HLW) to 60,000 years (spent fuel). Overpack failure is likely to occur either as a result of localised corrosion (e.g., stress corrosion cracking), or when general corrosion has reduced the overpack thickness to the point where it is unable to support the stresses exerted by the lithostatic overburden. Overpack failure will lead to pore water ingress, initiation of anaerobic corrosion of steel components within the overpack, and eventually the start of waste dissolution and the first possibility for outward radionuclide migration into the buffer (Section 8.3). After overpack failure and pore fluid ingress, further corrosion and production of hydrogen occurs, particularly in the case of spent fuel supercontainers, due to corrosion of carbon steel boxes and fuel rod baskets (see Figure not illustrated in Figure 3.11). The buffer maintains high-ph conditions (~ph 12.4) and acts to delay and attenuate any outward radionuclide migration (Section 7.4.3). The sorption and solubility characteristics of radionuclides within the EBS is constrained by the ph of the buffer pore fluid but no specific functional requirements are placed on the EBS at this stage. The ph of the buffer pore fluid will slowly decrease as the portlandite in the concrete becomes exhausted. Remaining radionuclides begin to dissolve in groundwater and are dispersed. NIROND-TR E, April

54 3.1.2 Expected Evolution with an Initially Intact Envelope Present Figure Processes occurring in the supercontainer between fabrication and emplacement Figure Processes occurring in the repository tunnels during the operational phase prior to supercontainer emplacement and backfilling 40 NIROND-TR E, April 2008

55 Table The expected sequence of near-field processes and events focusing on a single disposal tunnel, for a supercontainer having an envelope that is initially intact. The timescale is normalised to closure of the tunnel after a ten year operating period at T0. All timings are indicative and do not correspond to any formalised safety assessment timeframe Time (approximate time in years relative to the time of supercontainer emplacement and backfilling) T-10, i.e., during repository excavation / construction T-10 to T0, i.e., during the operational phase up to supercontainer emplacement and backfilling. The operational period is considered by ONDRAF/NIRAS to include the period of construction, waste emplacement and closure of the repository. During the period of emplacement, it is assumed that each supercontainer is retrievable until the tunnel in which it is emplaced is backfilled and sealed. T0, supercontainer emplacement and backfilling (see also Figures 3.14 and 3.15) Key processes / events Fracturing of the Boom Clay and formation of a ~1 m-thick EDZ containing conjugate fracture zones. The EDZ has enhanced hydraulic conductivity due to decompression and variable stress (Section 2.6.1). Emplacement of concrete tunnel liner. Introduction of oxygen to the repository (Section 7.3.1). Oxidation of pyrite and organic matter in the EDZ and formation of sulphate and thiosulphate ions (Section 7.3.1). Precipitation of sulphate salts on the tunnel liner (Section 7.3.1). Possible reduction of sulphate and generation of reduced sulphur species by sulphate reducing bacteria (SRB). Suction-generated desaturation of the Boom Clay in the region close to the tunnel walls. Sealing of fractures in the EDZ (Section 7.3.1). Introduction of waste. Heating begins. Minor redistribution of sulphur species in the EDZ by diffusion (Section 7.3.1). Re-saturation of initially partially-saturated backfill begins (Section 5.2.1). If the envelope is absent or initially perforated, re-saturation of the concrete buffer will also begin at this time. Corrosion of the outer surface of the envelope commences (Section 8.2). Partially-saturated conditions may develop in the Boom Clay near the repository (Section 5.2.1). NIROND-TR E, April

56 Figure Transverse cross section through the disposal tunnel showing key processes and events occurring during the peak thermal phase Figure Longitudinal cross section through the disposal tunnel showing key processes and events occurring during the peak thermal phase 42 NIROND-TR E, April 2008

57 Time (approximate time in years relative to the time of supercontainer emplacement and backfilling) T0 to T+5 Key processes / events Intact envelope prevents resaturation of the buffer during this period. Radiolysis of buffer concrete and pore fluid occurs, but is limited by the unsaturated conditions. Temperature peaks at close to 100 C near to overpack surface. Generation of water vapour in response to heating buffer concrete. If the Boom Clay has become undersaturated near to the repository, saturated conditions will begin to return (Section 5.2.1). T+5 to T+20 Oxygen trapped within the EDZ, the repository tunnels and the supercontainer is consumed by corrosion, oxidation of pyrite and organic matter, and to a lesser extent by microbial activity. Passive layer on the outer surface of the envelope continues to evolve. Very limited overpack corrosion occurs - the overpack uniform corrosion rate is typically < 0.1 µm per year. Localised corrosion of the envelope may occur, possibly leading to early perforation. Within the buffer, the radiation field will decay but the oxidising effect of radiolysis may still persist over this period. Chemical conditions within the EBS will become anaerobic but will not be well poised (Section 7.4.1). Temperatures within the buffer reach a maximum of 80 C to 100 C. Temperatures within the rest of the EBS reach at least 50 C to 60 C (Section 4.2). Thermal expansion of EBS components. Re-mobilisation of sulphur species may occur through dissolution of precipitated sulphate salts (Section 7.4.2). The Boom Clay near the repository may reach temperatures as high as C, potentially liberating carbon dioxide and hydrogen sulphide gases (Sections 4.2 and 7.4.2). Hydration of the buffer concrete continues, but the development of crystalline cement phases is hindered by slow kinetics (Section 5.2.4). Complex coupled THM processes may occur. Thermal expansion may lead to cracking of EBS components and changes in porosity and permeability that may, in turn, influence heat transfer. NIROND-TR E, April

58 Figure Transverse cross section through the disposal tunnel showing key processes occurring approximately 300 years after emplacement Figure Longitudinal cross section through the disposal tunnel showing key processes occurring approximately 300 years after emplacement 44 NIROND-TR E, April 2008

59 Time (approximate time in years relative to the time of supercontainer emplacement and backfilling) Key processes / events T+20 to T+100 Hydration of the buffer concrete continues (Section 7.3.4). By about 30 years, all of the EBS is beginning to cool slowly (Section 4.2). T+100 to T+200 An initially sealed envelope may, by this stage, have become perforated. In this case, waters from outside the supercontainer can enter the buffer and the buffer will rapidly become fully saturated (Section 5.2). Any trapped air will move outward by diffusion as a dissolved phase Various chemical species in the Boom Clay waters may react with the buffer and its pore waters, and start to migrate towards the overpack (Section 7.4.3). Calcite is expected to precipitate in the outer part of the backfill, and possibly the buffer, as a result of the chemical reaction between the HCO - 3 waters from the Boom Clay and the more alkaline waters of the concrete buffer. This may lead to porosity reduction (Section 7.4.3). T+100 to T+3,000 Overpack and uniform corrosion continues at a very slow rate, leading to the production of hydrogen gas (Section 8.2 and 8.3). Localised corrosion of the envelope may continue. Further reaction between chemical species in pore waters and buffer phases occurs leading to additional calcite precipitation (Section 7.4.3). NIROND-TR E, April

60 Figure Transverse cross section through the disposal tunnel showing key processes occurring approximately 10,000 years after emplacement Figure Longitudinal cross section through the disposal tunnel showing key processes occurring approximately 10,000 years after emplacement 46 NIROND-TR E, April 2008

61 Time (approximate time in years relative to the time of supercontainer emplacement and backfilling) T+3,000 T+3,000 to T+100,000 T+100,000 or longer Key processes / events The earliest possible time of overpack failure is uncertain and is being addressed through ongoing research. The time suggested here (T+3,000) is pessimistic and likely to be revised. If a uniform corrosion rate of <0.1 µm/year is assumed, the overpack lifetime could be as long as 40,000 years (HLW) to 60,000 years (spent fuel). Overpack failure is likely to occur either as a result of localised corrosion (e.g., stress corrosion cracking), or when general corrosion has reduced the overpack thickness to the point where it is unable to support the stresses exerted by the lithostatic overburden. Overpack failure will lead to pore water ingress, initiation of anaerobic corrosion of steel components within the overpack, and eventually the start of waste dissolution and the first possibility for outward radionuclide migration into the buffer (Section 8.3). After overpack failure and pore fluid ingress, further corrosion and production of hydrogen occurs, particularly in the case of spent fuel supercontainers, due to corrosion of carbon steel boxes and fuel rod baskets (see Figure 3.19). The buffer maintains high-ph conditions (~ph 12.4) and acts to delay and attenuate any outward radionuclide migration (Section 7.4.3). The sorption and solubility characteristics of radionuclides within the EBS is constrained by the ph of the buffer pore fluid but no specific functional requirements are placed on the EBS at this stage. The ph of the buffer pore fluid will slowly decrease as the portlandite in the concrete becomes exhausted. Remaining radionuclides begin to dissolve in groundwater and are dispersed. 3.2 Expected Evolution at Key Near-Field Locations This section summarises the range of key Thermal, Hydrologic, Mechanical, Chemical, Biological and Radiolytic (THMCBR) effects expected at the three locations identified in Figure 3.1 (Points A, B and C). Further detail on the events and processes causing these effects is then provided in Sections 4 through 7. NIROND-TR E, April

62 3.2.1 Thermal Effects Radiogenic heating caused by decay of waste radionuclides will raise temperatures throughout the near-field. The thermal evolution of the near-field differs depending on the length of the waste storage period, the time of supercontainer assembly and emplacement, and depending on whether the waste is spent fuel or HLW. ONDRAF/NIRAS has adopted a design constraint that requires the overpack surface temperature within the supercontainer to remain below 100 C. Based on preliminary thermal modelling it is likely that a cooling period of at least 60 years between vitrification of HLW, or reactor defuelling, and supercontainer assembly will be needed to satisfy this constraint [77]. Based on preliminary thermal modelling, a 60-year cooling period, and regardless of waste type, after 5 years most of the interior of the supercontainer is likely to be above 65 C and peak temperatures within the EBS are likely to occur in the period between 5 and 20 years after supercontainer assembly and emplacement. Maximum temperatures will be experienced at slightly later times with increasing distance from the waste. Because the buffer concrete will mostly be pre-cast and given time to cure before emplacement of the overpack, the initial temperature of the buffer prior to insertion of the overpack is likely to be close to ambient. Only the sealing layer for the Phase 3 concrete might be cast in situ. The relatively small volume of Phase 3 concrete means that any heat generated during its cure will not raise the temperature significantly throughout the surrounding concrete. It is believed that heat generated in the repository during hydration of the backfill will not cause a significant perturbation to the temperature field (R. Webers, personal communication). This assumption will be tested during the ESDRED large-scale backfill test when temperature increase due to backfill hydration will be measured. Although it is based on generic input parameters, the preliminary modelling of Weetjens and Sillen [77] allows a first estimate to be made of the likely thermal evolution at specific points within the EBS. Close to the overpack (Point A in Figure 3.1). After waste emplacement, temperatures close to the overpack are likely to be in the range C for about 15 years for disposal of HLW, and in the range C for about 30 years for disposal of spent fuel. After 50 years the temperature of the surface of the overpack will have fallen to approximately 60 C for HLW and to about 80 C for spent fuel, and the temperature of the buffer concrete will mostly be in the range C. Close to the supercontainer surface (Point B in Figure 3.1). The temperature close to the supercontainer surface will rise to a maximum of ~65 C (after 5 years for HLW) and ~80 C (after 15 years for spent fuel). Close to the backfill/tunnel liner/boom Clay interfaces (Point C in Figure 3.1). The temperature close to the backfill/liner/boom Clay interfaces will rise to a maximum of ~55 C (after 10 years for HLW) and ~70 C (after 15 years for spent fuel). 48 NIROND-TR E, April 2008

63 These modelling results will be revised in future as a more complete set of input parameters becomes available Hydraulic and Hydrothermal Effects The heat emanating from the waste may influence the distribution of water in the near-field. The evolution of hydraulic saturation in the near-field has been modelled by Weetjens et al. [78]. Hydraulic effects within the supercontainer will differ markedly depending upon whether the envelope is sealed, perforated or absent. Close to the overpack. Free water initially contained within the concrete buffer when the supercontainer is first assembled will be heated through contact with the overpack assembly. Some water may migrate outwards towards the edge of the buffer driven by the initially steep thermal gradient. Some water may also be absorbed into the Phase 2 filler if this is emplaced dry, or be released by the filler if a cementitious grout is used. The hydrothermal effects are likely to have a limited effect on buffer saturation and are not likely to generate a dry zone [58]. Preliminary scoping calculations to estimate the time required for the buffer to become saturated after the resaturation of the EDZ and attainment of hydrostatic pressures in the backfill, and in the absence of an envelope were undertaken by Weetjens et al. [78]. On the basis of the assumed material properties and unsaturated volume and a 25 cm-thick backfill, saturation of the buffer will take ~2 years. Once saturated, there will be no further driving force for advection of pore fluid through the buffer, save perhaps if gas overpressures develop, and any further transport will be by diffusion alone. Close to the supercontainer surface. Free pore water will migrate outwards away from the waste and towards the outer part of the concrete buffer in response to heating of the inner concrete. With a sealed envelope, in the period before envelope perforation these effects may lead to slightly higher saturations in the outer part of the buffer than in the inner part of the buffer. As noted above, however, if the envelope is absent the buffer will saturate throughout within ~2 years. Close to the backfill/tunnel liner/boom Clay interfaces. If it is not close to being saturated after its emplacement, the backfill would exert a significant suction potential in which case its saturation state would increase quite rapidly. However, saturation of the EBS is mainly controlled by the availability of pore water provided by the Boom Clay and will probably occur at the expense of the saturation of the Boom Clay close to the tunnel lining. Preliminary calculations based on a 25 cm-thick backfill indicate that about 1 year after backfilling, the concrete backfill materials will be ~99% saturated, but the degree of hydraulic saturation in the first decimetres of the surrounding Boom Clay will be lower than this [78]. After about 2 years, however, both will become completely saturated. Upon reaching hydraulic saturation, the pore water pressure is initially at atmospheric pressure and it takes a relatively long time, about 100 years, until this pressure is equilibrated with the in situ hydrostatic water pressure. The lowering of water pressure is primarily related to excavation and construction in the Boom Clay rather than by the EBS saturation process. The 100 year recovery time reflects the NIROND-TR E, April

64 extremely low hydraulic conductivity of the Boom Clay, which limits the rate at which water can be supplied from the surrounding region. In the future, hydraulic and hydrothermal effects will be further constrained by the results of in-situ experiments at the HADES underground laboratory Mechanical Effects The heat from the waste and other processes, such as mineralogical and porosity changes, may potentially lead to various mechanical effects within the near-field, including expansion and contraction and cracking of the buffer. The tensile stresses expected as a result of the heat of concrete buffer hydration are well below the tensile strength of the concrete [6]. However, the heat generated by the waste is the prime concern as a possible cause of cracking. Close to the overpack. Rapid heating of the concrete buffer close to the overpack may cause expansion cracking and may lead to the formation of higher-temperature cement solid phases with different chemical and physical properties. Close to the supercontainer surface. If the envelope is absent, or if it becomes perforated, waters from outside the supercontainer may enter and react with the buffer. Calcite will - precipitate in the outer part of the buffer as a result of the chemical reaction between the HCO 3 waters from the Boom Clay and the more alkaline waters of the concrete buffer, and this may lead to porosity reduction [28] [76]. Close to the backfill/tunnel liner/boom Clay interfaces. Swelling and creep of the Boom Clay will gradually seal any excavation-induced fractures and will subsequently raise the pressures back toward the ambient lithostatic pressure. This latter effect may cause some mechanical deformation of the liner and backfill, which may themselves also experience creep. Compared with such deformations, the mechanical effects of chemical precipitation reactions (carbonation, sulphate precipitation) are likely to be of lesser significance. Later in the evolution of the system, gas generation may potentially to cause local transient deformation of the Boom Clay (see Section 3.2.4) Chemical Effects After initial hydration, the solid phase assemblage within the buffer concrete will mostly comprise portlandite (Ca(OH) 2 ) and Calcium-Silicate-Hydrate (CSH) gel. Initial buffer pore fluid ph values at 25 C may be well in excess of 12.4 due to the presence of alkali metal hydroxides. The initial redox conditions within the buffer concrete will be oxidising, but poorly poised. Similar initial conditions are expected within the backfill. However, a wide range of chemical effects may occur subsequently to alter the initial chemical conditions. Close to the overpack. The initially very high ph of the pore waters in the cement barriers may decrease to slightly lower values (e.g., ph ~12.5) if the readily soluble alkali metal hydroxides are leached or diffuse away and heating of the concrete buffer close to the overpack 50 NIROND-TR E, April 2008

65 leads to the formation of high-temperature cement solid phases such as afwillite. However, the large mass of portlandite in the buffer will continue to buffer ph at levels >11 throughout the thermal phase and for a long time thereafter. Radiolysis may initially generate oxidising radicals [11] [12]. Calculations based on the expected average maximum irradiation rate within the buffer (~10-4 Gy s -1 ), and on the expected maximum initial irradiation rate at the buffer overpack interface (~ Gy s -1 ), suggest that radiolysis may influence the duration of oxidising conditions by generating oxidising radicals over a period of 300 years [11] [12]. However, experimental observations suggest that such irradiation rates will generally have a minor, short-lived impact on corrosion ([66]; [29]). Modelling of overpack corrosion in the absence of radiolysis has indicated that redox conditions are likely to become anaerobic within at most a few years, and the influence of oxygen on corrosion processes over longer timescales is likely to be negligible [41]. However, preliminary calculations by MacDonald et al. (in progress) indicate that the cathodic reactions under irradiation might be controlled by O 2 or H 2 O 2, rather than H 2 O, and further studies are underway to confirm these scoping calculations. Good quality concrete is one of the most favourable environments for preventing corrosion of steel. The high ph and initially oxidising conditions will lead to the development of a passive film on the carbon steel overpack. Oxygen will be removed rapidly by reaction with the overpack steel, leading to anaerobic conditions and negative corrosion potentials that will be sustained over the long-term. Under such anaerobic, high-ph conditions, the long-term uniform corrosion rate is likely to be low (~<0.1 µm per year - [65] [44]. The risk of localised corrosion is negligible because the corrosion potential of the overpack (and the stainless steel envelope, if present) is predicted to be much lower than the pitting potential [41]. Low chloride and sulphide content of the buffer concrete are also favourable for preservation of the passive film on the overpack, assuming that this low content can be maintained and that significant concentrations of aggressive species are not able to reach the overpack. This assumption will depend, to some extent, on the level of SRB activity in the EDZ and tunnel liner. H 2 and other gases may be produced through radiolysis of water, primarily free pore water. However, the pressure rise due to radiolysis is likely to be small, probably 0.1 to 0.3 MPa over the first 100 years [11]. After the onset of anaerobic conditions, H 2 will also be generated by anaerobic corrosion of the overpack and this is likely to generate more significant quantities of gas [41]. However, based on recently obtained electrochemical measurement of the passive current density of carbon steel in a saturated Ca(OH) 2 solution (~10-8 A cm -2 ), the amount of hydrogen generated is not expected to pose a threat to the integrity of a sealed stainless steel envelope [42]. Weetjens et al. [78] have analysed gas transport within the supercontainer buffer. Their calculations also demonstrate that the hydrogen production due to gamma radiolysis is likely to be quite small compared with hydrogen production due to anaerobic corrosion of the steel EBS components. Their scoping calculations suggest that gas could be generated at a rate that is too fast for it to be removed by molecular diffusion of dissolved gas and therefore a free gas phase might develop. In the model, the highest calculated pressure is 3.4 MPa and this occurs close to the overpack after ~100 years. However, the calculations were based on a relatively high NIROND-TR E, April

