ML19205A508

NER038 - Gocevski, V., Pathologies/Degradation Mechanisms Experienced by Hydro-Quebec During the Evaluation of Gentilly-2 NPP, Report Submitted to Ascet, (June 2015)

ML19205A508
Person / Time
Site:	Seabrook
Issue date:	07/24/2019
From:	Morgan, Morgan, Lewis & Bockius, LLP, NextEra Energy Seabrook
To:	Atomic Safety and Licensing Board Panel
SECY RAS
References
50-443-LA-2, ASLBP 17-953-02-LA-BD01, RAS 55113
Download: ML19205A508 (51)

Text

UNITED STATES OF AMERICA NUCLEAR REGULATORY COMMISSION ATOMIC SAFETY AND LICENSING BOARD In the Matter of Docket No. 50-443-LA-2 NEXTERA ENERGY SEABROOK, LLC ASLBP No. 17-953-02-LA-BD01 (Seabrook Station, Unit 1)

Hearing Exhibit Exhibit Number: NER038 Exhibit

Title:

Gocevski, V., Pathologies/Degradation Mechanisms Experienced by Hydro-Quebec During the Evaluation of Gentilly-2 NPP, Report Submitted to ASCET, (June 2015)

Gentilly-2 Nuclear Power Plant PATHOLOGIES/DEGRADATION MECHANISMS EXPERIENCED BY HYDRO-QUEBEC DURING THE EVALUATION OF GENTILLY-2 NPP Report Submitted to ASCET by:

Vladimir Gocevski Hydro-Quebec Second Draft Montreal, June 2015

FORWORD Guidelines for the objectives, the contents and the format of this report were provided by Mr. Neb Orbovic in a teleconference that was held on February 3, 2015 and was followed by an e-mail (by ASCET Organizing Committee summarizing the OECD/NEA/CSNI ASCET report preparation instructions. The following is an excerpt from this e-mail:

"I would like to summarize in few words our discussion. The participants reports should be submitted by April 15th and they should contain 9 Chapters: 1) introduction, 2) material aspects and lab. Material testing, 3) Material modeling, 4) Lab. structural components testing - destructive testing, 5) Structural modeling, 6) Non-destructive testing and in-situ condition assessment, 7) Structural acceptance criteria for structures with pathologies/degradation mechanisms, 8) structural repair and 9) Conclusions and recommendations for future work (conclusions and recommendations are related to each report)".

This report presents a brief summary of some particular concrete degradation mechanisms which have been encountered at some of the hydroelectric power plants located in Quebec in the last few decades as well as at Gentilly-2 Nuclear Power Plant (G-2 NPP). Many of the general pathologies/degradation mechanisms reported in the literature and which can be observed in various concrete structures, including NPPs, are not discussed here. The description of the degradation mechanisms presented in this report is brief and is intended only for sharing some aspects of our experience, gained mainly from work that was carried out for G-2 NPP, with those who will attend the meeting of June 29, 2015. Hopefully, time will be available to discuss some of these issues following the presentation at the meeting of June.

This incomplete second draft of the report is submitted in order to familiarize the conference participants with its contents and in order to provide the organizing committee with the conclusion summary so that it might be used as needed by the committee.

However, it is to be noted that several chapters of this draft report are incomplete or missing and many parts are still in raw format; i.e. they were not subjected to any proofreading, editing, or review. Nonetheless the draft is being submitted and distributed during the conference with the aspiration that it will be finalized in the near future.

Moreover, excerpts from another report are reproduced here for being pertinent. The report, which is entitled Impacts of ASR on Nuclear Power Plants; Problems and Research Needs, is being prepared by the author for RILEM as requested by Prof. Victor Saouma.

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TABLE OF CONTENTS FORWORD ............................................................................................................................... 1

1. INTRODUCTION ................................................................................................................. 3

2. MATERIAL ASPECTS AND LABORATORY MATERIAL TESTING ........................... 5

3. MATERIAL MODELING .................................................................................................. 13

4. LAB. STRUCTURAL COMPONENTS TESTING - DESTRUCTIVE TESTING .......... 14

5. STRUCTURAL MODELING............................................................................................. 14

6. NON-DESTRUCTIVE TESTING AND IN-SITU CONDITION ASSESSMENT............ 17 6.1 Introduction ............................................................................................................ 17 6.2 Non-Destructive Testing; Periodic Integrated Leakage-Rate Testing ............................... 17 6.3 In-Situ Condition Assessment ........................................................................................... 20 6.3.1 The consequences of AAR in the concrete structures of a NPP- introduction ............... 20 6.3.2 The change of post-tensioning due to combined effects of AAR and creep .................. 21 6.3.3 Micro cracking propagation throughout the prestressed concrete sections due to creep and/or combined effects of AAR and creep ...................................................... 25 6.3.4 Behaviour at the junction of the perimeter wall and the base slab caused by different expansion of the wall and the slab (due to different confinement) ............................. 26 6.3.5 Concrete splitting in the plane where the ducts with post-tensioning cables are located (due to creep and/or combined effects of AAR and creep) ........................................ 28 6.3.6 Effects of AAR on non-prestressed reinforced concrete in contact with water at G-2 NPP ........................................................................................................ 30 6.3.7 Effects of AAR on the turbo-generator foundation at G-2 NPP ................................... 32 6.3.8 Structural integrity under seismic excitation of AAR-affected containment ............... 34 6.3.9 Freeze-Thaw cycles - Concrete cracking and their propagation, delamination and concrete spalling .................................................................................................. 36 6.3.10 Initial cracking - construction of the containment structure ....................................... 37