66 assumed corrosion rate of 1 µm per year for the first 100 years, followed by 0.1 µm per year thereafter, and may therefore overestimate the gas pressure generated. Close to the supercontainer surface. The abundance of high-temperature solid cement phases within the buffer may be less than that close to the overpack, but high-ph conditions will still lead to the development of a high-quality passive film on the surface of the stainless steel envelope, if present. Under these conditions, the envelope corrosion rate at its internal surface is estimated to be significantly less than 0.01 µm per year [65]. Even at these low corrosion rates, anaerobic conditions are expected to develop rapidly [41]. Again, the low chloride and sulphide content of the buffer concrete are favourable for preservation of the passive film, assuming that this content can be maintained. Close to the supercontainer surface, the irradiation rate will be too low for radiolysis to be significant. A negligible quantity of gas is likely to be generated by internal corrosion of the stainless steel envelope under anaerobic conditions. If the envelope is absent, or after envelope perforation, calcite will precipitate in the outer part of the buffer as a result of the chemical reaction between the HCO - 3 waters from the Boom Clay and the more alkaline waters of the concrete buffer [76]. Close to the backfill/tunnel liner/boom Clay interfaces. The external surface of the supercontainer may be exposed to various solutes, including chloride, carbonate, bicarbonate and reduced sulphur species in Boom Clay pore waters. Reaction with these solutes could lead to corrosion and perforation of a sealed envelope (if present) within a few years if aggressive ions (e.g., Cl -, S 2 O 2-3 ) reach the stainless steel/backfill interface before the oxygen in this region is consumed. However, the backfill will saturate relatively quickly (>95% within a few weeks - Weetjens et al. [78], and anaerobic conditions are likely to prevail at the supercontainer outer surface after a brief transition period, as is expected at the envelope inner surface, if present [41]. The combined effect will be effectively to protect the envelope from significant corrosion under aerobic conditions by aggressive ionic species present in Boom Clay pore fluid. It is therefore likely that soon after closure, anaerobic conditions will prevail at the supercontainer outer surface. Under such conditions, corrosion of any envelope that may be present is likely to be slow and the time to perforation could exceed 1,000 years. However, if sufficient concentration of aggressive reduced sulphur species, such as sulphide or polysulphide, are present, corrosion (and envelope perforation) could be more rapid. ONDRAF/NIRAS is evaluating further the potential role of such species Biological Effects Dark, hot, high-ph conditions will tend to suppress microbial activity, but the presence and possible persistence of microbes cannot be fully discounted [55]. It is possible that superplasticisers, as well as dissolved organic carbon in Boom Clay pore water, could form a nutrient source for microbes within the EBS. 52 NIROND-TR E, April 2008

67 Close to the overpack. Close to the overpack, conditions are most unfavourable for microbial activity owing to the high temperatures and high ph. Close to the supercontainer surface. Any surviving microbes will be largely immobile owing to the small size of the pores in the buffer concrete as compared to the typical size of most microbes. Close to the backfill/tunnel liner/boom Clay interfaces. Conditions close to the backfill/liner/boom Clay interfaces, while still harsh, may be the most favourable for microbial growth and activity. It will therefore be important to ensure that any void space in this region, such as created by roof fallouts, or spaces between the wedge blocks and the Boom Clay, is filled with cementitious grout. Although void space will be minimised, after saturation the supply of nutrients will be by diffusion, which may limit microbial activity Radiation Effects The average gamma irradiation rate in the buffer after supercontainer fabrication will be about Gy s -1 or 0.6 Gy hour -1 [11]. Close to the overpack, the irradiation rate will be higher (initially approximately 25 Gy hour -1 ), but still much lower than at the surface of the canisters because of shielding by the overpack. Radiolysis will influence redox conditions and will generate hydrogen, oxygen, and other oxidising radicals whenever pore water is present. The effects of these oxidising radicals could last for a few hundred years in the case of vitrified HLW. Close to the supercontainer surface and close to the backfill/tunnel liner/boom Clay interfaces radiation effects will be negligible. NIROND-TR E, April

68 54 NIROND-TR E, April 2008

69 4 Thermal Evolution 4.1 Thermal Processes in the EBS Prior to repository construction, the temperature of the rocks at the repository location will depend on the prevailing geothermal gradient. During repository construction but before waste emplacement, the temperature within the excavations may be influenced by any tunnel ventilation used, but the temperature of the repository host rocks themselves is expected to be little different from those prior to construction. The thermal evolution experienced by the EBS will depend on the heat sources present and the processes by which thermal gradients within the system are established and subsequently relax. The principal heat source in the system is provided by the waste, which will liberate heat as it decays, but any exothermic chemical reactions that occur (e.g., cement hydration reactions) will also contribute to the heat budget. The predominance and nature of the radioactive decay process means that the thermal evolution of the EBS can, in general terms, be described as an initial period of rapid heating, followed by a slower exponential cooling, as the system returns to a state close to that of the pre-excavation conditions. The heating effect of cement hydration reactions in the repository is expected to be relatively minor because much of the concrete to be emplaced in the repository is contained in the buffer and the tunnel linings, which will have been formed, cured and cooled at the surface before emplacement (although it is acknowledged that some slow hydration reactions may continue after supercontainer emplacement). Depending on its composition and means of emplacement, exothermic reactions within the cementitious backfill may also contribute as a heat source in the first few years after closure. In detail, the precise thermal evolution experienced will differ from location to location. Higher absolute peak temperatures are expected closer to the wastes, and the time at which peak temperatures are achieved will be later at greater distances from the heat source. The thermal evolution will also be constrained by a potentially complex set of thermo-hydromechanical-chemical processes that will influence rates of heat transfer. Thermal interaction between adjacent tunnels will not occur for tens of years and will therefore post-date peak near-field temperatures. However, such interactions will have a negligible effect on peak temperatures [77]. Thermal processes affecting heat transfer include conduction, convection and radiation. Heat transfer is expected to occur mainly by conduction but convection may also play a significant role. The design of the EBS is such that there will be very few or no voids, and this means that radiative heat transfer will be minor. For example, the Phase 2 filler is specifically designed to eliminate voids and ensure good contact between the overpack and the Phase 1 buffer concrete. Similarly, cementitious grout will be used to minimise any void space in the region between the tunnel liner and the Boom Clay, and the backfill will occupy the space between the overpack and the tunnel liner. NIROND-TR E, April

70 The significance of heat transfer by conduction will depend to a great extent on the thermal conductivities of the EBS materials and the surrounding rocks. Modelling studies have demonstrated that, if a sealed envelope is included in the design, then significant water flow and convective heat transfer would not be expected within the supercontainer during the initial period after repository tunnel closure, when significantly elevated temperatures will occur [58] [59]. Outside the supercontainer, however, heat transfer by convection may occur in the backfill and in the tunnel liner during hydrological resaturation. If the envelope is absent or the design includes perforations, repository resaturation up to the surface of the overpack will be a fairly rapid process and the buffer will become fully saturated from an early stage. This would greatly simplify the analysis of conditions within the EBS and heat and mass transfer within the buffer. However, even if the envelope is absent or perforated, convection is still unlikely to play a significant role in heat transfer within the supercontainer due to the low permeability of the concrete buffer. Thermal expansion owing to the presence of time-varying thermal gradients within the EBS could lead to changes in stress and potentially to cracking of EBS components. These and other changes in the physical characteristics of EBS components over time (e.g., changes in porosity and permeability) could, in turn, feedback to influence heat transfer and thermal evolution. Assessing the thermal effects of waste emplacement is an important aspect of the repository design process; particularly, given the uncertainties associated with evaluating the coupled thermo-hydro-mechanical-chemical processes that may occur in the early stages after waste emplacement. However, ONDRAF/NIRAS is following an iterative approach to repository design and assessment, which allows the disposal system to be designed in such a way that the thermal evolution is acceptable. For example, the thermal evolution of the repository can be managed by allowing an appropriate waste cooling time before supercontainer assembly and by using appropriate spacing between the wastes. 4.2 Expected Thermal Evolution Preliminary studies of the thermal evolution of the near-field have been made by Weetjens and Sillen [77] and Weetjens et al. [78]. These preliminary studies were based on generic values of thermal parameters and will be updated in the future based on a revised set of more realistic input values. The preliminary studies involved two-dimensional axi-symmetric conductive heat flow calculations, using the code PORFLOW [61]. The importance of other heat transport mechanisms, such as convection or radiation during the operational phase, was evaluated but the influence on the subsequent thermal evolution within and around the supercontainer was found to be minimal. The objectives of the study of thermal evolution were: To determine the maximum temperatures to which the overpack and buffer materials might be exposed. 56 NIROND-TR E, April 2008

71 To characterise the temperature field around the disposal tunnels in order to constrain thermal boundary conditions for geochemical modelling exercises. To investigate the main influences on the maximum temperatures experienced in the nearfield, so that the design choices affecting thermal behaviour can be identified. The calculations were made for cases involving vitrified HLW and UOX spent fuel. The calculations assumed realistic values for the thermal conductivity of buffer (1.7 W/mK), envelope, backfill, tunnel liner and Boom Clay, and an exaggerated, high conductivity for the overpack and its interior, in order to ensure a uniform internal temperature distribution. Weetjens and Sillen [77] presented calculations assuming the use of normal concrete and, as an alternative, heavy concrete containing hematite aggregate for the buffer. However, because heavy concrete is not presently considered by ONDRAF/NIRAS as a reference buffer material, the results for heavy concrete are not discussed here. The only heat source term in the scoping calculations was heat generation by radioactive decay. The thermal output of vitrified HLW was based on the empirical relationship suggested by Put and Henrion [60]. The thermal output of spent fuel was based on calculations with the ORIGEN code using data supplied by Belgonucléaire [77]. For spent fuel, a relatively hot type of UOX assembly was chosen (burnup of 55GWd/tHM) to be conservative. EBS material properties were provided by ONDRAF/NIRAS and are believed to be well characterised. All material properties, including those of the Boom Clay, are assumed to be isotropic. In the main series of calculations, it was assumed that a storage/cooling period of 50 years elapses between vitrification (for HLW) or reactor defuelling (for spent fuel) and supercontainer assembly. Subsequent cases using storage/cooling periods of 60 and 70 years were also considered. Calculations considered the thermal perturbation due to a long tunnel filled with supercontainers, and included the impact of heating by waste emplaced in adjacent tunnels. It was concluded that the requirement for the temperature of the outer surface of the overpack to remain below 100 C could only be satisfied if the cooling period was 60 years or longer [77] Vitrified HLW The calculated temperature evolution is shown in Figures 4.1 and 4.2 for vitrified HLW assuming a cooling time of 60 years. NIROND-TR E, April

72 2.5 1 year 2 years 5 years Boom Clay waste gallery lining backfill buffer 10 years 20 years 50 years Figure Predicted temperature field in the repository near-field for the case of vitrified HLW disposal with a normal concrete buffer, assuming a cooling time of 60 years [78]. Axes show distance in metres from the mid-point of the supercontainer 58 NIROND-TR E, April 2008

73 T ( C) time (years) Figure Predicted evolution of the absolute temperature with time in the repository near-field for the case of vitrified HLW disposal with a normal concrete buffer, assuming a cooling time of 60 years [78]. The absolute temperature for HLW (T abs) is calculated as follows: T abs = T + T i + 2 T ng where the initial in-situ temperature, T i, is 15.7 C and the temperature increment due to neighbouring tunnels, T ng, is negligible for the first 50 years. Line 1 represents the calculated temperature at the surface of the overpack. Line 3 represents the temperature at the surface of the envelope. Line 5 represents the temperature within the tunnel liner. Line 7 represents the temperature in the Boom Clay 2.75 metres from the middle of the supercontainer. Lines 1, 3, 5 and 7 represent the temperature evolution at points on a line through the mid-point of the supercontainer The preliminary results shown in Figures 4.1 and 4.2 indicate that for the case of vitrified HLW disposal with a normal concrete buffer, and assuming a cooling time of 60 years, the temperature close to the overpack surface peaks at about 92 C ( T ~76 C) after about 4 years. Within the tunnel liner the temperature peaks at about 54 C ( T ~38 C) after about 10 years Spent Fuel The preliminary calculations of Weetjens and Sillen [77] for spent fuel assumed UNE-type fuel with a burn-up of 55 GWd thm -1, because this fuel has the highest heat output per assembly, and can therefore be expected to generate the highest temperatures within the EBS of any Belgian UOX spent fuel type. Calculated maximum temperatures are similar to the results for vitrified HLW, although generally a few degrees higher, and the spatial thermal gradients generated within the buffer are also similar for both waste types. However, when the waste is spent fuel, the temperature remains high for a longer period than in the case of vitrified HLW. NIROND-TR E, April

74 The calculated temperature evolution for spent fuel, assuming a cooling time of 60 years, is shown in Figure T ( C) time (years) Figure Predicted evolution of the absolute temperature with time in the repository near-field for the case of spent fuel disposal with a normal concrete buffer, assuming a cooling time of 60 years [78]. The absolute temperature for spent fuel (T abs) is calculated as follows: T abs = T + T i + 2 T ng where the initial in-situ temperature, T i, is 15.7 C and the temperature increment due to neighbouring tunnels, T ng, is negligible for the first 50 years. In the case of spent fuel, the tunnel spacing is 120 m. The key to the numbered lines is given in Figure 4.2 The preliminary results shown in Figure 4.3 indicate that for the case of spent fuel disposal with a normal concrete buffer, and assuming a cooling time of 60 years, the temperature close to the overpack surface peaks at about 98 C ( T ~82 C) after about 10 years. Within the tunnel liner the temperature peaks at about 72 C ( T ~56 C) after about 15 years. 4.3 Main Uncertainties As noted above, heat transfer within the EBS is dominated by conduction. Therefore, apart from the heat output of the waste, there are two main parameters associated with each of the materials of the EBS that influence the thermal evolution thermal conductivity and volumetric heat capacity. In the studies of Weetjens and Sillen [77], the heat capacity was taken to be the arithmetic mean of the constituent phases for each EBS material [69] [10]. Weetjens and Sillen [77] undertook sensitivity studies to evaluate the influence of the thermal conductivity of the buffer and the Boom Clay on the results. Note that the thermal conductivity of the buffer can be influenced by design choice. On the other hand, the thermal conductivity 60 NIROND-TR E, April 2008

75 of the Boom Clay cannot be changed, but it can be determined with greater accuracy. The sensitivity studies showed that: Varying the thermal conductivity of the buffer between 1.4 and 2.0 W m -1 K -1 has a significant influence on the overpack temperature but negligible influence on buffer, lining and backfill temperatures. Varying the thermal conductivity of the Boom Clay between 1.3 and 2.2 W m -1 K -1 has a significant influence on the whole of the temperature field, and the impact of this uncertainty is much greater than the uncertainty in buffer conductivity. Although the thermal conductivity of the Boom Clay has been measured on several occasions there is still uncertainty over both the precise value and the anisotropy of this parameter. The strong dependence of the thermal evolution on the Boom Clay thermal conductivity make it essential that it be determined with a high degree of accuracy in order to underpin the predicted thermal evolution. The influence of cooling time, the time between waste emplacement in canisters, or reactor defuelling, and the moment when supercontainers are assembled and emplaced in the repository, has also been investigated. The main set of calculations by Weetjens and Sillen [77] were made assuming that the cooling time was 50 years and this assumption results in maximum temperatures at the overpack surface exceeding 100 C. Other calculations were made assuming a cooling time of 60 years and indicate lower maximum overpack surface temperatures (see Figures 4.2 and 4.3). A cooling time of 60 years or longer would result in the overpack surface temperature for vitrified HLW and spent fuel supercontainers remaining below 100 C. Note that similar scoping calculations for the smaller MOX supercontainer are yet to be carried out. The calculations were performed with a tunnel geometry involving only a 25-cm backfilled annulus around the supercontainer. A thicker backfill with relatively low thermal conductivity (1W/mK) could result in somewhat higher temperatures within the supercontainer, particularly in the case of the smaller MOX supercontainer. Depending on the final EBS geometry, the calculations may need to be repeated to verify if the overpack surface temperature criterion will still be met. Weetjens and Sillen [77] also evaluated a situation in which no backfill was emplaced around the supercontainer and there was instead an air gap. Calculations indicate that there is little difference in temperature gradient across the gap in the backfilled and the unbackfilled cases. However, for the unbackfilled case, results are strongly dependent on the emissivity of stainless steel. When the emissivity is small, convection is the main heat transport process and when it is larger, radiation becomes more important. Weetjens and Sillen [77] consider that the emissivity of stainless steel will be of the same order of magnitude as concrete, and therefore radiation is likely to be an important heat transfer process if no backfill was emplaced. A general conclusion is that in terms of thermal consequences, there is little difference between a backfilled and an unbackfilled situation, assuming that the distance between the outside of the supercontainer and the tunnel liner is not significantly greater than 25 cm. However, a NIROND-TR E, April

76 backfilled situation is more favourable for limiting corrosion of a stainless steel envelope, if such an envelope is present. A potential additional heat source term is represented by the heat of hydration of the backfill. However, it is believed that this contribution is unlikely to have a significant influence on the temperature field (R. Webers, personal communication). Another source of potential uncertainty is the influence of the initial supercontainer temperature. The scoping calculations of Weetjens and Sillen [77] assumed that the supercontainer was initially at the temperature of the repository surroundings. This assumption is not strictly correct because the concrete parts of each supercontainer will begin to heat up the moment the overpack is inserted into the buffer, during the supercontainer fabrication process. However, the thermal storage capacity of the supercontainer is limited, compared to the thermal output of the waste. Sensitivity studies assuming different initial temperatures revealed negligible influence on the overall thermal evolution. 62 NIROND-TR E, April 2008

77 5 Hydraulic Evolution 5.1 Hydraulic Processes Within the EBS The repository tunnels will be backfilled and closed soon after supercontainer emplacement, and the repository will become resaturated fairly rapidly after closure. At the point of resaturation, the relative pressure within the backfill and acting on the outside of the supercontainer will be zero and there will be no further suction. Over the next hundred years, the pressure will rise and will approach the hydrostatic pressure for the waste emplacement depth, and depending on temperature increase, could even exceed it. The most significant hydraulic processes acting within the EBS during the post-closure phase will be water transport under unsaturated and saturated conditions, and gas transport and dissolution: Water transport under unsaturated conditions will be complex and affected by several different processes including pressure gradients, capillarity and thermal effects. For example, within a sealed supercontainer with an intact envelope, pore water that is initially present in the concrete buffer may be driven outwards by thermally-driven vapour flow. Additional vapour may be generated by dehydration of hydrous cementitious phases within the buffer. The vapour may condense within the cooler part of the buffer increasing the degree of saturation there. If the supercontainer envelope is initially sealed, unsaturated conditions will prevail within the buffer until the envelope becomes perforated and external pore fluids are able to infiltrate and saturate the buffer. Unsaturated conditions are likely to prevail for a much shorter period in the EBS outside the supercontainer. The backfill and tunnel liner are both expected to saturate within a few years of tunnel closure, and the buffer within a few years of envelope perforation, thus these scoping calculations were based on a 25 cm-thick backfill and preliminary hydraulic parameters. Water transport under saturated conditions is constrained by the hydraulic gradient, and also the hydraulic conductivity of the buffer, backfill and other EBS materials. The hydraulic conductivity is chiefly a function of the porosity of the relevant porous medium, and is also influenced by pore fluid salinity and temperature. Gas transport and dissolution within the EBS results mainly from the generation of gases near to the overpack and envelope surfaces by anaerobic corrosion, and within the buffer close to the overpack due to radiolysis. Late in the evolution of the EBS, after perforation of the overpack, there may be release of gases generated by radioactive decay of the waste. Gas transport through the EBS is most likely to involve dissolution and diffusion in pore fluid or advection with displacement of pore fluid (i.e., immiscible two-phase flow.) NIROND-TR E, April

78 5.2 Expected Hydraulic Evolution Saturation of the Backfill and Buffer The heat flux from the waste may have a pronounced influence on the duration of unsaturated conditions in the backfill. To investigate this, Weetjens et al. [78] performed hydraulic and coupled thermo-hydraulic simulations to investigate the effect of elevated temperatures on the saturation of the backfill. Calculations were performed for a 1D radial geometry using the PORFLOW code [61]. An initial saturation state of 70% was assumed for all cementitious EBS materials. A uniform backfill thickness of 25 cm was assumed and all materials were considered to be continuous and isotropic. For the Boom Clay, unsaturated parameters were obtained based on fitting observations from a single column experiment to Van Genuchten retention and relative conductivity characteristics [74]. For concrete and cementitious backfill materials, literature data gathered in the context of Category A studies were used. The boundary condition applied for the model of resaturation of the buffer was hydrostatic pressure at the envelope. This represents the hydraulic state after the backfill has become saturated and the pressure in the backfill has equilibrated to hydrostatic pressure. At this point the envelope is assumed to vanish, so that hydrostatic pressure is experienced by the outer surface of the buffer. In a realistic situation without an envelope, buffer and backfill would resaturate simultaneously and the pressure gradient would tend to be lower and take longer to equilibrate. In the case of an envelope that is perforated by corrosion, the 3D configuration of resaturation through a small hole would probably be complex and the resaturation time could be much longer. Additional calculations would be needed for detailed evaluation of resaturation after envelope perforation. The saturation profiles resulting from the hydraulic simulation of the water flow into the buffer are shown in Figure 5.1. The calculations are based on preliminary material characterisation data, and are strongly dependent on the interplay between the unsaturated characteristics of the concrete and the Boom Clay. The backfill exerts quite a large suction potential and its saturation state increases quite rapidly. However, this occurs at the expense of the saturation of the Boom Clay close to the tunnel lining. After about 1 year, the concrete backfill materials are ~99% saturated with water, but the first decimetres of Boom Clay become only partially saturated. Desaturation of the Boom Clay could temporarily open cracks and create space for microbes. After about 2 years, however, the near-field becomes completely saturated. Upon reaching hydraulic saturation, the pore water pressure is initially at 0 m relative (no further suction), and it takes about 100 years until this pressure is equilibrated with the in-situ hydrostatic water pressure (Figure 5.2). 64 NIROND-TR E, April 2008

79 Figure Saturation across the outer part of the EBS as a function of time after backfilling, but before envelope perforation [78] Figure Water pressures as a function of time in the outer part of the EBS [78] If the heat flux is also taken into account by making a thermo-hydraulic simulation, the water flow (and hence, the saturation process) occurs approximately twice as quickly as in the NIROND-TR E, April

80 simulations that only considered hydraulic forces, as a result of viscosity effects (Figures 5.1 and 5.2). The increase from 70% saturation to approximately 95% occurs within just one month, but the saturation of the residual 5% takes longer, between one and two years [78]. Calculations to estimate the time required for the buffer to become saturated after the resaturation and attainment of hydrostatic pressures in the backfill, and following removal of the stainless steel envelope, were also undertaken by Weetjens et al. [78]. In these calculations the buffer was assumed to have the same hydraulic properties as the backfill and tunnel lining. The resulting saturation profiles are illustrated in Figure 5.3. As was expected on the basis of the assumed material properties and the unsaturated volume, the saturation of the buffer takes approximately two years. As was seen for the backfill, a buffer material exerting a large suction can lead locally to some temporary de-saturation of the Boom Clay close to the repository tunnels. Given the assumed hydraulic properties, these calculations suggest that if the supercontainer envelope was not present, saturation of the EBS including backfill and buffer, right up to the overpack surface, would be complete within a few years of tunnel closure. Such rapid saturation would tend to validate assumptions regarding the electrochemical evolution and corrosion of overpack [41]. However, this result depends crucially on the assumptions made regarding the unsaturated material characteristics the retention curve (capillary pressure function) and conductivity curve (relative conductivity as a function of pressure or saturation), as well as the porosity of the EBS materials and their initial saturation state, the hydraulic conductivity of the Boom Clay and the pressure gradient around the tunnel. The BACCHUS2 and RESEAL test results indicate that the hydro-mechanical behaviour at the interfaces between the backfill, tunnel liner and Boom Clay also have an important influence. The specific storage parameters do not influence the saturation transient but they do influence the time to reach hydrostatic pressure after saturation. Unless EBS materials with a permeability much smaller than that of the Boom Clay are used, it is expected that the EBS resaturation time will be controlled by the hydraulic characteristics of the Boom Clay. Although the clay is capable of providing the necessary water to fully saturate the EBS in a few years, due to its low hydraulic conductivity, the Boom Clay is likely to limit the rate at which water is supplied. An EBS with high suction will be unable to draw water quickly from the Boom Clay because as the clay desaturates, its hydraulic conductivity falls rapidly. In the simulations of Weetjens et al. [78], a relatively large but temporary underpressure was developed in the first metre of Boom Clay outside the tunnel lining (see Figure 5.3). Note that trapped air is likely to have a negligible influence on the EBS saturation process. Calculations carried out in the context of the BENIPA project [7] have demonstrated that the saturation time and associated saturation profiles are similar for calculations involving unsaturated flow only taking into account the movement of pore water (made using PORFLOW), and also for two-phase flow taking into account air and water (made using TOUGH2). This result holds for cool or hot conditions (that take into account heating due to 66 NIROND-TR E, April 2008

81 radioactive decay) because in either case trapped air can dissolve rapidly enough in pore water so that it does not hinder the resaturation process. The BENIPA calculations were made for saturation of a bentonite buffer but the implications remain valid for the case of the supercontainer concrete buffer. Figure Saturation profiles within the EBS as a function of time after the attainment of hydrostatic pressures in the backfill and envelope removal [78] In conclusion, backfill and buffer saturation will commence immediately after backfilling and sealing of the tunnels and both are expected to be complete in just a few years if no envelope is present. The exact timing of perforation of an initially sealed envelope is uncertain. Weetjens et al. [78] assumed that a sealed envelope would become perforated within a few hundred years by localized corrosion. This is because under high-ph conditions stainless steel corrosion is dominated by localised phenomena such as pitting corrosion and stress corrosion cracking. It is possible that the corrosion potential of the envelope will be sufficiently low that localised corrosion due to chloride ions does not occur [42], although in such circumstances it may still be possible for reduced sulphur species, if present, to promote localised corrosion. In the absence of localised corrosion, the uniform corrosion rate of stainless steel is so low that perforation of the envelope would take thousands of years. Note that localised envelope failure (e.g., the development of a small pinhole) would be likely to lead to a more complex, heterogeneous buffer saturation scenario than that modelled by Weetjens et al. [78], and resaturation times in this case could be much longer. Based on currently available knowledge, and disregarding the potential consequences of gas generation by corrosion of the overpack (see Section 8.3), if the envelope is initially sealed, it NIROND-TR E, April

82 seems likely that the buffer will only become fully saturated hundreds to thousands of years after repository closure Thermal Dehydration Effects Changes to the content and spatial distribution of water in the supercontainer buffer concrete during heating, and the potential water vapour pressures generated if a sealed envelope is present, are key uncertainties in the evolution of the EBS. Therefore, preliminary calculations have been carried out by Poyet [58] [59] in order to evaluate the thermal dehydration of the concrete buffer after emplacement of the overpack. The calculations were made for a two-dimensional section perpendicular to the axis of the supercontainer from the overpack, across the EBS and into the Boom Clay. The calculations of Poyet [58] assumed constant boundary conditions of 100 C at the overpack surface and 16 C at a distance of 23 metres into the Boom Clay. Although an overpack surface temperature of 100 C will not be sustained, this assumption was used to obtain an indication of the evolution of the buffer under extreme conditions. Subsequent calculations involved a time-varying thermal output to simulate more accurately the evolution of heat from a supercontainer with vitrified HLW [59]. Two possible supercontainer designs were considered (see Table 5.1 and Appendix A for input data and boundary conditions): 1) Ordinary concrete buffer, cured at 20 C with no water loss (water:cement mix ratio of 0.43); Phase 2 filler of portlandite (Ca(OH) 2 ). 2) Ordinary concrete buffer, partially dried at 60 C and having a reduced moisture content; Phase 2 filler of lime (CaO). Table Properties of the supercontainer concrete assumed in the two calculation cases considered for the thermal dehydration modelling [58] Property Units 20 C cure concrete 60 C dried concrete Porosity Intrinsic permeability m 2 2.3E E-17 Thermal conductivity W m -1 K Free water kg m Bound water kg m Saturation The thermal dehydration modelling was conducted for a closed system with a sealed envelope. The system was represented as a rigid concrete matrix with a pore space containing a mixture of liquid water and water vapour (no air). The simulations of Poyet [58] considered only the heating phase. Subsequently, Poyet [59] considered both heating and cooling of the buffer assuming a vitrified HLW heat source. It is assumed that similar conclusions can also be applied to a supercontainer containing a spent fuel heat source. 68 NIROND-TR E, April 2008

83 Both models account for the equilibrium between liquid water pore fluid and water vapour, and for the release of bound water from the concrete. The quantity of water released from the concrete, and the consequent change in porosity, were modelled as a function of temperature. It was assumed that no bound water would be released until the temperature exceeds 60 C. Above 60 C, water release was assumed to increase linearly at a rate of kg m -3 per degree, based on experimental data. The porosity was assumed to increase to accommodate water released by dehydration. The Van Genuchten equations [74] were used to determine relative and effective permeabilities of the concrete for liquid water and water vapour, based on the degree of water saturation. The thermal dehydration calculations of Poyet [58] suggested that a range of relatively minor changes may occur in response to heating as follows: For both concretes considered, small amounts of bound water may be released from the supercontainer concrete buffer as a result of heating by the waste. The maximum amount released will be from the concrete closest to the overpack and will comprise 13% of the total amount of bound water (equivalent to 9 kg m -3 ). A small and non-uniform increase in porosity will be created by the dehydration of the concrete, up to a maximum of 10.4% to 10.8%. A small increase in the saturation will be caused by the release of bound water into the buffer pore space. For Case 1 (concrete cured at 20 C), saturation increases by about 8%. For Case 2 (concrete partially dried at 60 C), saturation increases by about 28%. A small increase in vapour pressure associated with the elevated temperatures. For Case 1 (concrete cured at 20 C), vapour pressure reaches around 0.1 MPa. For Case 2 (concrete partially dried at 60 C), vapour pressure reaches around 0.02 MPa. Similar results were generated by the more detailed modelling of Poyet [59]. The more recent analysis suggests that the temperature increase will induce little dehydration of the concrete (from 0.4 to 4.5 kg m -3 ). There will be a small porosity increase (from 10.4% to 10.6%) and limited vapour generation (the maximum pressures reached are and MPa for the two cases). For the two cases, the heating is not expected to generate a dry zone close to the overpack. The evolution of saturation at various points within the supercontainer is shown in Figure 5.4 for the case in which the buffer is cured without water loss and has an initial saturation of 0.806, and for the case in which the buffer is dried and has an initial saturation of For Case 1 (concrete cured at 20 C), saturation increases from to (~4%). For Case 2 (concrete partially dried at 60 C), saturation increases from to (~23%). The change is probably less than the uncertainty in buffer porosity. NIROND-TR E, April

84 Figure The saturation state at the boundaries of the concrete buffer in response to heating [59]. The right hand diagram shows the location of points within the supercontainer, three at the surface of the overpack and three at the surface of the envelope. The evolution of saturation at each point with time is given on the graphs to the left. The upper graph gives results for a buffer with an initial saturation of while the lower graph gives results for a dried buffer with an initial saturation of The graphs indicate a rapid increase in saturation due to heating (and dehydration of hydrous cementitious phases) and a slower decrease as the saturation state within the buffer homogenises For both cases the patterns are similar a rapid increase in saturation during the heating phase followed by a slower decrease leading to homogenisation of the saturation within the buffer. 70 NIROND-TR E, April 2008

85 The difference in saturation for the two cases causes a significant difference in water transfer rate, and therefore the rate of change of saturation. This is because the permeability, as calculated from the Van Genuchten equations, is ~1,500 times faster in the case with a higher initial saturation. The main conclusions to be drawn from the studies of Poyet [58] [59] are that the evolution within the supercontainer is similar for both of the cases considered (a concrete buffer cured with no water loss, and a buffer that is partly dried before supercontainer assembly). Moreover, the temperature increase in the concrete buffer will induce little dehydration and porosity increase and will not generate a dry zone close to the overpack. Various simplifications have been made in the modelling and uncertainties remain. Key uncertainties are associated with the thermal conductivity data, the assumed cement dehydration relationships, and the potential impact of considering the thermal expansivity of water. Note that the study assumed significantly higher buffer thermal conductivity (2.3 and 2.7 W/mK) than the thermal modelling of Weetjens and Sillen [77] (1.7 W/mK) Gas Effects Weetjens et al. [78] have analysed gas transport within the supercontainer buffer focusing on the production of hydrogen. The objectives of the study were to quantify the amount of gas that could be formed in the repository near-field, and to evaluate whether diffusive gas migration through the EBS is fast enough to prevent the build-up of large gas pressures. It was found that the main gas production mechanisms within the repository near-field are likely to be anaerobic corrosion of steel and radiolysis. However, hydrogen production due to gamma radiolysis is likely to be quite small in volumetric terms compared with hydrogen production due to anaerobic corrosion of the steel EBS components. Gases (e.g., helium, radon) will also be generated by radioactive decay, but these will only be released following overpack perforation. The key parameter influencing gas effects is the rate of gas generation, constrained principally by the uniform corrosion rate at the overpack surface (corrosion), and by the gamma radiation field at the overpack surface and the water content of the buffer (radiolysis). The solubility of hydrogen and its diffusivity in buffer pore fluid determine how quickly any gas generated can be removed in solution. Although some uncertainty surrounds the rate of gas generation by radiolysis, due, for example, to the potentially complex effect of temperature on the kinetics of radiolysis, following the onset of anaerobic conditions, the anaerobic corrosion of the steel EBS components will be the main source of hydrogen gas. Despite its larger surface area, gas produced by corrosion of the stainless steel envelope, if present, will be less significant than gas produced by corrosion of the carbon steel overpack, due to the very low uniform corrosion rate of stainless steel. An estimate of the hydrogen production due to anaerobic carbon steel corrosion was made, assuming a corrosion rate of 1 µm per year for the first 100 years, followed by 0.1 µm per year thereafter. Note that these values are conservative compared with the likely long-term uniform corrosion rate of carbon steel under high-ph, anaerobic conditions (~<0.1 µm per year; [65]; see also Section 8 and Kursten [40]). Note also that uncertainty NIROND-TR E, April

86 remains concerning any potential couplings between corrosion and radiolysis at the overpack surface. A series of 1D radial transport calculations were made using the code PORFLOW to evaluate the rate of removal of hydrogen by diffusion of dissolved gas in Boom Clay pore fluid for a range of diffusion coefficients [78]. The scoping calculations assumed a 25 cm-thick backfill. Results of these transport simulations were compared with the hydrogen production rate to see if hydrogen could be removed purely by a diffusive flux. The results showed that a free gas phase will be generated soon after repository closure for the lower range of diffusion coefficients considered. Multiphase gas migration calculations were made using the TOUGH2 code in order to evaluate the pressure distribution within the repository near-field assuming that the envelope was perforated [78]. The pressure increase was dominantly caused by gas generation at the overpack surface and the only significant pressure increase was within the concrete buffer (Figure 5.5). The highest calculated pressure was 3.4 MPa at a point close to the overpack after ~100 years and at this time the gas occupied 24% of the porosity. This pressure may be compared with lithostatic pressure in the Boom Clay at the expected repository depth (~4.5 MPa), and the expected tensile strength of the buffer (~2 MPa). Note that this time marks the end of the period when the assumed corrosion rate was 1 µm per year. 72 NIROND-TR E, April 2008

87 Figure Pressure distribution in the near-field calculated by Weetjens et al. [78] assuming that the envelope is perforated and that corrosion of overpack and envelope both start at time = 0. The white object to the left of each diagram represents the overpack and the envelope is situated at a distance r = 1 m on the x-axis The likely effects of radiolytic gas production are discussed in Section Main Uncertainties The range of processes that will affect the hydraulic evolution of the near-field is complex, and further more detailed modelling and experimental work is underway to further constrain the magnitudes and uncertainties associated with the various processes, the couplings between them, and the rates at which they will occur. Examples of uncertainties bearing on EBS evolution include: There should be better characterization of the unsaturated behaviour of the near-field materials, both governing resaturation processes and governing gas generation and twophase flow. A good estimate of the initial degree of saturation of near field materials should also be made. The expected relationship between cement dehydration and temperature needs to be better defined, including possible kinetic effects, especially for the concrete of the buffer, as this will govern the quantity of water liberated during buffer heating. There are large uncertainties associated with the gas transport parameters for the Boom Clay and EBS materials and results are highly dependent on these parameter values, especially for the two-phase flow calculations. This applies to the hydrogen diffusion NIROND-TR E, April

88 coefficient, and to the unsaturated characteristics of the media, especially gas entry pressure, liquid permeability and amount of storage given by porosity Gas generation, as a result of corrosion of the overpack, could have a significant influence on the mechanical evolution of the supercontainer, potentially promoting cracking in the cementitious components of the EBS, and threatening the integrity of a sealed envelope. Further research is attempting to constrain the expected corrosion behaviour and the rate and total quantity of gas generated by overpack corrosion. There is uncertainty in the likely mode of gas transport whether a continuum approach is appropriate or whether fingering or local breakthrough may be more relevant. In the gas transport modelling a two-phase flow model is used, assuming a continuous, isotropic porous media. The behaviour of the solid phase is reduced to a storage coefficient and assumes elastic behaviour with negligible displacements. Such a model may not be valid if computed gas pressures become comparable to in-situ stresses (a few tens of bars). The hydraulic evolution of the buffer will differ greatly depending on whether the envelope is initially sealed, perforated or absent. If the envelope is sealed and the buffer has a high initial saturation state, overpack corrosion may still occur, causing gas pressure to build up within the supercontainer and potentially even leading to rupture of a sealed envelope and sudden release of gas. Alternatively, an initially unsealed or absent envelope would allow gas to be continuously released so that there would be less tendency for pressure build-up within the supercontainer. In either case, the mechanical couplings and effects, on both the EBS and on the Boom Clay, must be evaluated. 74 NIROND-TR E, April 2008

89 6 Mechanical Evolution 6.1 Mechanical Processes Within the EBS Repository excavation will change the regional stress field at the repository and may create a zone adjacent to the excavations in which the properties of the host rock are altered as a result of stress modification, leading to creep and/or the formation of new micro-fractures. Excavation-induced deformation effects will be mitigated to some extent, at least initially, by the emplacement of tunnel liners, which will resist the external lithostatic pressure. The highly coupled hydromechanical behaviour and the viscoplastic nature of the Boom Clay is such that excavation-induced micro-fractures are expected to heal and become sealed. The tunnels will be backfilled and closed soon after supercontainer emplacement, and the repository will become hydrologically resaturated fairly rapidly after closure. Approximately 100 years after resaturation, pressures within the backfill and acting on the outside of the supercontainer will approach the hydrostatic pressure for the waste emplacement depth, and depending on temperature increase, could exceed it (see Section 5.1). Once the tunnel liners begin to degrade and fail structurally, the clay may creep further and cause pressures in the backfill and acting on the outside of the supercontainer to increase and approach ambient lithostatic pressures. Depending on the composition of the backfill material, these effects may be accompanied by a decrease in backfill porosity. Several relevant mechanical effects derive from couplings with thermal or hydraulic processes. For example, heat from the waste, pore pressure changes during saturation of the backfill and buffer, and chemical processes leading to gas generation may all drive mechanical changes within the EBS and the immediately surrounding rocks including: Thermal expansion and contraction of EBS components leading to tensile and compressional stresses. Volume and porosity changes associated with the development of corrosion products at the outer surface of the overpack, and to a lesser extent, the envelope, if present. Mineralogical and porosity changes within the buffer concrete, backfill and tunnel liner components of the EBS associated with the development of high-temperature more crystalline cementitious phases, which often have lower molar volumes than their lowtemperature equivalents. Chemical reaction between HCO 3 - waters from the Boom Clay and the more alkaline waters of the concrete buffer may lead to porosity reduction through calcite precipitation. Gas generation, which has the potential to lead to increased pressure within the supercontainer, regardless of whether or not an envelope is present. If the envelope is perforated or absent, external pore fluids will quickly saturate the buffer. Gas pressure close to the overpack may rise depending on the interplay between gas generation by corrosion and radiolysis, and gas transport by diffusion and/or two-phase flow. If the envelope is sealed, pore water initially present in the buffer concrete may promote corrosion, resulting in gas generation and increase in internal pressure. NIROND-TR E, April

90 Cracking of the buffer concrete in response to thermal stresses and gas generation. Fracturing or disturbance of the Boom Clay resulting from high gas pressure, suction during near-field resaturation, or thermal effects. These effects are most likely to occur during the early period after closure (perhaps the first few hundred to a few thousand years), when the thermal gradients and the potential for chemical reaction are expected to be at their greatest. Mineralogical changes, cracking of EBS materials, faulting and fracturing (including reactivation of pre-existing features) could affect groundwater flow and radionuclide transport pathways, and could cause minor displacement of EBS components. The interaction of high-ph, EBS-derived fluids with the Boom Clay adjacent to the disposal system could cause mineralogical changes leading to a reduction in the swelling properties of the clay. Desaturation of the Boom Clay may open up cracks and create void space in which microbes may develop. 6.2 Expected Mechanical Evolution Thermal Stresses Stresses will be generated within the buffer concrete as a result of heating caused by cement hydration reactions during concrete curing, and also by radioactive decay of the waste. Belgatom [6] have made 2D model calculations, using the SYSTUS code, of the evolution of stress in the ONDRAF/NIRAS supercontainer concrete buffer in response to the two different forms of heating. The model corresponded to an axisymmetric cross section at the supercontainer mid-point. The calculations were intended to provide an order of magnitude estimate of the thickness of the zone within which fracturing of the buffer might occur, so that the potential need for reinforcement could be evaluated. It was recognised that the potential for fracturing would be greater nearer to the ends of the supercontainer. However, consideration of the whole of the buffer requires more sophisticated 3D calculations that have yet to be carried out. Calculations were made in two stages, the first corresponding to heating during concrete curing following casting of the Phase 1 concrete buffer, the second due to heating by radioactive decay of the waste following insertion of the overpack (see Figure 2.6). The second calculation assumed that heat is generated at 300 W per metre length of the overpack. Most of the parameters used in the modelling were taken from literature review. Both calculations were made independent of each other by assuming a long resting time between the two stages. The stresses generated by each stage were added to obtain the final stress state. Concrete Curing The stress calculation due to concrete curing was made in two stages: 76 NIROND-TR E, April 2008

91 In the first stage, the temperature distribution due to heat of hydration was computed. In the second stage, the stress distribution corresponding to this temperature distribution was inferred. For a heat transfer coefficient equal to 5 W m -2 C -1 on both walls of the supercontainer (adiabatic cure), compression is experienced throughout the buffer 16 days after casting. For a heat transfer coefficient equal to 5 W m -2 C -1 on the inner wall of the supercontainer and 16 W m -2 C -1 on the outer wall of the supercontainer (ventilated store), there is a tensile zone at the outside perimeter of the buffer. However, this tensile zone is limited to about 10 to 12 cm in thickness with a maximum stress of about 0.4 MPa. Adiabatic curing therefore results in a better contact between the concrete and the envelope. Overpack Thermal Loading The stress calculation due to the thermal loading of the overpack was made in two steps: In the first stage, the temperature distribution due to the thermal power of the waste was computed. In the second stage, the stress distribution corresponding to this temperature distribution was inferred based on a modified Young s Modulus curve [6]. The final stress state was obtained by adding together the stress due to concrete curing and due to overpack thermal loading. Results indicated a narrow tensile zone near the outer surface of the buffer, meaning that tensile cracks may appear in the buffer to a depth of about 5 cm. Belgatom [6] considers that the calculated stress levels throughout the buffer are acceptable, both in tension and compression. However, stress levels at the ends of the buffer have yet to be evaluated Corrosion Effects The potential mechanical impact of corrosion within the supercontainer, resulting from the generation of corrosion products at the overpack and envelope surfaces, is at present uncertain. At the expected low corrosion rate, the formation of corrosion products is likely to reach a steady state limited by diffusion [13] [44]. Observation of corrosion of steel within concrete fence posts is equivocal with some evidence suggesting that the volume change is unlikely to be enough to disrupt the concrete structure [62], while other observations suggest that the volume change may be enough to cause spallation [67]. This section will be developed further when more information is available concerning the potential mechanical impact of corrosion within the supercontainer. 6.3 Main Uncertainties The range of processes that will affect the mechanical evolution of the near-field is complex, and further more detailed modelling and experimental work is underway to further constrain the magnitudes and uncertainties associated with the various processes, the couplings between NIROND-TR E, April

92 them, and the rates at which they will occur. Examples of uncertainties bearing on EBS evolution include: Gas generation, as a result of corrosion of the overpack and envelope, could have a significant influence on the mechanical evolution of the supercontainer, potentially promoting cracking in the cementitious components of the EBS and threatening the integrity of a sealed envelope. The mechanical impact of gas generation on the EBS must be further evaluated. The impact of thermal stresses on the concrete buffer has yet to be evaluated in 3D, and the potential for fracturing near the ends of the supercontainer is therefore not yet known. Complex mechanical effects are expected in the Boom Clay and EDZ surrounding the disposal tunnels, including strong hydro-mechanical coupling. Further research is focusing on fracture development in the EDZ, viscous effects, anisotropy and the effects of suction on the near-field, which could play a role in fracture creation [79]. 78 NIROND-TR E, April 2008

93 7 Chemical Evolution This section of the report focuses on the chemical evolution of the EBS components, and also includes a description of processes related to radiological (radiolysis) and biological (microbes) processes. A detailed account of the electrochemical evolution of the metallic barriers, including a description of likely corrosion rates and processes, is given in Section Chemical Processes in the EBS Prior to repository construction, the rocks and groundwaters at the repository location are assumed to be typical of the Boom Clay, a silty clay or argillaceous silt with a high pyrite and glauconite content in the more silty layers, and variable carbonate and organic matter contents. The more reactive mineralogical components of the Boom Clay are calcite, siderite and pyrite, and these will control the concentrations of dissolved calcium and iron, and the redox potential of the groundwaters. Boom Clay groundwaters are typically dilute NaHCO 3 solutions with a relatively high concentration of dissolved organic matter, a ph of ~8.5, and a redox potential (Eh) of around -300 mv [16]. Chemical conditions in the cementitious EBS components will initially be highly alkaline (their pore waters will have a ph of 12.5 or higher). These conditions will be affected by the elevated temperatures associated with the wastes, leading to some mineralogical changes. Redox conditions inside the supercontainer are expected to be anaerobic at all but the earliest time after supercontainer closure. The repository tunnels are also expected to return to anaerobic conditions shortly after they are backfilled and sealed, owing to the consumption of oxygen by corrosion reactions and the oxidation of Boom Clay pyrite and organic matter. Although microbial activity will occur in the repository while oxidising conditions exist prior to, and for a period shortly after, closure, their activity is expected to decline as a result of the high ph, high temperatures, and lack of light and space. Chemical conditions in the EBS may also evolve as a result of interactions between the materials of the EBS and ingressing Boom Clay groundwaters. These interactions are expected to include a gradual and slow reduction in ph, precipitation in the cementitious regions of the EBS environment of various minerals (e.g., calcium carbonate), and slow corrosion reactions at the envelope and, eventually, at the outer surface of the overpack. The following sections describe the chemical conditions and effects that are expected to exist and occur within the near-field during the following periods in repository evolution: The initial conditions, for example, in the Boom Clay immediately surrounding the EBS and in the newly-formed buffer concrete. Conditions and chemical evolution during repository construction and operation. Conditions and chemical evolution after repository closure. NIROND-TR E, April

94 Although the description has been divided according to these three periods, it should be remembered that, in reality, some effects and processes will continue from one period to the next. 7.2 Initial Chemical Conditions Initial Chemical Conditions in the Boom Clay The mineralogical composition of Boom Clay surrounding the EBS has been summarised in Table 2.5 and Table A7.2 (data from [15]). Viable microbes are most probably present within the undisturbed Boom Clay, but it is thought that microbial activity is limited because of the small size of the pores in the clay [80]. The composition of a reference pore water from the Boom Clay is shown in Table 7.1. Details of the experimental and modelling procedures used in deriving this reference water composition are presented in [16]. 80 NIROND-TR E, April 2008

95 Table A reference Boom Clay pore water composition [16] Species or Parameter Concentration Units Na K Ca Mg Fe Si 0.1 mmole litre -1 Al HCO Cl Total S 0.02 SO Total Inorganic Carbon (TIC) 15.1 Dissolved Organic Carbon (DOC) mg C litre -1 ph 8.5 -log 10 [a + H ] pco atm Eh -274 mv Temperature 16 C Initial Chemical Conditions in the Buffer Two possible concrete formulations that are being considered for use as the buffer in the supercontainer have been presented in Table 2.3. The likely mineralogical composition of the buffer following the initial hydration reactions that will occur during supercontainer fabrication is summarised in Table 7.2 (after [76]). NIROND-TR E, April

96 Table Likely mineralogical composition of the newly formed buffer Phase Representative formula Portlandite Ca(OH) 2 Afwillite Ca 3 Si 2 O 4 (OH) 6 Amorphous CSH Ettringite e.g., Ca 1.8 SiO 4.6 : 3.6H 2 O Ca 6 Al 2 (SO 4 ) 3 (OH) 12 :26H 2 O Hydrogarnet Ca 3 Al 2 (OH) 12 Hydrotalcite Mg 4 Al 2 (OH) 14 :3H 2 O Hematite Fe 2 O 3 Initially the proportion of CSH to afwillite is expected to be high and the ettringite content will be rather low or even zero [28]. The composition of pore water in the buffer will, in large part, be buffered by the mineral phases present, but may contain high concentrations of dissolved Na + and K +. Even though the cement specified for use as the buffer in the supercontainer design is a low-alkali cement ( 0.6% in weight Na 2 O equivalent 1 ), the expected concentrations of dissolved alkali metal hydroxides are such that the initial buffer pore waters are expected to have ph values >13 [76]. The composition of young pore water from the buffer has been estimated by Wang [76] from thermodynamic modelling studies, and is shown in Table 7.3. Details of the various assumptions made in deriving this water composition are presented in Wang [76]. 1 The weight Na 2O equivalent (Na 2O eq) is defined as follows: Na 2O eq = weight % Na 2O weight % K 2O. 82 NIROND-TR E, April 2008

97 Table Composition of young pore water from the buffer (after [76]). Element Concentration (mmole litre -1 ) Ca 0.7 Na 141 K 367 Al 0.06 Si Mg ~10-7 Fe 10-5 S 2 C 0.3 ph Expected Evolution During Repository Construction and Operation Construction and operation of the repository will lead to a range of chemical effects on the Boom Clay and the engineered barriers Oxidation Effects Oxidation effects are unavoidable during construction of the repository and ventilation of the tunnels. The principal oxidation effects on the Boom Clay are related to oxidation of pyrite and organic matter. Excavation of the repository tunnels will lead to fracturing of the Boom Clay in a 1 m zone around the tunnels the EDZ [8]. Although such excavation-induced fractures in the Boom Clay are known to close in a relatively short period [8] [72], oxygen from the atmosphere in the repository will cause some oxidation of the clay. For example, Van Geet [71] reports that petrographic analyses of samples of Boom Clay that were vacuum packed and protected from oxygen within a time span of 1 to 2 hours after sampling from the Northern Starting Chamber in the underground facility at Mol found precipitates of gypsum and Fe-(oxi)-hydroxides in the fractures, while intact framboidal pyrite remained within the matrix of clay blocks exposed to the air. Similar brown precipitates of Fe- (oxi)-hydroxides were also observed on Boom Clay samples retrieved from the excavation of the Connecting Gallery at Mol within one hour after excavation. Van Geet et al. [71] indicate that these oxidation processes are fast and operate on a period of several hours. NIROND-TR E, April

98 Oxidation of pyrite leads to the production of higher concentrations of dissolved sulphate (SO 2-4 ) and thiosulphate (S 2 O 2-3 ) in pore waters within and close to the fractures and the excavations. Reduced sulphur species will be aggressive to the metallic supercontainer components [71]. Experiments performed by Van Geet [71] showed high sulphate concentrations (up to 2,700 mg l -1 ) in the clay matrix next to a major fracture (Figure 7.1). A heterogeneous pattern of increased sulphate concentrations was observed up to at least 25 mm perpendicularly from this fracture surface. This latter effect was attributed by Van Geet et al. [72] to the presence of microfractures in the clay. Figure Dissolved pore water sulphate concentration as a function of distance from a fracture wall for a sample containing a brown precipitate (Sample a) and another sample without visible effects of oxidation (Sample b) Once the excavation-induced fractures in the Boom Clay seal, ingress of oxidising species to the clay will slow, although some oxidation processes will continue throughout the period while the repository is open and atmospheric oxygen remains. After the excavation-induced fractures close, the main processes affecting the spatial distribution of sulphate and thiosulphate in the clay will be diffusion, which will act to reduce chemical potential gradients, and advection, which will act towards the excavations in response to the strong inwardsdirected hydraulic gradient. 84 NIROND-TR E, April 2008

99 Despite evidence of a minor enhancement of Boom Clay hydraulic conductivity over a region up to 6-8 metres around the repository tunnels [8], chemical alteration will be much more limited. Studies of the extent of the oxidised zone around the tunnels suggest that it will be limited to about 1 m around each tunnel after 20 years [72] [78]. The studies modelled the advective-diffusive transport of oxygen dissolved in pore fluid using the code PORFLOW. The findings are supported by experimental data from the HADES underground research laboratory. The concept of sulphate migration towards the repository from the EDZ is consistent with observations made in the underground facility at Mol of sulphate salt precipitation on the concrete tunnel liner (Figure 7.2). Figure Photographs of sulphate salts precipitated in the Connecting Gallery at Mol, mainly along the joints of concrete blocks [72] NIROND-TR E, April

100 Sulphate salts that have been observed in the underground facility at Mol include thenardite (Na 2 SO 4 ), apthitalite (K 3 Na(SO 4 ) 2 ), trona (Na 3 (CO 3 )(HCO 3 ) 2H 2 O), burkeite (Na 6 (CO 3 )(SO 4 ) 2 ), and schoenite (K 2 Mg(SO 4 ) 2 6H 2 O) [72]. Such sulphate salt precipitation processes are expected to continue to occur on the liner during the period in which the repository is open, prior to hydraulic resaturation Introduction of Microbes Microbes will be introduced to the disposal system during repository excavation and operation. Microbial activity is, therefore, expected to occur in the repository and the EDZ prior to, and for a period shortly after, closure [55]. Increased microbial activity has been found within the EDZ as compared with the undisturbed Boom Clay [72]. High concentrations of Sulphate- Reducing Bacteria (SRB) have been observed in the Swiss Mont Terri research programme. In a repository environment, SRB may promote the return to reducing chemical conditions (see below), although other causes, primarily corrosion, are likely to be more important in the ONDRAF/NIRAS repository (this is discussed in more detail in Section 8). The activity of SRB will influence the evolution of sulphate, sulphide, polysulphide and thiosulphate concentrations in the disposal system. Microbial communities may form in cavities, for example between the tunnel liner and the Boom Clay, or as biofilms, for example on the surfaces of the concrete tunnel liners. It is therefore important to ensure that any void space adjacent to the tunnel liner is eliminated, wherever possible. Some species of microbe are known to persist under conditions of elevated ph, but their activity tends to be limited by a lack of available nutrients [43] [80]. Microbial degradation of concrete under anaerobic conditions is not expected to occur [43]. The buffer concrete is expected to be fairly sterile initially, but will normally contain a few per cent sulphur as SO 3, which could potentially be assimilated by SRB. Indeed, Fukunaga et al. [23] observed a tendency for low-eh to promote SRB activity under a ph as high as 10. Microbial activity within the supercontainer cannot, therefore, be ruled out. However, as noted by Herbert [32], microbial activity within the supercontainer is not expected to be significant because much of its interior will experience temperatures in excess of 70ºC - a temperature sufficient to kill most microbes. The high-ph conditions within the buffer will also tend to inhibit microbial activity [55]. The radiation field within the supercontainer will be too low to have any significant effect on microbes present within the buffer. The concrete buffer pore water will be irradiated at a rate of ~0.6 Gy hr -1, although the rate will be higher nearer the overpack [11] Effect of Radiolysis The introduction of wastes to the repository will cause the parts of the disposal system nearest to the wastes to be exposed to radiation. However, irradiation rate within the supercontainer will be relatively low and, particularly outside the overpack, the radiation field will be dominated by gamma radiation. As noted above, the average strength of the gamma radiation 86 NIROND-TR E, April 2008

101 field within the buffer will initially be ~0.6 Gy hr -1 [11]. The irradiation rate will decrease with time as a result of radioactive decay and, even at the highest initial irradiation rates, the effects of radiation on the materials of the EBS are not expected to be significant. For example: A review of the effects of radiation on the likely rates of overpack corrosion suggests that irradiation rates would need to be in excess of 3 Gy hr -1 for overpack corrosion rates to be significantly affected by radiolysis [29]. Based on an experimental study of the influence of radiation on the corrosion of carbon steel in artificial groundwaters, Smart and Rance [65] found that a small increase in the corrosion rate of carbon steel at an irradiation rate of 11 Gy hr -1 was sustained for less than one year. These experiments involved steels that had had their surface coating removed by pickling. It is important to remember that the steels of the supercontainer system will not be pickled but, instead, will have surface coatings that will render them less reactive and less prone to corrosion. The effect of radiolysis on corrosion is discussed in more detail in Section 8.4, and the influence of the surface condition of the steel is discussed in Section 8.5. The effects of radiation on buffer concrete mineralogy may include some local disruption of crystalline mineral frameworks, but such effects are not expected to be significant and, indeed, are likely to be counteracted by the effect of elevated temperatures on cement phase mineralogy, which tends to increase the abundance of crystalline phases at the expense of amorphous phases such as CSH gels. The supercontainer will initially contain oxygen in the air trapped within the pores of the concrete buffer. Chemical transport modelling suggests that following saturation of the buffer, the oxygen trapped in buffer pore space will be consumed by corrosion of the carbon steel overpack as fast as it can diffuse from the concrete buffer to the overpack. This will lead rapidly to reducing conditions near the overpack and within the buffer. Several sources indicate that reducing conditions at the overpack and envelope surfaces will be attained after about 1 year (e.g., [46] [41]). Radiolysis has the potential to prolong the period in which more oxidising conditions might persist. Radiolysis calculations indicate that for an initial period after the trapped oxygen has been consumed, the environment at the buffer/overpack interface may be more oxidising than -450 mv she, (the passivation limit for carbon steel see Section 8) [11]. However, these calculations of radiolytic gas production did not take account of corrosion reactions going on in the system. MacDonald et al. [41] consider that although radiolysis could shift corrosion potentials to more positive values (and could therefore reduce the rate of hydrogen evolution), it should not influence the corrosion rate because the passive current densities for both overpack and envelope are independent of potential. Other lines of evidence suggest that the effect of radiolysis of buffer pore water in the supercontainer may not be significant. For example, scoping calculations made to quantify the production of oxidising species as a result of the radiolysis of water, following an approach derived from studies of radiolysis effects in Boiling Water Reactor (BWR) coolant circuits [81], suggest that the effects of radiolysis on the redox conditions in the supercontainer system may only last for a period considerably shorter than 300 years [41] [42]. On the other hand, recent modelling by Bouniol [12] suggests that the concentration of many of the principal radiolytic species derived from oxygen (e.g., H 2 O 2, O 2-2 ) remains fairly constant over 300 NIROND-TR E, April

102 years. ONDRAF/NIRAS has an ongoing experimental and modelling programme to evaluate more closely the potential impact of radiolysis on the corrosion and general chemical evolution of the supercontainer system. Radiolysis within the supercontainer during the operational phase will cause the generation of only small amounts of gas. The hydrogen concentration in buffer pore water due to gamma radiolysis was estimated to be 0.8 mmol l -1 after 100 years in a closed and saturated system [12]. This corresponds to a production rate of mol year -1 [12], which is about 10% of the hydrogen production due to anaerobic corrosion of steel at a corrosion rate of 0.1 µm per year [78]. A more sophisticated analysis has been made by Bouniol [12] that considers radiolysis within the supercontainer under variable temperature and radiation fields. The new analysis has confirmed the earlier results [11] and suggests that radiolysis may be chemically self-regulating, as hydrogen may be removed by reaction with oxidizing radicals. However, there are still uncertainties in the analysis. For example, the coupling of radiolysis with corrosion processes has not yet been taken into account. As noted in Section 7.3.2, radiolysis is expected to have a limiting effect on microbial activity Effect of Elevated Temperatures As discussed in Section 4, the introduction of wastes to the repository will cause the temperature of the disposal system to rise for a period as heat migrates outwards from the wastes. This heating may have several chemical effects: Elevated temperatures in the supercontainer will tend to favour higher rates of chemical reaction for most reactions. Elevated temperatures in the supercontainer will favour the formation within the buffer concrete of mineral phase assemblages that include a higher proportion of crystalline phases, such as afwillite, than would be the case at 25 C. The afwillite-portlandite assemblage dominates the equilibrium phase chemistry from 25 C up to temperatures of about C (Figure 7.3). However, the slow kinetics of afwillite formation means that the CSH gel phase initially present in the buffer concrete may persist for many years. As noted in Section 7.3.2, elevated temperatures and high ph conditions are expected to have a limiting effect on microbial activity [55]. 88 NIROND-TR E, April 2008

103 Figure The CaO-SiO 2-H 2O system at saturated vapour pressure, 80 to 220 C, after Hong and Glasser [33]. Solid lines show phase relationships obtained by reversible reaction. Dashed lines are obtained from monotropic reaction. AFW = afwillite; POR = portlandite. Note that this figure illustrates equilibrium phase assemblages and relationships, and does not show metastable phases such as CSH gel that can co-exist with crystalline phases such as portlandite. The AFW + POR field extends to low temperature (25 C) 7.4 Expected Chemical Evolution and Effects after Repository Closure Backfilling of waste deposition tunnels and closure of the repository will lead to a range of chemical changes and effects. The volume of backfill placed between the supercontainer and the tunnel liner will consist of an annular ring varying between ~20 cm and ~50 cm thick. It is expected that the backfill will become hydrologically saturated within just a few years [78] The Return to Anaerobic Chemical Conditions Redox conditions in the repository tunnels are expected to become anaerobic shortly after the repository tunnels are backfilled and sealed. Redox conditions inside the supercontainer are expected to become anaerobic after the influence of radiolysis has ceased. Oxygen initially trapped within the EDZ, the repository tunnels and the supercontainer will be consumed by corrosion processes and the oxidation of Boom Clay pyrite and organic matter. However, radiolysis will tend to delay the return to anaerobic conditions, particularly in the NIROND-TR E, April

104 environment adjacent to the overpack. Once chemical conditions have become anaerobic, hydrogen gas will be produced from the corrosion of steel. Assuming an initially saturated buffer, MacDonald et al. [41] predicted a transition to anaerobic conditions at the overpack surface in as little as a few weeks, although this period is likely to be extended when radiolysis is taken into account. Other scientific literature on similar systems includes a wider range of predictions of the duration of oxidising conditions, ranging from 0.5 years to several hundred years, although these estimates refer to conditions in the repository as a whole. For example, performance assessments of the Nirex deep disposal concept for low-level and intermediate-level wastes (LILW) in steel and concrete containers surrounded by a concrete backfill assume that the duration of oxidising conditions in the repository after repository closure will be in the range between 1 and 100 years, with a best estimate of 10 years [47]. In the Belgian supercontainer for HLW and spent fuel, there will be greater levels of radiolysis than in the backfill of the Nirex LILW repository. Radiolysis of water produces hydrogen, oxygen and other oxidising radicals and, in the supercontainer environment, these may prolong the existence of mildly oxidising conditions close to the overpack. However, in the supercontainer for both HLW and spent fuel, radiolysis will probably cease at times significantly greater than 300 years, which is ten times the half life of Cs-137 (the main source of gamma radiation in Category C waste) [11]. Corrosion leading to gas production is discussed in further detail in Section 8. ONDRAF/NIRAS has an ongoing experimental and modelling programme to evaluate more closely the potential impact of radiolysis on the evolution of the supercontainer system (e.g., [12]). Gas generation and changes in near-field saturation are discussed in Sections and Effect of Elevated Temperatures The supercontainer will continue to experience elevated temperatures in the early post-closure period and, as described in Section 4, heat will be transferred outwards from the waste and the supercontainer into the Boom Clay through the backfill. Effect of elevated temperatures on the buffer. CSH gel is a major component of cements formed at low temperatures. Even though CSH gel is metastable, particularly at the elevated temperatures expected in the supercontainer, CSH gel is expected to persist for thousands of years because the rates of the crystallisation processes involved in producing phases such as afwillite are relatively slow. Although the more crystalline phases that will form tend to have lower solubilities and lower ph buffering capacities than the corresponding amorphous, lower-temperature gel phases, the persistence of CSH gel and portlandite within the buffer means that the ph of the coexisting pore waters is expected to remain high for thousands of years. The solubility of portlandite is well known and although the solubility of portlandite decreases as temperature rises in the range up to ~100 C, the decrease in solubility is slight and pore water ph levels will remain high 90 NIROND-TR E, April 2008

105 throughout the thermal phase. As noted above, high-temperature cement phases tend to be better crystallised and, therefore, denser than low-temperature phases. As temperatures increase, further crystallisation may occur and this may cause the cement matrix to become increasingly porous. In the supercontainer system, enhanced porosity may provide more space for gas storage. Effect of elevated temperatures on sulphate salt solubilities. The occurrence of sulphate salt precipitation processes both on the liner and, after backfilling, within the cementitious backfill would tend to limit the total amount of sulphur species that reach the supercontainer. After repository closure, hydrological resaturation of the backfill may, however, in conjunction with the elevated temperatures, lead to dissolution of the sulphate salts and re-mobilisation of sulphur species. These sulphur species may then interact with and move through the materials of the EBS. Sulphate movement through the cementitious backfill and buffer will be retarded by ettringite precipitation, whereas there are no obvious retardation mechanisms for sulphide or thiosulphate ions. Effect of elevated temperatures on corrosion. Temperature will have a significant impact on corrosion processes. Corrosion rate increases at higher temperature and all field boundaries in an Eh ph diagram are temperature dependent, as is the passive corrosion current density [41]. Further work to evaluate the influence of temperature increase on corrosion processes is ongoing, including consideration of the influence of temperature on chloride corrosion threshold. Effect of elevated temperatures on Boom Clay. A recent study by Deniau et al. [18] indicates that gases, such as CO 2, may be produced as a result of heating of Boom Clay kerogen. H 2 S may also be produced in small quantities [18]. These and other small organic molecules formed as a result of heating Boom Clay kerogen might also serve as nutrients for microbes Interactions with Envelope Absent or after Envelope Perforation If an envelope is present, corrosion will occur primarily as a result of the access of waterbearing corrosive species to its outer surface following backfill resaturation, a process expected to occur in just a few years (see Section 5.2.1). The inner surface of the envelope will be in close contact with the concrete of the buffer and is, therefore, expected to be passivated; it will be exposed to less corrosive pore waters but may undergo localised corrosion. Envelope corrosion is described in further detail in Section 8.2. When the envelope becomes perforated, or if the supercontainer is fabricated with no envelope, a range of additional chemical interactions may occur: Influx of water and chemical species. Waters from outside the supercontainer may enter the buffer, carrying various chemical species, and these may react with the buffer and its pore waters, and migrate towards the overpack. For example, Wang [76] suggests on the basis of diffusive transport calculations that, if there was early perforation of the envelope, rapid ingress of high-chloride Boom Clay pore water, and no chloride retardation, a chloride concentration of 12 mmol l -1 (400 ppm) might reach the overpack in 400 years. NIROND-TR E, April

106 More conservative scenarios could involve advection of pore fluid through the buffer, due to capillarity and hydraulic saturation, resulting in reduced sulphur species reaching the surface of the overpack within a few years of envelope perforation. Such scenarios would be particularly relevant to EBS evolution if the envelope was absent or initially perforated. Bicarbonate in Boom Clay pore fluids will react with the buffer concrete. However, scoping calculations suggest that the concrete will be effective in buffering ph to high values for as long as 80,000 years. Microbes. Microbes may be sufficiently mobile outside the supercontainer to migrate into the concrete buffer. However, the high ph within the buffer will tend to suppress microbial activity, as will the initially high temperatures. Moreover, the distribution of pore sizes within the buffer may restrict microbes from entering the buffer unless significant cracking occurs. The same effect may prevent significant amounts of colloids, such as Boom Clay humic and fulvic acids, from entering the buffer. These conclusions remain to be confirmed by more detailed information on buffer pore size distribution. Carbonation. Simplified coupled chemical transport modelling performed by Wang [76] suggests that calcite will precipitate in the outer part of the buffer as a result of the chemical reaction between the HCO - 3 waters from the Boom Clay and the more alkaline waters of the concrete buffer. The precipitation of calcite in the outer part of the buffer may lead to a reduction in the porosity of the buffer [76]. Wang [76] suggests that this porosity reduction could be a beneficial effect with respect to the safety of the repository because the sealing effect could reduce diffusion through the affected zone, thus limiting both any perturbation of the far-field by alkaline waters and the migration of radionuclides. However, this sealing effect is unlikely to be complete. Leaching. Some of the more soluble of the species within the buffer, such as sodium and potassium, may be leached from the concrete [76]. If this were to occur to a significant degree, there would be a lowering of the pore water ph conditions within the buffer. However, as noted above, the amount of portlandite present coupled with the low overall potential for significant water fluxes through the system (which might leach Na, K and Ca out of the concrete) mean that high-ph conditions are expected to be maintained within the supercontainer for tens of thousands of years. Gas release. Gas storage within, and release from, the supercontainer are discussed in Section Overpack corrosion. Once aggressive ions from the Boom Clay have entered the supercontainer, they can diffuse through the concrete buffer towards the overpack surface. However, the normal concentration of chloride in the pore waters of the Boom Clay is lower than the critical concentration of chloride necessary for significant corrosion of carbon steel in concrete ( ppm chloride at C under most redox conditions - [37]). However, note that redox conditions, temperature and degree of saturation exert an important influence on corrosion in the presence of chloride ions, and corrosion thresholds will be different if other species are present together with chloride [65] [17]. Further information on corrosion processes is presented in Section NIROND-TR E, April 2008

107 7.5 Main Uncertainties The range of processes that will affect the chemistry of the near-field is extremely complex, and further more detailed modelling and experimental work is underway to further constrain the magnitudes and uncertainties associated with the various processes, the couplings between them, and the rates at which they will occur. Examples of uncertainties include: The coupled chemical transport simulations made thus far have excluded the effect of temperature changes. During the thermal phase there may be a very complex interplay of processes in the EBS, particularly if the envelope is absent, or after it has perforated, with different fluids (water and gas) and different chemical species are moving in opposite directions in response to thermal, hydraulic and chemical potential gradients. It is important to know the degree of saturation in the buffer concrete near the overpack in order to evaluate the tendency for overpack pitting during the initial, short-duration aerobic period. The timing of envelope perforation and the degree of chemical communication thereafter between the outside and inside of the supercontainer is uncertain. It may make the modelling of EBS evolution simpler and less uncertain if the envelope is initially perforated or absent. In this way, the buffer can be assumed to saturate rapidly after closure of the tunnels and there is no need to make assumptions about the timing and nature of envelope perforation. The occurrence and effects of cracking of the buffer concrete are uncertain and difficult to predict and may lead to chemical conditions and transport effects that are distributed spatially in a heterogeneous fashion. The chemical modelling conducted to-date has not considered the potential effects of organic matter within the Boom Clay or the potential effects of superplasticisers and corrosion products from the steel EBS materials. The potential for superplasticisers to generate hydrogen under the effects of radiolysis within the buffer must also be evaluated. NIROND-TR E, April

108 94 NIROND-TR E, April 2008

109 8 Electrochemical Evolution of Metallic Barriers 8.1 Electrochemical Processes Within the EBS Aqueous corrosion is an electrochemical process that involves the oxidation of metal to form metal ions. These ions are initially neither in the aqueous environment nor incorporated into a solid corrosion product. Oxidation is balanced by simultaneous reduction on the surface of the corroding metal. In aqueous systems, the balancing reaction is normally the reduction of oxygen or water [48]. For steel embedded in concrete, the anodic reactions of interest are: + 3Fe + 4H O Fe3O4 + 8H + 8e 2 (1) + 2Fe + 3H O Fe2O3 + 6H + 6e 2 (2) + Fe + 2H 2 O HFeO2 + 3H + 2e (3) 2+ Fe Fe + 2 e (4) The possible cathodic reactions depend on the availability of oxygen and on the ph in the vicinity of the steel surface. The most likely reactions are: 2H O + O2 + 4e 4OH 2 (5) + 2H + 2e H 2 (6) Reactions (1), (2), (3) and (4) represent the dissolution or oxidation of iron and in good quality concrete reactions (1) and (2) are of primary interest. Typically the oxides Fe 3 O 4 and Fe 2 O 3 will form a thin protective passive layer on the surface. The metallic components of the EBS, in particular the carbon steel overpack, will begin to corrode on contact with water or water vapour. Corrosion rates will depend on the prevailing chemical and electrochemical conditions. The rates will also depend on the nature of the corrosion processes that occur (e.g., uniform corrosion or pitting corrosion). In the repository disposal tunnels, after backfilling and sealing, corrosion reactions will contribute to the removal of oxygen introduced to the repository during excavation and operations, and will lead to a rapid return to anaerobic chemical conditions. Once oxygen has been removed, further corrosion can only occur by uniform anaerobic corrosion, unless a sufficient concentration of electroactive species, such as polysulphide or thiosulphate, are present. Corrosion potentials at the envelope and overpack are likely to be sufficiently negative that the risk of pitting corrosion will be very low [41], excepting the possible impact of sulphide, polysulphide and elemental sulphur. Rates of general corrosion are also expected to be low, and the overpack is not expected to be perforated by corrosion for thousands of years, well after the thermal pulse associated with the wastes has declined. NIROND-TR E, April

110 Corrosion products will be formed close to the surfaces of the supercontainer envelope, if present, and at the outer surface of the overpack. These corrosion products may also act as sorption substrates that retard any radionuclides eventually released from the overpack. Anaerobic corrosion of iron and steel liberates hydrogen gas and may raise pore fluid pressure within the buffer porosity while a sealed envelope remains intact. After the envelope is perforated or if it is absent, gas escaping the supercontainer may displace waters in the porosity of the buffer and backfill. Gas production due to gamma radiolysis and microbial activity is likely to be minor compared with hydrogen production due to anaerobic corrosion of the steel EBS components. Further analysis will be required to reduce uncertainties in the calculated gas pressures and to assess the implications for gas migration in the EBS, including potential couplings with hydraulic and mechanical processes. A key factor influencing corrosion processes within the supercontainer concerns the timing of perforation of the envelope, if this is present and initially sealed. Before the envelope is perforated it is likely that hydraulic conditions within the buffer will be partially saturated, especially close to the overpack where heating is most intense. However, under such partially saturated conditions the corrosion risk will be minimal. Preliminary hydraulic calculations indicate that once the envelope is perforated, or if it is absent, the buffer will resaturate rapidly, within just a few years (see above, Section 5). Once the buffer becomes resaturated, it is likely that corrosion reactions at the overpack and envelope surfaces will result in anaerobic conditions being achieved. Although there are steel components within the overpack, such as the stainless steel HLW canisters and carbon steel boxes for spent fuel, there is little risk of significant corrosion of these components prior to overpack failure. This is because the overpack will be evacuated prior to sealing, will contain only trace amounts of water, and in the case of spent fuel, the carbon steel boxes will be filled with sand or an inert gas. Once the overpack has failed and buffer pore fluids enter, anaerobic, high-ph corrosion of the boxes and stainless steel HLW canisters can begin to occur. Such corrosion of metals contained within the overpack, particularly the carbon steel boxes, will lead to additional gas generation. However, overpack failure is not likely until at least several thousand years after closure of the repository. 8.2 Corrosion of the Envelope It is assumed that if present, the outer supercontainer envelope will be manufactured from a low-carbon stainless steel with an enhanced Mo content (AISI 316L hmo) as recommended by the ONDRAF/NIRAS Corrosion Panel [50]. Stainless steel, particularly the grade recommended, has good corrosion resistance. It benefits from extremely low uniform corrosion rates in atmospheric environments and is highly resistant to corrosion in cement-based environments [48]. The presence of a passive film on stainless steel restricts uniform corrosion over the whole surface. Ambient temperature atmospheric corrosion only occurs if a thin layer of moisture is present, and corrosion rates in atmospheric environments are expected to be well below 0.1 µm per year under controlled conditions [65]. Breakdown of the passive film in small areas can 96 NIROND-TR E, April 2008

111 lead to localised corrosion. Such localised corrosion on a free surface will most likely take the form of small depressions or pits (pitting corrosion). The most common cause of localised corrosion is the presence of salts, such as chloride, which are able to disrupt the passive film and prevent it reforming. Localised corrosion is especially a problem under aerobic conditions [65]. Future ONDRAF/NIRAS studies are seeking to constrain the likely electrochemical evolution of the envelope, in particular the likelihood of external pitting in response to the presence of electroactive species such as polysulphides or thiosulphates at the surface of the envelope after backfill resaturation (see Table 3.1). 8.3 Corrosion of the Overpack The overpack will surround the vitrified HLW canisters or spent fuel assemblies and will provide the main barrier to radionuclide release during the thermal phase. Based on recommendations of the ONDRAF/NIRAS corrosion panel [50], the overpack will be manufactured from carbon steel [5]. The surface of the overpack will be in contact with the supercontainer filler and concrete buffer and will therefore be subjected to a high-ph environment from the time of supercontainer assembly onwards. Kursten [40] has reviewed uniform corrosion data pertaining to carbon steel and mild steel corrosion in high-ph environments. Data from corrosion measurements in aerobic and anaerobic solutions were reviewed as well as data from the corrosion of carbon steel in highph cementitious environments: In aerobic solutions, uniform corrosion rates in the range 1 to 10 µm per year have typically been reported at ambient temperature, and corrosion rate tends to increase at higher temperature and higher chloride concentration ([40], Table 1). Miserque et al. [44] and Féron [21] have recently carried out carbon steel immersion tests under aerobic conditions in the absence of aggressive species. These experiments indicated a rate of growth of the corrosion layer of less than 0.01 µm per year based on XPS measurement of layer thickness. This rate cannot necessarily be directly translated into a carbon steel corrosion rate because during passive dissolution, corrosion of the metal surface and dissolution of the corrosion layer both occur simultaneously. In anaerobic solutions, the most relevant data are those generated over very long periods by hydrogen evolution experiments ([40], Table 2). Experiments show that the corrosion rate may fall as low as 0.07 nm/year at 20 C [65]. This is a reflection of the low solubility of iron oxide at high ph and the low concentration of H + at high ph. Repassivation of carbon steel has been reported even in the presence of chloride under anaerobic conditions [65]. In aerobic cementitious material, a range of corrosion rates has been observed, depending on porosity, cement type, ph, chloride concentration and humidity. In the absence of chloride and before carbonation of the cement has occurred, corrosion rates may be as low as 0.08 µm/year ([40], Table 4). NIROND-TR E, April

112 In anaerobic concrete without added chloride, corrosion rates in the range 0.23 to 1.44 µm/year have been observed, but these are believed to represent maximum values ([40], Table 6). Kursten [40] summarises approaches that have been taken to quantify the variation of uniform corrosion rate with time. The experimental techniques have involved the measurement of weight loss on test coupons, electrochemical measurement techniques, or the measurement of hydrogen gas evolution. The main disadvantage of weight loss measurements is that these experiments only provide an indication of average corrosion rate over long time periods instead of instantaneous corrosion rates, and are also prone to weighing errors. Electrochemical techniques are not particularly well suited for the measurement of low anaerobic corrosion rates over long timescales (>1 year) since results can be affected by oxygen ingress through electrical connections to the test cell and tend to have detection limits >0.1 µm/year. The best technique for measuring steel corrosion rates in simulated repository environments are those involving measurement of the volume of hydrogen produced as a result of the corrosion process [40]. The rate of uniform corrosion generally decreases with time either because of the depletion of reactants or, more usually, because of the formation and growth of a protective corrosion product film [36]. Observation of the corrosion of iron at low potentials contains strong evidence that ferrous hydroxide (Fe(OH) 2 ) and magnetite (Fe 3 O 4 ) surface films thicken until their growth rate is balanced by their rate of dissolution at the oxide/electrolyte interface. In such circumstances, the metal dissolution rate is determined by the rate of diffusion of metal ions from the metal interface. In the absence of dissolution, film thickening will decrease with time to a negligible rate [64]. A general trend of uniform corrosion rate decreasing with increasing exposure time has been observed by many researchers when carbon steel has been exposed to alkaline media, representative of the environment surrounding the carbon steel overpack within the supercontainer. Smart et al [65] used gas measurements to monitor the instantaneous uniform corrosion rate of carbon steel in aqueous alkaline environments (saturated Ca(OH) 2, 0.1M NaOH), and also in cementitious environments (Nirex Reference Vault Backfill). The experiments also considered various temperatures (30 C, 50 C, 80 C), and chloride concentrations over a period of ten years. For all tests carried out on carbon steel, the corrosion rate was observed to decrease with increasing exposure time (see Figure 8.1). The decrease in the corrosion rate was associated with the development of a protective magnetite film on the surface of the carbon steel samples. 98 NIROND-TR E, April 2008

113 Figure Anaerobic corrosion rate of carbon steel as a function of time, derived from volumetric gas measurements in various alkaline media at 50 C [65]. Decreasing uniform corrosion rate in alkaline solutions and cementitious environments has also been reported [22]. The evolution of the uniform corrosion rate of carbon steel in saturated Ca(OH) 2 at 35 C was monitored by measuring the evolved hydrogen and showed progressive decrease to low levels over several hundred days. Kreis [38] [39], Simpson et al. [63], and Grauer et al. [27] have also observed the rate of corrosion on mild steel in anaerobic alkaline conditions and detected a decrease over similar periods. Andrade and Gonzalez [1] and Gonzalez et al. [25] have carried out a range of experiments at ambient temperature to determine the corrosion rate of mild steel in OPC mortar using a linear polarisation resistance technique. They recorded a pronounced decrease in corrosion rate as a function of time, followed by a rise when the cement had eventually become carbonated. Based on these and similar observations, Gras [26] has suggested that the variation in the rate of uniform corrosion reactions obeys a power law of the form X = kt -n (where X is corrosion rate, t is time, and k and n are constants) representing the slowing of reaction rate over time as a protective film develops on the metal surface. A similar power law was observed by Dridi [20] to fit the variation of corrosion rate with time, compiled from uniform corrosion rate data for carbon steel exposed to argillaceous environments and reducing conditions over 10 years. Crossland [14] has compiled corrosion data from both laboratory experiments and archaeological artefacts up to 1,700 years old (Figure 8.2) to suggest that individual datasets all appear to follow parabolic rate law kinetics (X = kt -0.5 ). The effect of soil type is to change the parabolic rate constant, k. NIROND-TR E, April

114 Figure Corrosion rates of carbon steel and cast iron in soil [14] Miserque et al. [44] investigated the surface layer developed on mild steel grade FeE500 immersed in 0.1 M NaOH. They observed that the oxide layer reached a constant thickness of less than 10 nm. This stationary behaviour was interpreted to indicate that a combined corrosion/dissolution mechanism was occurring simultaneously at the interface. The evolution towards a stationary thickness was interpreted to indicate that the effective rate of corrosion reduced with time. Smart et al. [65] investigated the stability of passive films formed on carbon steel using techniques in which steel surfaces were mechanically depassivated in the test solutions by scratching, cutting or abrasion. The rate with which the passive film reformed was monitored electrochemically. It was observed that after depassivation, carbon steel surfaces were able to repassivate rapidly in deaerated alkaline conditions (0.1 M NaOH), both in the absence and presence of chloride ions. Isotopic tracers were used to demonstrate that the passive film was formed by direct reaction between the metal and water. Further experiments in 0.5 M NaCl solutions showed that chloride ions tend to reduce the rate of repassivation and to react with existing passive surface films, but the presence of chloride does not prevent the formation of a passive film. These observations imply that the carbon steel overpack surrounded by an anaerobic cement-based buffer would maintain a protective passive film indefinitely, leading to low corrosion rates. From the data summarised in Section 8.3 it is clear that the uniform corrosion rate of carbon steel under disposal system conditions (a high-ph cementitious environment and long-term reducing conditions) will tend to decrease with time and within a few years reach a very low constant value of less than 0.1 µm per year. Such a conclusion is supported by numerous experimental and natural observations. 100 NIROND-TR E, April 2008

115 As noted above, once oxygen has been consumed, further corrosion of the overpack can only occur by uniform anaerobic corrosion, unless a sufficient concentration of electroactive species are present. Corrosion potentials at the envelope and overpack are likely to be sufficiently negative that the risk of pitting corrosion will be very low [41], excepting the possible impact of sulphide, polysulphide and elemental sulphur. It should be noted, however, that radiolysis can generate oxidising species that could increase the corrosion potential of metals. 8.4 Effect of Radiation on Anaerobic Corrosion of the Overpack As described in Section 7.3.3, the surface of the overpack will be subjected to a radiation field that could initially be as high as ~ Gy s -1 (~22 Gy h -1 ). Radiolysis could influence the electrochemical environment at the overpack surface by generating oxidising species. However, Kursten [40] found little information in the literature relevant to the effect of radiolysis on the anaerobic corrosion rate. Moreover, the available data is somewhat contradictory: Smart et al. [65] reported corrosion potential values for carbon steel embedded in various deaerated cement mixes in the range of -685 mv SHE to -750 mv SHE. These values are close to or below the hydrogen equilibrium potential (-737 mv SHE at ph 12.5 and 1 atmosphere hydrogen). This indicates that in the absence of radiolysis, hydrogen evolution would be possible at the potentials observed for carbon steel embedded in cement. Macdonald et al. [42] used the BWR (Boiling Water Reactor) radiolysis code FOCUS, modified to allow variations of ph, to calculate the concentration of O 2, H 2 and H 2 O 2 at the overpack surface. A mixed potential model was used to calculate the corrosion potential for irradiation rates in the range Gy h -1 and various temperatures (30, 60 and 90 C). Results predict that at the expected initial irradiation rate at the overpack/concrete interface (~25 Gy h -1 ), the corrosion potential is expected to vary from about -300 mv SHE (30 C) to about -350 mv SHE (90 C). These values are well above the hydrogen equilibrium potential indicating that hydrogen evolution could not occur. However, Macdonald et al. [42] consider that the effects of radiolysis on the redox conditions in the supercontainer system may last for a period considerably shorter than 300 years. Macdonald et al. [42] also concluded that when the decrease in temperature and irradiation rate with time is taken into account, the corrosion potential may increase from about -350 mv SHE at 90 C to -260 mv SHE at 30 C. The susceptibility of carbon steel to localised corrosion can be assessed by comparing these values with the repassivation potential. Pourbaix and L'Hostis [57] have suggested that -300 to -350 mv SHE is a conservative estimate of the repassivation potential for carbon steel in passive systems such as concrete, where chloride is the main aggressive species. The calculated corrosion potential values of MacDonald et al [42] are generally equal to or slightly more positive than the repassivation potential showing that the possible occurrence of localised corrosion cannot be excluded. Bouniol [12] conducted radiolysis simulations examining the evolution of radiolysis species assuming an initial gamma irradiation rate of 25 Gy h -1 at the overpack surface. These calculations predict a high and almost constant oxygen concentration of mol litre -1 at the overpack/concrete buffer interface over a 300-year period. However, no attempt has yet been made to couple these radiolysis calculations with corrosion reactions. NIROND-TR E, April

116 Smart and Rance [66] found that a small increase in the corrosion rate of pickled carbon steel at an irradiation rate of 11 Gy hr -1 was sustained for less than one year. However, for radiation doses up to 300 Gy hr -1 and a ph range of 9 to 10, they reported the production of hydrogen by anaerobic corrosion. This indicates that low corrosion potentials were established, even in an environment with a much higher irradiation rate than that expected in the supercontainer. The impact of radiolysis on the anaerobic corrosion of carbon steel in alkaline media, representative of the concrete buffer environment, forms part of an ONDRAF/NIRAS experimental programme to be carried out over the period Main Uncertainties A number of factors influence the rate of overpack corrosion and contribute uncertainty to the long-term corrosion estimate: The initial surface condition of the carbon steel overpack will influence the subsequent corrosion behaviour. Most published corrosion rate data refer to the corrosion behaviour of freshly prepared metal surfaces. However, it is more realistic to assume that, under disposal conditions, an aerobically formed corrosion product layer will be present on the surface of the steel before anaerobic corrosion commences. It is therefore important to determine how a corrosion film formed under one set of conditions (e.g. aerobic) affect the corrosion behaviour under a subsequent set of conditions (e.g. anaerobic) [65]. The effect of corrosion product layers on the subsequent corrosion behaviour of carbon steel has been studied by Naish et al. [45] who suggest that the presence of aerobically-formed corrosion product films will delay, but not prevent the onset of anaerobic corrosion of carbon steel. Further uncertainties relate to the exact method of fabrication and assembly of the supercontainer, and the type and mode of emplacement of the filler, both of which may influence the initial nature of the overpack surface. However, regardless of the initial nature of the overpack surface, long-term, uniform corrosion rates will evolve towards similar low values, as outlined in Section 8.3. The initial degree of saturation of the concrete buffer, and the speed with which it resaturates, will influence the speed with which conditions at the overpack surface become anaerobic. Buffer resaturation is related to whether an envelope is present, and if it is present, whether it is sealed or perforated. The rate of corrosion reactions is influenced by the temperature and in general, the higher the temperature the faster a low uniform corrosion rate is reached. However, a detailed understanding of corrosion kinetics at the overpack surface and the exact time of transition to a low uniform corrosion rate is not yet available. The concentration of chloride and reduced sulphur species at the overpack surface will influence the corrosion behaviour. The likely concentration of aggressive species at the interface between the Boom Clay and tunnel liner is being determined in greater detail, and this information is being used to estimate the likely concentration of chloride and reduced sulphur at the overpack surface. 102 NIROND-TR E, April 2008

117 There is significant variation in the experimentally measured uniform corrosion rate of carbon steel in aerobic solutions or aerobic concrete, and this makes it difficult to specify a single threshold value in the case of the overpack [40]. However, in the supercontainer, the aerobic period is likely to be short and the contribution of aerobic uniform corrosion correspondingly small. The likelihood of localised corrosion of the overpack during the early aerobic phase is uncertain. An ongoing ONDRAF/NIRAS study is seeking to address this issue. The influence of radiolysis on the electrochemical conditions at the overpack surface is uncertain. No theoretical study of coupled radiolysis corrosion behaviour has yet been undertaken, although the impact of radiolysis on the anaerobic corrosion of carbon steel in alkaline media forms part of an ongoing ONDRAF/NIRAS experimental study. Further research and development by ONDRAF/NIRAS is seeking to reduce these uncertainties as the SFC-1 programme moves forward. NIROND-TR E, April

118 104 NIROND-TR E, April 2008

119 9 Conclusions The work summarised in this report leads to the following conclusions concerning the expected evolution of the near-field in the Belgian concept for disposal of Category C wastes: Thermal Evolution The thermal evolution of the near-field will depend on the waste inventory, the cooling time before supercontainer assembly and disposal, and the thermal parameters of the EBS and Boom Clay host formation. The maximum temperature at the overpack surface may be limited by the cooling time, the time between waste formation and supercontainer assembly. A cooling time of a minimum of 60 years is likely to be necessary for maximum overpack surface temperatures to remain below 100 C for HLW and spent fuel. Calculated maximum temperatures are similar for both waste types, although generally a few degrees higher for spent fuel. The spatial thermal gradients generated within the buffer are also similar for both waste types. However, when the waste is spent fuel, the temperature remains high for a longer period than in the case of vitrified HLW. Based on preliminary modelling using generic input parameters, and assuming a 60-year cooling time, peak EBS temperatures generally occur in the first 5 to 15 years after waste emplacement and are not significantly influenced by heat from surrounding tunnels. For vitrified HLW, the influence from two closest neighbouring tunnels is +~7 C after 50 years, and for spent fuel the effect is negligible. However, the thermal pulse from neighbouring tunnels arrives too late to have any effect on peak EBS temperatures. For both waste types, the buffer will experience peak temperatures that are typically in the range C assuming a 60-year cooling time, whereas the Boom Clay adjacent to the tunnel liner will experience peak temperatures in the range C. Sensitivity studies have shown that the thermal evolution of the EBS is particularly sensitive to the influence of the thermal conductivity of the Boom Clay. In future it will be important that the value of this parameter is determined with the best precision possible in order that accurate predictions of thermal evolution can be made. Hydraulic Evolution The hydraulic evolution of the repository has been constrained by evaluating the time needed for the backfill and buffer to saturate after tunnel closure. The timescale is mainly dependent on the unsaturated behaviour of the EBS materials and Boom Clay, porosity of the backfill, the hydraulic conductivity of the Boom Clay and the hydraulic gradient. A key uncertainty relates to assumptions regarding the unsaturated material characteristics, as well as the initial saturation state of EBS materials. A further uncertainty relates to whether or not a sealed envelope is present, and to the timing of sealed envelope perforation, and the subsequent degree of hydraulic communication between the inside and outside of the supercontainer after perforation occurs. Based on preliminary scoping calculations, the following conclusions are drawn: NIROND-TR E, April

120 After backfilling and closure of the repository, the backfill will saturate rapidly, within a few years. The timescale depends primarily on the hydraulic conductivity of the Boom Clay. Equilibration of the pressure heads will take up to 100 years. After envelope perforation or in the absence of an envelope, buffer saturation will also occur within a few years. If the envelope is absent or designed with pre-existing perforations, the entire EBS up to the outside surface of the overpack will become saturated within a few years of backfilling and sealing of the disposal tunnels. If the envelope is initially sealed, the timing of perforation is highly uncertain and this introduces uncertainty into the hydraulic evolution of the buffer. In order to improve confidence in these calculations, there should be better characterization of the unsaturated behaviour of the near-field materials, both governing saturation processes and governing the effects of gas generation and two-phase flow. In order to evaluate changes to the content and spatial distribution of pore water in the supercontainer concrete buffer during heating, and the potential water vapour pressures generated, the hydrothermal evolution of the supercontainer buffer was modelled. The main conclusion was: Small amounts of bound water may be released from the supercontainer concrete buffer as a result of heating by the waste. However, this is only likely to lead to modest changes in saturation state and water vapour pressure within the buffer pore space and will not generate a dry zone close to the overpack. The modelled evolution is similar for both a concrete buffer cured with no water loss and a buffer that is partly dried before supercontainer assembly. Scoping calculations on gas generation and migration demonstrate that diffusive removal of hydrogen in pore fluid is not rapid enough to disperse the hydrogen produced by anaerobic corrosion of the envelope and overpack. These calculations suggest that a free gas phase is likely to develop within the supercontainer. However, it must be remembered that calculations to date have been based on an unrealistically high overpack corrosion rate. Mechanical Evolution The mechanical evolution of the EBS is influenced by the disturbance to the regional stress field caused by excavation of the repository. Other relevant mechanical effects derive from couplings with thermal or hydraulic processes. For example, heat from the waste, pore pressure changes during saturation of the backfill and buffer, and chemical processes leading to gas generation may all drive mechanical changes within the EBS and the immediately surrounding rocks including: Thermal expansion and contraction of EBS components leading to tensile and compressional stresses. Volume and porosity changes associated with the development of corrosion products at the outer surface of the overpack, and to a lesser extent, the envelope. 106 NIROND-TR E, April 2008

121 Volume and porosity changes associated with mineralogical changes within the buffer concrete, backfill and tunnel liner components of the EBS. Cracking of the buffer concrete in response to thermal stresses and gas generation. Fracturing or disturbance of the Boom Clay resulting from high gas pressure, suction during near-field resaturation, or thermal effects. A 2D analysis of the buffer stress evolution caused by thermal loading indicates that calculated stress levels throughout the buffer, both in tension and compression, are unlikely to cause cracking. However, stress levels in the buffer at the ends of the supercontainer have yet to be evaluated. The mechanical effects of corrosion have also yet to be evaluated. Chemical and Electrochemical Evolution Understanding the chemical and electrochemical evolution of the EBS is complex but essential to the demonstration of the safety of the repository system. In particular, the containment safety function relies on there being no releases from the overpack during the thermal period, and this relies on the maintenance of a high-ph chemical environment at the overpack surface. The following concluding remarks are made: Oxidising effects are unavoidable during repository construction and operation and mostly involve oxidation of Boom Clay pyrite and organic matter by introduced air. Scoping calculations and experimental data suggest that during repository operation, oxidising conditions are not likely to develop more than a metre into the host clay. For the irradiation rates expected within the supercontainer, the possible effects of radiation on the materials of the EBS, on corrosion, on buffer mineralogy, and on gas production are not expected to be significant. Further evaluation of the influence of radiolysis on redox evolution is in progress. Scoping calculations suggest that radiolysis could prolong the duration of the aerobic phase within the supercontainer. The elevated temperatures that will be experienced by the concrete buffer will cause the mineralogy to evolve towards an assemblage dominated by portlandite and afwillite. This assemblage will buffer ph to values in excess of 12 (as measured at 25 C) for tens of thousands of years. Assuming rapid saturation of backfill and buffer, anaerobic conditions will likely prevail throughout the entire EBS soon after repository closure but will remain poorly poised in the concrete barriers. A critical uncertainty in the development of redox conditions in the supercontainer is whether the envelope will be sealed, initially perforated or absent. o o If it is initially perforated or absent, both buffer and backfill will probably saturate within a few years of tunnel sealing and anaerobic conditions will soon prevail everywhere. If it is initially sealed, the evolution of saturation state in the concrete pore space of the buffer will be complex, particularly in the hottest region close to the overpack. Depending on the initial saturation state, corrosion may still occur and lead to anaerobic conditions, but on an uncertain timescale. NIROND-TR E, April

122 Following perforation, the buffer will saturate rapidly. Future research will need to reduce the uncertainties associated with supercontainer evolution involving a sealed envelope and an unsaturated buffer. In the absence of an envelope, or following envelope perforation, waters from outside the supercontainer may enter the buffer carrying various chemical species, and these may react with the buffer and its pore waters, and migrate towards the overpack. Review of the scientific literature suggests that the uniform corrosion rate of carbon steel under disposal system conditions (a high-ph cementitious environment and long-term reducing conditions) will tend to decrease with time and within a few years reach a very low constant value of less than 0.1 µm per year. This conclusion is supported by numerous experimental and natural observations. Preliminary geochemical modelling suggests that the concentration of chloride ions at the overpack surface may remain too low throughout the thermal period to threaten overpack integrity by enhancement of corrosion processes. Further modelling is attempting to provide additional constraints on chloride evolution. The potential impact of reduced sulphur species to enhance localised overpack corrosion has yet to be evaluated in detail. External fluids entering the buffer will react with the concrete mineral assemblage, ultimately converting all of the portlandite to calcite. However, this process is expected to be diffusion-controlled, and is not predicted to be complete for many tens of thousands of years or longer. Precipitation of calcite may reduce buffer porosity and further limit chemical transport, thereby preserving portlandite in the internal part of the buffer for longer. Microbial activity will be suppressed in the EBS, mainly due to high temperature and high ph. The range of processes that will affect the chemistry of the near-field is extremely complex, and further more detailed modelling and experimental work is required in order to constrain the magnitudes and uncertainties associated with the various processes. These uncertainties include: o o o The coupled chemical transport simulations made thus far have excluded the effects of temperature changes and the presence of corrosion products. There may be a very complex interplay of processes in the EBS, both before and after the envelope has perforated, and even if it is absent, with different fluids (water and gas) and different chemical species moving in opposite directions in response to hydraulic and chemical potential gradients. Further complexities to the chemistry of the system may be introduced by the inclusion of superplasticisers, and the final selection of superplasticiser in the supercontainer concrete specification will need to take into consideration the implications for the chemical evolution. The occurrence and effects of cracking of the buffer concrete are uncertain and difficult to predict and may lead to chemical conditions and effects that are distributed spatially in a heterogeneous fashion. A preliminary study of 108 NIROND-TR E, April 2008

123 the susceptibility of the buffer to cracking has been carried out and will be developed further [5]. However, the main importance of the buffer is to condition the chemical composition of the supercontainer pore fluid, and this function is expected to be fulfilled regardless of the occurrence of cracking. NIROND-TR E, April

124 10 References [1] Andrade C., González J.A., Quantitative Measurements of Corrosion Rate of Reinforcing Steels Embedded in Concrete Using Polarisation Resistance Measurements, Werkstoffe und Korrosion 29, pp , [2] Baeyens, B., Maes, A., Cremers, A., Henrion, P. N., Aging effects in Boom Clay, Radioactive Waste Management and the Nuclear Fuel Cycle, 6, pp , [3] Baeyens, B., Maes, A., Cremers, A., Henrion, P. N., In situ physicochemical characterization of Boom Clay, Radioactive Waste Management and the Nuclear Fuel Cycle, 6, pp , [4] Belgatom. Design for Overpacks for Vitrified Waste Canisters and Spent Fuel Assemblies. Belgatom Report /950 2, [5] Belgatom. Demonstration of the Supercontainer Design for Spent Fuel Waste Disposal. Stage 1 Concrete Characterization Interim Report, [6] Belgatom, Demonstration of the Supercontainer Design for Spent Fuel Waste Disposal. Stage 2 Concrete Characterization Adaptation and large scale tests, 2007 [7] BENIPA, Numerical Experiments for the Verification of Bentonite Barrier Assessment Models - Deliverable D-7 (Revision 3). Bentonite Barriers in Integrated Performance Assessment - BENIPA Project EC Contract Nº FIKW , [8] Bernier, F., Li, X.L., Bastiaens, W., Ortiz, L., Van Geet, M., Wouters, L., Frieg, B., Blümling, P., Desrues, J., Viaggiani, G., Coll, C., Chanchole, S., De Greef, V., Hamza, R., Malinsky, L., Vervoort, A., Vanbrabant, Y., Debecker, B., Verstraelen, J., Govaerts, A., Wevers, M., Labiouse, V., Escoffier, S., Mathier, J-F., Gastaldo, L., Bühler, Ch., Fractures and Self-healing Within the Excavation Disturbed Zone in Clays (SELFRAC). European Commission Report, EUR22585, p. 62, 2007 [9] Bogue, R. H., The Chemistry of Portland Cements. 2nd ed., Reinhold Publishing, New York, [10] Bouniol, P., Bétons speciaux de protection. CEA SACLAY, BN 3740, [11] Bouniol, P., Evaluation du Terme Source Hydrogene par Simulation de la Radiolyse dans le Beton d un Superconteneur Comportant deux Colis Primaires. CEA Report RT DPC/SCCME A, [12] Bouniol, P., Radiolysis in the concrete of a supercontainer including 2 primary waste forms. Simulation at the concrete/steel interface for a variable temperature. CEA Report RT DPC/SCCME/ A, NIROND-TR E, April 2008

125 [13] Crossland, I., Long-Term Corrosion of Iron and Copper. ICEM 05 Conference, Glasgow, Scotland, [14] Crossland, I., Corrosion of Iron-Based Alloys Evidence from Nature and Archaeology, Report No. Crossland Report CCL/2006/02, Nirex Ltd, Harwell, UK, [15] De Craen, M., Wang, L., Weetjens E., Natural evidence on the long-term behaviour of trace elements and radionuclides in the Boom Clay, SCK CEN Final report to NIRAS/ONDRAF for the period , Contract nr. CCHO /00/00, KNT R-3926, Mol, Belgium, [16] De Craen, M., Wang, L., Van Geet, M. and Moors, H., Geochemistry of Boom Clay pore water at the Mol site, SCK-CEN, Scientific report BLG-990 (04/MDC/P-48), [17] De Viedma, P.G., Castellote, M. And Andrade, C. Comparison between several methods for determining the depassivation threshold value for corrosion onset. In: Corrosion and Long Term Performance of Concrete in NPP and Waste Facilities (Ed. V. L'Hostis, F. Foct and D. Féron) J. Phys. IV France 136, 79-88, [18] Deniau, I., Derenne, S., Beaucaire, C., Pitsch, H., and Largeau, C., Simulation of thermal stress influence on the Boom Clay kerogen (Oligocene, Belgium) in relation to long-term storage of high activity nuclear waste I. Study of generated soluble compounds, Applied Geochemistry, 20, , [19] Dierckx, A., Cool, W., Lalieux, P., De Preter, P. and Smith, P. The ONDRAF/NIRAS Safety Strategy for the Disposal of Category B and C Wastes. International Symposium on the Safety Case, January 2007, NEA, Paris, [20] Dridi, W., (2005). Couplage entre corrosion et comportement diphasique dans un milieu poreux: Application à l'évolution d'un stockage des déchets radioactifs, thèse présentée pour l'obtention du diplôme de docteur de l'école Nationale des Ponts et Chaussées. [21] Féron, D., Corrosion Issues in Nuclear Waste Disposal. Third International Workshop on Long Term Prediction of Corrosion Damage in Nuclear Waste Systems. May , The Pennsylvania State University State College, [22] Fujiwara, A., Yasutomi, I., Fukudome, K., Tateishi, T. Fujiwara, K., Influence of Oxygen Concentration and Alkalinity on the Hydrogen Gas Generation by Corrosion of Carbon Steel, Mat. Res. Soc. Symp. Proc. 663, pp , [23] Fukunaga, S., Yoshikawa, H., Fujiki, K., Asano, H., Experimental Investigation of the Active Range of Sulphate-Reducing Bacteria for Geological Disposal. Mat. Res. Soc. Symp. Proc. Vol 353, p , NIROND-TR E, April

126 [24] Gens, R., Bel, J., Pourbaix, A., Hélie, M. Wickham, S.M., Bennett, D.G., Corrosion processes and the expected evolution of the BSC-1 Supercontainer design for disposal of Belgian HLW and spent fuel. NUCPERF, March [25] González, J.A., Algaba, S., Andrade, C., Corrosion of Reinforcing Bars in Carbonated Concrete, Br. Corros. J. 15(3), pp , [26] Gras, J.-M., Résistance à la Corrosion des Matériaux de Conteneur Envisagés pour le Stockage Profond des Déchets Nucléaires 1ère partie: Aciers non ou faiblement alliés, fontes, Report No. EDF HT-40/96/002/A, Electricité de France, Moret-sur- Loing, France, [27] Grauer, R., Knecht, B. Kreis, P. Simpson, J.P., The Long Term Corrosion Rate of Passive Iron in Anaerobic Alkaline Solutions, Werkstoffe und Korrosion 42, p , [28] GSL, Belgian Supercontainer Design for HLW and Spent Fuel Disposal: Evaluation of the Reference Design. Galson Sciences Report , [29] GSL, Effect of Gamma-Radiolysis on Corrosion of Steel Containers for Spent Fuel and High-Level Radioactive Waste Disposal. Galson Sciences Report , [30] GSL, Requirements Management in the ONDRAF/NIRAS Programme. Galson Sciences Report , [31] Henrion, P.N., Monsecour, M., Fonteyne, A., Put, M., De Regge, P. Migration of radionuclides in Boom Clay. In: Radioactive Waste Management and the Nuclear Fuel Cycle, Harwood Academic Publishers, New York, USA, Vol.6, pp , [32] Herbert, B.N., The Potential for Microbial Activity Affecting the Integrity of Buried High Level Radioactive Waste in the Boom Clay at Mol, Belgium. Unpublished ONDRAF/NIRAS Report, 2002 [33] Hong S.-Y., Glasser. F.P., Phase Relations in the CaO-SiO2-H2O system to 200 C at Saturated Steam Pressure. Cement and Concrete Research, 34, pp , 2004 [34] IAEA, The principles of radioactive waste management. IAEA Safety Series No F, Vienna, [35] ICRP, Radiation Protection Recommendations as Applied to the Disposal of Long Lived Solid Radioactive Waste. Annals of the ICRP. ICRP Publication 81, Pergamon Press., [36] King F., Overview of a Carbon Steel Container Corrosion Model for a Deep Geological Repository in Sedimentary Rock, Report No. NWMO TR , Nuclear Waste Management Organization, Toronto, Canada, NIROND-TR E, April 2008

127 [37] Kissel, J., Pourbaix, A., Les effets combinés de la teneur en chlorure et de l'alcalinité des bétons sur la corrosion de l'acier. Rapports Techniques CEBELCOR, Vol.165, RT.315, [38] Kreis, P., Hydrogen Evolution from Corrosion of Iron and Steel in Low/Intermediate Level Waste Repositories, Report No. NAGRA NTB 91-21, NAGRA, Wettingen, Switzerland, [39] Kreis, P., Simpson, J.P., Hydrogen Gas Generation from the Corrosion of Iron in Cementitious Environments, in Corrosion problems Related to Nuclear Waste Disposal, European Federation of Corrosion Publication number 7, published by Institute of Materials, London, UK, 1992 [40] Kursten, B., Uniform Corrosion Rate Data of Carbon Steel in Cementitious Environments Relevant to the Supercontainer Design Best Estimate from Literature Data. SCK CEN External Report, [41] MacDonald, D.D., Urquidi-MacDonald, M., Engelhardt, G.R., Simulation of Hydrogen Production in the Annulus of a Supercontainer for the Disposal of High Level Nuclear Waste in a Belgian Boom Clay Repository. Report to ONDRAF/NIRAS, [42] MacDonald, D.D., Urquidi-MacDonald, M., Engelhardt, G.R., Simulation of Hydrogen Production in the Annulus of a Supercontainer. Powerpoint presentation, Brussels, 17 January [43] Miller, W., Alexander, R., Chapman, N. McKinley, I., Smellie, J., Geological disposal of radioactive wastes and natural analogues. Elsevier Science Waste Management Series, Volume 2. ISBN , [44] Miserque, F., Huet, B., Bendjaballah, D., Azou, G., L Hostis, V., X-Ray Photoelectric Spectroscopy and Electrochemical Studies of Mild Steel FeE500 Passivation in Concrete Simulated Water. NUCPERF, March [45] Naish C.C., Blackwood D.J., Thomas M.I., Rance A.P., The Anaerobic Corrosion of Carbon Steel and Strainless Steel, Report No. AEAT/R/ENV/0224, AEA Technology, Harwell, UK, 2001 [46] Nirex, The Evolution of the Eh in the Pore Water of a Radioactive Waste Repository. Nirex Report NSS/R308, [47] Nirex, Nirex 97: An Assessment of the Post-Closure Performance of a Deep Waste Disposal facility at Sellafield. United Kingdom Nirex Limited, Science Report S/97/012, [48] Nirex, Corrosion resistance of stainless steel radioactive waste packages. Nirex Report N/110, 2004 NIROND-TR E, April

128 [49] ONDRAF/NIRAS, Safety Assessment and Feasibility Interim Report 2. ONDRAF Report Ref: NIROND F, [50] ONDRAF/NIRAS. A Review of Corrosion and Material Selection Issues Pertinent to Underground Disposal of Highly Active Nuclear Waste in Belgium. ONDRAF Report Ref: NIROND , [51] ONDRAF/NIRAS, Multi-Criteria Analysis on the Selection of a Reference EBS Design for Vitrified High Level Waste. ONDRAF Report Ref: NIROND , 2004b. [52] ONDRAF/NIRAS, Working group "Repository Conceptual Design" - GTA. Minutes of Meeting June 13th ONDRAF/NIRAS note , 2005 [53] ONDRAF/NIRAS. Long-term safety functions within the disposal programmes of ONDRAF/NIRAS. ONDRAF/NIRAS Report, 9 June [54] ONDRAF/NIRAS, The ONDRAF/NIRAS Safety Strategy for the Disposal of Category B&C Wastes. NIROND-TR report E, 2007 [55] Pedersen, K., Microbial Processes in Radioactive Waste Disposal. SKB Technical Report TR-00-04, 2000 [56] Poinssot, C., Ferry, C., Grambow, B., Kelm, M., Spahiu, K., Martinez, A., Johnson, L., Cera, E., de Pablo, J., Quinones, J., Wegen, D., Lemmens, K., McMenamin, T. Mechanisms Governing the Release of Radionuclides from Spent Nuclear Fuel in Geological Repository: Major Outcomes of the European Project SFS. Mat. Res. Soc. Symp. Proc. Vol. 932, [57] Pourbaix A., L'Hostis V., Passivation, Localised Corrosion and General Corrosion of Steel in Concrete and Bentonite. Theory and Experimentals. Journal de Physique IV 136, Proceedings of the International Workshop NUCPERF 2006 Corrosion and Long Term Performance of Concrete in NPP and Waste Facilities (Eds. L'Hostis V., Foct F. and Féron D.) March, 2006, Cadarache, France, pp , [58] Poyet, S., Conception du Superconteneur ONDRAF-NIRAS Phase 2: Simulation du Comportement Thermo-Hydrique du Tampon en Béton en Service. CEA Report RT DPC/SCCME A, [59] Poyet, S., Design of the ONDRAF Supercontainer Concept for Vitrified HLW Disposal in Belgium : Study of the Thermo-Hydric Behaviour of the Concrete Buffer. CEA Report RT DPC/SCCME/ A, [60] Put, M., Henrion, P., Modelling of radionuclide migration and heat transport from an HLW repository in Boom Clay. European Commission Report EUR 14156, Nuclear Science and Technology, Luxembourg, NIROND-TR E, April 2008

129 [61] Runchal, A.K., PORFLOW, a software tool for multiphase fluid flow, heat and mass transport in fractured porous media. User s Manual, Version ACRi, Bel Air, California, USA, [62] Sagoe-Crentsil, K.K., Glasser, F.P., Analysis of the Steel-Concrete Interface. In: Corrosion of Reinforcement in Concrete, Editors, C.L Page, K.W.J.Treadway and P.B.Bamforth, Society for Chemical Industry, Elsevier Applied Science, London and New York, ISBN , [63] Simpson J.P., Schenk R., Knecht B., Corrosion Rate of Unalloyed Steels and Cast Irons in Reducing Granitic Groundwaters and Chloride Solutions, Mat. Res. Soc. Symp. Proc. 50, pp , [64] Smart N.R., Blackwood D.J., Werme L., Anaerobic Corrosion of Carbon Steel and Cast Iron in Artificial Groundwaters: Part 2 Gas Generation, Corrosion 58(8), pp , [65] Smart, N.R., Blackwood, D.J., Marsh, G.P., Naish, C.C., O Brien, T.M., Rance, A.P., Thomas, M.I., The Anaerobic Corrosion of Carbon and Stainless Steels in Simulated Cementitious Repository Environments: A Summary Review of Nirex Research. AEA Technology Report AEAT/ERRA-0313, [66] Smart, N.R., Rance, A.P., Effect of Radiation on Anaerobic Corrosion of Iron. SKB Technical Report TR-05-05, [67] Taylor, K.J., Smart, N.R., Porter, F.M., The propagation of localised corrosion of carbon steel in cementitious environments. AEAT Report AEAT/ERRA-0309, Harwell, UK, [68] Tractebel Engineering. Category B & C Disposal Facility Design-Stage 2, performed in the context of the 2007 Cost Evaluation. Part 1 Closed Cycle No excess plutonium, [69] Van Cauteren, L., Evaluatie van de warmtegeeidbaarheid en de warmtecapaciteit van de Boomse klei. Nota Ref Niras, Brussel, [70] Van Keer, I., De Craen, M., Sedimentology and diagenetic evolution of the Boom Clay: State of the art. Long-Term Performance Studies of the Geological Disposal of Conditioned High-Level and Long-Lived Radioactive Waste. Report to NIRAS/ONDRAF contract CCHO-98/332 / KNT , R-3483, [71] Van Geet, M., Oxidation phenomena in Boom Clay : case study from the Northern Starting Chamber in the second shaft. In: Bastiaens, W. et al., The Connecting Gallery. EURIDICE Internal Report R-3483, p. 51, 2003 [72] Van Geet, M., Wang, L., De Boever, P. and De Craene, M., Geochemical boundary conditions relevant for assessing corrosion processes for the supercontainer design. SCK-CEN Report ER-14, NIROND-TR E, April

130 [73] Van Geet, M., Bernier, F., Sillen, X. And Li, X., (Excavation) Damaged and disturbed zone. ONDRAF/NIRAS Technical Note, [74] Van Genuchten, M. T., A closed form equation for predicting the hydraulic conductivity of unsaturated soils. Soils Science Society of America, 44, pp , [75] Volckaert, G., Neerdael, B., Manfroy, P., Lalieux, Ph., De Cannière, P. and Labiouse, V., Characteristics of Argillaceous Rocks: A Catalogue of the Characteristics of Argillaceous Rocks Studied with Respect to Radioactive Waste Disposal Issues: Belgium, Canada, France, Germany, Italy, Japan, Spain, Switzerland, United Kingdom, and United States. - Boom Clay. Revision number 2-05/5/1997. [76] Wang, L., Near-field chemistry of a HLW/SF repository in Boom Clay scoping calculations relevant to the supercontainer design. SCK-CEN Report KNT /CCHO /00/00, [77] Weetjens, E. Sillen, X., Thermal analysis of the Supercontainer concept. 2D axisymmetric heat transport calculations. SCK-CEN Report R-4277, [78] Weetjens, E., Sillen, X., Van Geet, M., Mass and energy balance calculations for the VHLW/Iron/(concrete)/Clay reference concept. NF-PRO Deliverable 5.1.2, [79] Weetjens, E., Van Geet, M. Phenomenological Description: Reference Concept (VHLW/Iron/(concrete)/Clay). NF-PRO Deliverable 5.1.1, [80] West, J.M., McKinley, I.G., Stroes-Gascoyne, S., Microbial effects on waste repository materials. In: M.J. Keith-Roach and F.R. Livens, Radioactivity in the Environment, Volume 2. Elsevier Science, pp , [81] Yeh, T.K., Macdonald, D.D., Motta, A.T.. Modeling Water Chemistry, Electrochemical Corrosion Potential, and Crack Growth Rate in the Boiling Water Reactor Heat Transport Circuits. 1: The DAMAGE-PREDICTOR Algorithm, Nuclear Science and Engineering, 121 Issue 3, pp , NIROND-TR E, April 2008

131 Appendix A Input Data and Boundary Conditions to Support Modelling Studies of the Belgian EBS Design for HLW Disposal A1 Introduction A1.1 Background Following the selection of the supercontainer with OPC buffer as the design for disposal of Category C wastes, subsequent ONDRAF/NIRAS research and design efforts are focused on elaborating and building confidence in this design. This Appendix is intended to provide a set of input data and focuses on the design for disposal of vitrified HLW (ZAGALC). A1.2 Structure of Appendix A The document is structured as follows: Dimensions of the supercontainer are provided in Section A2. The radiological inventory for vitrified HLW is provided in Section A3. The concrete buffer specification is provided in Section A4. The boundary conditions for thermal-hydrological (T-H) modelling are provided in Section A5. The specification for the carbon steel overpack is provided in Section A6. A reference composition for the Boom Clay and Boom Clay pore fluid is provided in Section A7. References are provided in Section A8. A2 Supercontainer Dimensions The dimensions of the principal components of the supercontainer are provided in Table A2.1. Diagrams illustrating the supercontainer and its position within the repository are provided in Figures. A2.1 and A2.2. NIROND-TR E, April

132 Table A2.1 - Dimensions of the supercontainer for ZAGALC vitrified HLW and its various components. All dimensions correspond to room or surface temperatures. Overpack and supercontainer dimensions are those reported by Belgatom ([3] Table 5) for a HLW supercontainer with a 30 mm-thick overpack Dimensions - canisters (metres) Canister length without lid Canister length with lid Nominal height of glass inside canister 1.1 Canister outer diameter Thickness of canister wall Dimensions - overpack (metres) Overpack length, L op Overpack thickness, T op Space between overpack and canisters (sides) Space between overpack and canisters (ends) Overpack outer diameter, D op Dimensions - filler (metres) Thickness 2nd phase filler T Dimensions - buffer (metres) Buffer length Buffer inner diameter Dimensions - supercontainer (metres) Length supercontainer Supercontainer outer diameter D sc Stainless steel outer liner thickness Dimensions - tunnels (metres) Tunnel length 1000 Spacing between 2 disposal tunnels 50 Inner diameter tunnel D g 3.0 Outer diameter tunnel D ex NIROND-TR E, April 2008

133 It is assumed that the canisters will be positioned symmetrically within the overpack using guides, but the guides may be ignored for Phase 2 modelling purposes. The canisters are assumed to touch each other. For modelling purposes, the space between canisters and overpack is metres along the canister sides, and metres at either end. This space is filled with glass frit. The composition of the glass frit is given in Table A3.2. filler Figure A2.1 - A diagram illustrating the near-field engineered barriers for repository disposal of vitrified HLW and illustrating the geometry of the supercontainer within the repository. The position of the supercontainer in relation to the tunnel liner is shown schematically. It is as yet undecided whether the supercontainer will be fabricated with a sealed envelope, with a perforated envelope, or with no envelope at all NIROND-TR E, April

134 6mm stainless steel envelope Figure A2.2 - A longitudinal cross section through the supercontainer. The position of the supercontainer in relation to the tunnel liner is shown schematically. It is as yet undecided whether the supercontainer will be fabricated with a sealed envelope, with a perforated envelope, or with no envelope at all 120 NIROND-TR E, April 2008

135 A3 HLW Inventory and Radiolysis Calculations The baseline ZAGALC vitrified HLW composition is provided in Table A3.1 and the composition of the glass frit in Table A3.2. Table A3.1 - Composition of ZAGALC vitrified HLW. Elemental composition [wt %] SiO B 2 O Al 2 O Na 2 O 9.8 CaO 4.0 Fe 2 O NiO 0.4 Cr 2 O P 2 O Li 2 O 2.0 ZnO 2.5 Oxides FP 10.6 Oxides actinides 0.9 MoO ZrO Fines (metallic particles) 0.7 TOTAL 100 Density of glass is 2750 kg m -3 For further details of the glass composition, ONDRAF/NIRAS refers to the standard AREVA vitrified product specification. NIROND-TR E, April

136 Table A3.2 - Composition of glass frit Elemental composition [wt %] SiO B 2 O Al 2 O Na 2 O 7.0 CaO 5.2 Li 2 O 2.6 ZnO 3.2 ZrO TOTAL 100 Density 1600 kg m -3 Detailed compositional data about fission product oxides are not available. The glass composition is computed by considering that actinide oxides are represented by UO 2, and that oxides of fission products are represented by (7.35% Cs 2 O % SrO). The baseline ZAGALC HLW radiological inventory at the time of vitrification is provided in Tables A3.3 and A NIROND-TR E, April 2008

137 Table A3.3 - A nominal radiological inventory of ZAGALC vitrified HLW fission products at the time of vitrification. Parent-daughter couples are shaded alternately in yellow and blue; calculated values, and values not considered, are in red. Bq per canister is calculated based on a glass volume of m 3 per canister Isotope [Bq m - ³] [Bq per canister] Se *E *E+10 Sr *E *E+15 Y *E *E+15 Zr *E *E+10 Tc *E *E+11 Ru *E *E+14 Rh *E *E+14 Pd *E *E+09 Ag-110m 4.44*E *E+12 Sn *E *E+10 Sb *E *E+10 Sb *E *E+14 Te-125m 2.17*E *E+13 I *E *E+07 Cs *E *E+14 Cs *E *E+10 Cs *E *E+15 Ba-137m 2.38*E *E+15 Ce *E *E+14 Pr *E *E+14 Pm-147 Not included Sm *E *E+13 Eu *E *E+14 Eu-155 Not included Probably overestimated as most I-129 expected to be lost during the vitrification process. Data for these radionuclides were not available. However, these are short-lived radionuclides and also low-energy beta emitters, and may be neglected when considering the total gamma emission. In Table A3.3, the activities of six parent-daughter couples were calculated as follows: Activity Pr-144 = activity Ce Activity Ba-137m = activity Cs Activity Rh-106 = activity Ru Activity Te-125m = activity Sb Activity Y-90 = activity Sr-90 1 Activity Sb-126 = activity Sn NIROND-TR E, April

138 Table A3.4 - A nominal radiological inventory of ZAGALC vitrified HLW actinides at time of vitrification. Decay chains are shaded alternately in yellow and blue; calculated values are in red. Bq per canister is calculated from Bq m -3 based on a glass volume of m 3 per canister Isotope [Bq m - ³] [Bq per canister] Cm *E *E+03 Am *E *E+11 Np *E *E+11 Pu *E *E+10 U *E *E+05 Th *E *E+06 Pa *E *E+05 Cm *E *E+09 Pu *E *E+07 U *E *E+06 Cm *E *E+10 Pu *E *E+11 U *E *E+05 Th *E *E+04 Cm *E *E+09 Pu *E *E+12 Am *E *E+13 Np *E *E+10 Pa *E *E+09 U *E *E+01 Th *E *E+03 Cm *E *E+04 Pu *E *E+00 Cm *E *E+13 Pu *E *E+10 U *E *E+06 Th *E *E+00 4N+ 3 4N+ 2 4N+ 1 Activities of these short-lived isotopes obtained by decay of long-lived parent isotopes assuming secular equilibrium. Information concerning the baseline spent fuel inventory will be provided in future versions of this report. 4N 124 NIROND TR E

139 A4 Concrete Buffer, Backfill and Wedge Blocks A4.1 Cement It is recommended that CEM I to European Standard BSENV 197-1:1992 be used with additional restriction that the cement has not been interground with materials other than gypsum (i.e., without slag) and has a low SO 3 content, preferably not exceeding 2% SO 3. A further limit is that the compound content of Ca 3 Al 2 O 6, as calculated from the Bogue formula, should preferably not exceed 5%. The organic content of interground gypsum should be low, and the cement should have been ground without the use of organic grinding aids. A4.2 Aggregate It is recommended that both fine and coarse aggregates should be limestone (calcium carbonate, calcite) containing not more than 2% each of magnesium, silicon and aluminium (as oxides). A4.3 Concrete Proportioning The concrete proportioning is based on a modification of the composition defined by Gallé et al. [10]. The proportions are approximately one part Portland Cement, two parts sand, four parts coarse aggregate (by weight). A4.4 Assumed Concrete Mix The assumed concrete mix is provided in Table A4.1. Table A4.1 - The composition of concrete buffer currently under consideration by ONDRAF/NIRAS. Two options are currently considered, a normal PC concrete, and a self-compacting concrete Component Content (kg m -3 ) Normal PC concrete Self-compacting concrete CEM I 42.5 N LA HSR LH Calcitec 2001 MS (ground CaCO 3 ) Limestone sand (0-4 mm) Calcareous aggregates (2-6 mm) Calcareous aggregates (6-14 mm) Calcareous aggregates (6-20 mm) Water : cement ratio Superplasticiser (Glenium ) 4.75% 14% NIROND-TR E, April

140 A4.5 Superplasticiser At present, it is more likely that a normal PC concrete will be chosen for the buffer on account of its lower superplasticiser content. The superplasticiser Glenium, a polycarboxylate etherbased material, will probably be used. A4.6 Nominal physical properties of concrete buffer Nominal physical properties of the supercontainer concrete are given in Table A4.2, based on data in Gallé et al. [10]. It is assumed that the proposed minor compositional modifications will have negligible impact on properties reported by Gallé et al. [10]. This is confirmed by unpublished data for a similar concrete made with calcareous aggregate, also shown in Table A4.2. We note that the aggregate density of the assumed concrete mix would be slightly increased, owing to the higher density of calcite (2710 kg m -3 ) compared with quartz (2650 kg m -3 ). However, the CEA formulation does not allow for entrained air, normally 0.6 to 1.0%. It is suggested that the two factors will approximately cancel. Heat capacity and thermal expansion are considered constant between 20 and 60 C although, in the TH model, the heat capacity is recalculated taking into account the dehydration of cement phases. Poisson s Ratio was calculated based on dynamic data (ultrasonic tests). The values at 20 and 60 C were found equal. The total water porosity was measured after the drying the samples at 60 C and the values at 20 C and 60 C are considered equivalent. The residual water permeability after a drying at 60 C was not measured. Note also that the concrete of Gallé et al. [10] was made with rounded fine and coarse aggregate, whereas the mineralogical restrictions placed on the preferred formulation (carbonate aggregate) mean that it would post probably have to be formulated with crushed (i.e. angular) fine and coarse aggregate. This may ultimately result in a higher water content being required than is specified in Table A NIROND TR E

141 Table A4.2 - Nominal physical properties of the supercontainer concrete (after Gallé et al. [10]) for siliceous and calcareous aggregates. The water:cement ratio is 0.43, identical to the assumed concrete mix in Table A4.1. Property Siliceous aggregate Calcareous 20 C 60 C 20 C Units Specific density (kg m -3 ) Thermal conductivity (W m -1 K -1 ) Specific heat capacity (J kg -1 K -1 ) Thermal expansion coefficient 1.10*E *E *E-05 (K -1 ) Elastic modulus (GPa) Compressive strength (MPa) Poisson's Ratio (-) Total porosity (%) Water permeability 4.00*E (m 2 ) Effective gas permeability 1.60*E *E *E-19 (m 2 ) Intrinsic gas permeability* 4.50*E *E *E-19 (m 2 ) After Gallé et al. [10]. Unpublished data of Gallé [10]. * Evaluated using the Klinkenberg approach. A4.7 Water content at 60 C The effect of heating on the concrete moisture content is not known. This arises for several reasons: (1) The stable moisture is a function of both temperature and relative humidity (2) Once drying commences, a gradient is established in the concrete and diffusion through the pore structure of a well made concrete may be extremely slow. (3) The bonding of water molecules in the cement solids is complex. In the most abundant solid, C-S-H gel, a range of sites exist differing in strength of bonding. Therefore water is not lost in discrete stages but is instead lost continuously, especially in the course of dynamic heating. However, methods have been developed for characterising the water content after drying (e.g. [4]). In order of increasing severity of drying, P drying and D drying have become standard. NIROND-TR E, April

142 These are achieved by drying over magnesium perchlorate at 23 C, and under vacuum at -79 C, respectively. These drying processes may have similar effects to what might be expected to happen after heating to 60 C and 90 C respectively. If this association is accepted, the effective water:cement ratio would decrease to the range 0.17 to 0.24 at 60 C (exact values depend on the cement analysis) and somewhat lower, by about 8%, for 90 C. Given that drying would produce a temperature gradient through the concrete, the slope of which is itself subject to calculation, it is recommended that the water:cement ratio 0.24 be used for calculation. This is likely to be a worst-case assumption. Numerical data for the calculations given above are taken form the review of Brouwers [4]. A4.8 Physical properties of filler It is assumed that the filler consists of portlandite, lime powder or cementitious grout. Nominal properties of portlandite or lime are provided in Tables A4.3 and A4.4. Heat capacities are from the National Institutes of Standards and Technology (NIST) Chemistry Web Book. Powder densities are based on information provided by Buxton Lime for as-supplied material. Table A4.3 - Nominal physical properties of portlandite filler (Ca(OH) 2 Density of portlandite 2,230 kg m -3 Density of portlandite powder 500 kg m -3 Porosity of emplaced filler 77.6% Specific heat capacity at 298 K 1180 J kg -1 K -1 Thermal conductivity 1 W m -1 K -1 Table A4.4 - Nominal physical properties of lime filler (CaO) Density of lime 3,345 kg m -3 Density of lime powder 1,000 kg m -3 Porosity of emplaced filler 70.0% Specific heat capacity at 298 K 1080 J kg -1 K -1 Thermal conductivity 1 W m -1 K -1 It is assumed that the thermal conductivity of both fillers is 1 W m -1 K NIROND TR E

143 A4.9 Backfill The exact composition of the cementitious backfill and its emplacement mechanism has yet to be determined. It is likely to be based on a PC with a carbonate aggregate, similar to the buffer concrete, although other aggregates, such as quartz sand, have been tested. A4.10 Tunnel Liner In order to provide support and stabilise the open tunnels during the operational phase of the repository, the tunnels are lined with concrete wedge blocks. The composition of the wedge block concrete is given in Table A4.5. Table A4.5 - The composition of the tunnel liner wedge block concrete currently under consideration by ONDRAF/NIRAS Component Content (kg m -3 ) PCCP CEM I 52.5N 430 PFA (Rugby Cement) mm graded crushed quartzite gravel 1055 Grade M washed quartzite grit sand 568 Content (litres m -3 ) Structuro *EMSAC 500 S 140 Water 30 (maximum) * EMSAC 500 S is an aqueous suspension of Elkem Microsilica manufactured by Elkem Materials. NIROND-TR E, April

144 A5 Boundary Conditions for T-H Modelling A5.1 Thermal conductivity of surrounding media The thermal conductivity of the surrounding media is given in Table A5.1. Table A5.1 - Properties of media surrounding the supercontainer. Thermal conductivity of backfill and Boom Clay after Cool [5]. Other backfill properties are assumed the same as for the concrete buffer. The data for the wedge blocks and Boom Clay are the same as for the Praclay Heater Test (ONDRAF/NIRAS, correspondence) Backfill (properties assumed same as concrete buffer except thermal conductivity) Thermal conductivity of backfill 1.0 W m -1 K -1 Dry density 2300 g cm -3 Specific heat capacity 1000 J kg -1 K -1 Wedge blocks Thermal conductivity of lining wedge blocks 1.5 W m -1 K -1 Dry density 2345 kg m -3 Specific heat capacity 770 J kg -3 K -1 Linear thermal expansion coefficient K -1 Boom Clay Thermal conductivity of Boom clay 1.69 W m -1 K -1 In Table A5.1, the backfill thermal conductivity (1 Wm -1 K -1 ) is the lowest acceptable value for this material, inasmuch as no specific material composition has yet been defined. The wedge block thermal conductivity (1.5 Wm -1 K -1 ) was the conservative lower limit used in scoping calculations for the Praclay Heater Test. The Boom Clay thermal conductivity (1.69 Wm -1 K -1 ) corresponds to undisturbed, saturated Boom Clay. A5.2 Thermal power of vitrifed HLW The variation of the power output of the waste with time is given by the red line in Figure 5.1 (after [60]). It can be written as: Q λit = Ai e (Q in W/tHM) i 130 NIROND TR E

145 with the coefficients A i and λ i defined in Table A Thermal output (W/tHM) Vitrified HLW (Put) Vitrified HLW (ORIGEN) NAGRA Vitrified HLW JNC-H12 Vitrified HLW 1 XSi/2003/ Time after waste production (a) Figure A5.1 - Power of vitrified HLW and its variation (red line) over time in years (a), after Put and Henrion [15]. The power is given in Watts per tonne of Heavy Metal equivalent (W thm -1 ) Table A5.2 - Coefficients for thermal power equation A 1 A 2 A 3 A 4 A λ 1 λ 2 λ 3 λ 4 λ E E E E E+00 NIROND-TR E, April

146 Each canister contains approximately 1.37 tonnes of heavy metal. In related work to predict the thermal evolution of the ONDRAF/NIRAS disposal concept, the vitrified waste content has been approximated as 1 thm to 0.75 of a canister. Therefore, in order to determine the power output per canister at any point in time after vitrification, it is necessary to divide the power values in Figure A5.1 by A6 Carbon Steel Overpack and Stainless Steel Envelope It is assumed that the overpack is constructed from P-235 carbon steel. Relevant properties are given in Tables A6.1 and A6.2. Data is taken from European Standards EN [8] and [9]. Table A6.1 - Chemical composition of P-235 steel (cast analysis) Steel grade Element % by mass Permitted deviation C 0.16 ± 0.02 Si 0.35 ± 0.05 Mn 0.60 to 1.20 ± 0.05 P max S max Al (total) 0.02 ± N Cr 0.30 ± 0.05 Cu 0.30 ± 0.05 Mo 0.08 ± 0.03 Nb ± 0.01 Ni Ti max 0.03 ± 0.01 V 0.02 ± 0.01 Others Cr + Cu + Mo + Nb NIROND TR E

147 Table A6.2 - Thermo-mechanical properties of P-235 steel of thickness 0.03 metres Property Density Value 7850 kg m -3 (Sect. 5 EN ) Specific heat capacity 4,800 J kg -1 K -1 Thermal conductivity 76 W m -1 K -1 Yield strength Tensile strength 225 MPa Up to 480 MPa It is assumed that the outer liner of the supercontainer is made of a low carbon stainless steel with enhanced Mo content (AISI 316L hmo). This type of steel has already been considered by the ONDRAF/NIRAS Corrosion Panel [14]. The 316L hmo stainless steel was recommended over normal 316L stainless steel, because an enhanced Mo content of 2.5 to 2.75 % Mo, compared with a normal Mo content of 2.0 to 2.5 %, significantly reduces the propagation of localised corrosion, such as pitting or corrosion beneath deposits. This is attributed to the precipitation of molybdenum oxide (probably MoO 2 ) in the acidic environment of pits. A 316L high-mo steel was used for extensive corrosion studies at SCK-CEN and its composition in wt % is given in Table A6.3. The standards that are the closest to this composition are Euronorm standards , (with addition of nitrogen) and All three standards specify Mo between 2.50 % and 3.00 % and accept S as high as %. Mechanical and thermal properties of 316L steel are given in Table A6.4 and are assumed to be equivalent to 316L hmo. NIROND-TR E, April

148 Table A6.3 - Chemical composition of 316L hmo steel Element % by mass C Si 0.61 Mn 1.16 P max S max Cr Mo 2.84 Ni Table A6.4 - Thermo-mechanical properties of 316L steel at room temperature. These properties are assumed to be equivalent to those of 316L hmo steel Property Value Density 8000 kg m -3 Specific heat capacity 500 J kg -1 K -1 Thermal conductivity 16.1 W m -1 K -1 at 100 C Hardness Yield strength Tensile strength Maximum 95 HRB (about 15.7 Rockwell C) 205 MPa 515 MPa Any coupling effects between radiolysis and corrosion mechanisms, such as production and consumption of oxygen, are neglected. 134 NIROND TR E

149 A7 Boom Clay Reference physical and mineralogical characteristics of the Boom Clay, and of Boom Clay pore fluids are provided in Tables A7.1, A7.2 and A7.3. These data are summarised in [6]. Table A7.1 -Petrophysical and hydraulic parameters of Boom Clay (compiled from [1] [2] [12] [18] [13] Property Value Units Bulk density (sat) kg m -3 Average grain density 2650 kg m -3 Water content % dry wt Total porosity (from migration experiments) vol % Specific surface area 44 m 2 g -1 In situ temperature 16 C Thermal conductivity 1.68 W m -1 K -1 Specific heat capacity 1400 J kg -1 K -1 Heat capacity 2.8 MJ m -2 K -1 Seismic velocity V p m s -1 Hydraulic conductivity Lab Field Vertical: m s -1 Horizontal: Vertical: m s -1 Horizontal: NIROND-TR E, April

150 Table A7.2 - Mineralogical composition of Boom Clay. Values in % total dry wt. [13] [17] [6] [7] Clay minerals Illite Smectite + illite/smectite ML Kaolinite Chlorite Chlorite/smectite ML 30-60% 10-45% 10-30% 5-20% 0-5% 0-5% Quartz 15-60% K-Feldspars Albite Carbonates Calcite Siderite Dolomite Ankerite 1-10% 1-10% 1-5% 1-5% present present present Pyrite 1-5% Organic Carbon 1-5% Others: Glauconite, apatite, rutile, anatase, Ilmenite, zircon, monazite, xenotime present present 136 NIROND TR E

151 Table A7.3 - A reference Boom Clay pore water composition Species or parameter Concentration Units Na K Ca Mg Fe Si 0.1 Al 2.40 E-05 mmole litre -1 HCO TIC (mg C litre -1 ) 15.1 Cl Total S 0.02 SO ph pco atm Eh -274 mv Temperature 16 C A more recent recommendation for a reference Boom Clay pore fluid and suggested reference compositions for two disturbed pore fluids are reproduced in Table A7.4 after Van Geet et al. [16]. It should be recognised that before reaching and interacting with the supercontainer, any Boom Clay fluids would first have to pass through and interact with the concrete of the wedge blocks and backfill. NIROND-TR E, April

152 Table A7.4 - Reference compositions for a worst-case disturbed Boom Clay pore fluid, expected disturbed pore fluid and a reference pore fluid [17] Case 1 Worst water ever measured Case 2 Expected perturbed water Case 3 Reference Boom Clay water mol/l mg/l mol/l mg/l mol/l mg/l SO , , S 2 O , HS - /S < Cl TIC Mg , ph Comments Maximum concentration ever observed/worst case [S 2- ] represents CERBERUS in-situ measurements. [SO 2-4 ] and [S 2 O 2-3 ] as observed in recent in-situ measurements Recent batch leaching experiments showed a Cl - concentration in the extract varying between 0.5 and 3 mg l -1 giving rise to pore water concentrations between 15 and 425 mg l -1, compared to 27 ppm measured by piezometer. At present, it is assumed that this high Cl - concentration is not freely mobile, but we cannot exclude a temporarily higher concentration, hence the value considered in Table A7.4. Note that the Boom clay is a marine clay deposited in an open marine environment. It is important to distinguish between mobile chloride (as measured in piezometers) and the total chloride content of Boom Clay (as measured in batch experiments). Most of the chloride content is not available for transport, probably due to the layer properties of compacted Boom Clay. 138 NIROND TR E

153 A8 References [1] Baeyens, B., Maes, A., Cremers, A., Henrion, P. N., Aging effects in Boom Clay, Radioactive Waste Management and the Nuclear Fuel Cycle, 6, , 1985a. [2] Baeyens, B., Maes, A., Cremers, A., Henrion, P. N., In situ physicochemical characterization of Boom Clay, Radioactive Waste Management and the Nuclear Fuel Cycle, 6, , 1985b. [3] Belgatom. Design for Overpacks for Vitrified Waste Canisters and Spent Fuel Assemblies. Belgatom Report /950 2, 2006 [4] Brouwers, H.J.H., The Work of Powers and Brownyard Revisited, Cement and Concrete Research, Vol. 34, No. 9, pages , [5] Cool, W., Thesis: "Quelques éléments de sûreté nucléaire opérationnelle pour un dépôt géologiquede déchets de haute activité et/ou longue durée de vie (catégorie B et C).", [6] De Craen, M., Wang, L., Van Geet, M., Moors, H. Geochemistry of Boom Clay pore water at the Mol site, SCK-CEN Scientific report BLG-990 (04/MDC/P-48), 2004a. [7] De Craen, M., Wang L., Weetjens E., Natural evidence on the long-term behaviour of trace elements and radionuclides in the Boom Clay, SCK CEN Final report to NIRAS/ONDRAF for the period , Contract nr. CCHO /00/00, KNT R-3926, Mol, Belgium, 2004b. [8] European Standard EN :2001 Specification for flat products made of steels for pressure purposes. Part 1: General requirements, [9] European Standard EN :2003, Specification for flat products made of steels for pressure purposes. Part 2: Non-alloy and alloy steels with specified elevated temperature properties, [10] Gallé, C., Sercombe, J., Pin, M., Arcier, G., Bouniol, P., Behaviour of high performance concrete under high temperature ( C) for surface long-term storage: thermo-hydro-mechanical residual properties. Material Research Society Symposium Proceedings, Vol. 663, [11] Gallé, C., Peycelon, H., Le Bescop, P.,. Effect of an accelerated chemical degradation on water permeability and pore structure of cement-based materials. Advances in Cement Research, Vol. 16, No. 3, pp , NIROND-TR E, April

154 [12] Henrion, P.N., Monsecour, M., Fonteyne, A., Put, M., De Regge, P., Migration of radionuclides in Boom Clay. In: Radioactive Waste Management and the Nuclear Fuel Cycle, Harwood Academic Publishers, New York, USA, Vol.6, pp , [13] ONDRAF/NIRAS, Safety Assessment and Feasibility Interim Report 2. (SAFIR 2). ONDRAF Report NIROND F, December 2001, [14] ONDRAF/NIRAS, A Review of Corrosion and Material Selection Issues Pertinent to Underground Disposal of Highly Active Nuclear Waste in Belgium. ONDRAF Report Ref: NIROND , [15] Put, M., P. Henrion, Modelling of radionuclide migration and heat transport from an HLW-repository in Boom Clay. EC, Nuclear Science and Technology, Luxembourg, EUR 14156, [16] Van Geet, M., Wang, L., De Boever, P., De Craene, M., Geochemical boundary conditions relevant for assessing corrosion processes for the supercontainer design. SCK-CEN Report ER-14, [17] Van Keer, I., De Craen, M., Sedimentology and diagenetic evolution of the Boom Clay: State of the art. Long-Term Performance Studies of the Geological Disposal of Conditioned High-Level and Long-Lived Radioactive Waste. Report to ONDRAF/NIRAS contract CCHO-98/332 / KNT , R-3483, [18] Volckaert, G., Neerdael, B., Manfroy, P., Lalieux, Ph., De Cannière, P., Labiouse, V., Characteristics of Argillaceous Rocks: A Catalogue of the Characteristics of Argillaceous Rocks Studied with Respect to Radioactive Waste Disposal Issues: Belgium, Canada, France, Germany, Italy, Japan, Spain, Switzerland, United Kingdom, and United States. - Boom Clay. Revision number 2-05/5/1997, NIROND TR E