7. STRUCTURAL ACCEPTANCE CRITERIA FOR STRUCTURES WITH PATHOLOGIES/DEGRADATION MECHANISMS.................................... 39

8. STRUCTURAL REPAIR .................................................................................................... 39 8.1 Introduction ............................................................................................................39 8.2 GENTILLY SPENT FUEL POOL LINING PROJECT ................................................... 40

9. CONCLUSIONS AND RECOMMENDATIONS FOR FUTURE WORK ....................... 41 BIBLIOGRAPHY ................................................................................................................... 45 2

1. INTRODUCTION The integrity of the various structures of nuclear power plants and nuclear facilities is essential to ensure the continuing operation of these installations without compromising public safety or environmental protection. As time pass by, the material properties of these structures continue to change due to environmental influences, various chemical processes and other factors.

The life cycle of existing nuclear power plants was often chosen to be 30-40 years at the time of the original design. The term for the initial 40-year-license, usually issued for nuclear reactor operators, is based mainly on economic considerations and not due to limitations imposed by the nuclear technology. It is the term of the initial license which has dictated the strategy adopted in the design of most structures and components of NPPs for an expected 30-40-year life cycle. In order to ensure the safe operation of nuclear power plants in subsequent life cycles, it is essential that the effects of age-related degradation on NPPs structures, systems and components, be assessed and managed during both the current and the subsequent life cycles.

The Nuclear Power Plant of G-2 owned by Hydro-Quebec, was approaching the end of its first life cycle of 30 years, in accordance with the initial design. Before the decommissioning of the plant, G-2 was poised for a major refurbishment project with the aim of extending its service life for another cycle of 25-30 years. The G-2 containment structure was constructed using concrete and mild carbon reinforcing steel. It was also pre-stressed using three different types of tendon anchorage arrangements. The ring beam accommodates tendon anchorage for the ellipsoidal dome and the vertical containment walls. Four vertical buttresses accommodate horizontal tendon anchorage for the walls (Figure 1).

Figure 1: Containment structure (reactor building envelope), vertical and horizontal sections 3

To evaluate the possibility of extending the service life of G-2, Hydro-Quebec (HQ) has had to prepare and execute an elaborate evaluation program, which included visual inspection of the accessible areas, in-situ measurements, laboratory material testing and an elaborate numerical analysis. Moreover, during the period from 1993 to 2003, HQ undertook several extensive research programs in collaboration with McGill University (McGill), McMaster University (McMaster), the Universality of Sherbrook (Sherbrook) as well as the Natural Sciences and Engineering Research Council of Canada (NSERC).

The contents of the joint research programs of HQ-Sherbrooke-McMaster-NSERC and HQ-McGill-NSERC were proposed by HQ and were approved by the NSERC. These University-Industry research programs were intended to define, by laboratory and in-situ testing, the properties of plain, reinforced and pre-stressed concrete affected by Alkali-Aggregate Reaction (AAR) and the behaviour of the affected concrete under the severe Canadian climate conditions as well as to develop an adequate analytical procedure for assessing the behaviour of AAR-affected concrete structures.

With the restrictions imposed on sample extraction from an operating NPP, conducting an elaborate numerical analysis was the more practical evaluation approach. During the years 2000 and 2003, Hydro-Quebec carried out an extensive numerical study of the G2 reactor building envelope, and later on, HQ conducted extensive analyses for all the structures of G2 during the five-year period from 2006 to 2010. These studies included evaluation of various aging/degradation mechanisms. The numerical simulations included a transient thermal analysis over the period of construction, a series of static analyses together with simulation of a continuing AAR effect, and a dynamic analysis simulating a seismic event.

By assessing the observed damage of concrete structures as well as the damage predicted by analyses, HQ was able to define a number of degradation mechanisms relevant to G-2 NPP. In general, the degradation mechanisms affecting concrete structures have been extensively studied and are relatively well known. Some of these mechanisms are listed below in groups according to the severity of the damage that they may produce in the affected structures:

1. Mechanisms referred to as Primary Degradation Mechanisms which are expected to cause significant damage:

• Alkali-Aggregate Reaction

• loss of pre-stressing (stress relaxation)

• corrosion of reinforcing and prestressing steel

• chemical attack

2. Mechanisms referred to as Secondary Degradation Mechanisms which are not expected to cause significant damage:

• construction deficiencies (initial hydration cracking)

• differential settlement

• shrinkage

• creep

• normal degradation over time

• frequent leakage-rate testing, accelerating the degradation over time

• freeze/thaw cycles (important for Canadian climate) 4

• thermal exposure

• radiation While many of the known degradation mechanisms - presented in the technical and scientific literature - are not pertinent to G-2 NPP; however, the broad research work conducted for G-2 NPP, for which an extensive technical documentation is available, was of help in identifying the mechanisms most relevant to the power plant. The structural damage due to these mechanisms may vary significantly ranging from minor to major damage. It should be noted that AAR may not be observed in many nuclear installations and is not an industry-wide phenomenon. However, most of the degradation mechanisms discussed in this report are related to concrete affected by the AAR; in other words, they are in combination with AAR. Some of these mechanisms are listed in what follows:

• Construction sequence: constraints related to thermal hydration, high cement content, resulting in initial cracking, mainly, to the exposed concrete. Although this is not an aging degradation mechanism because it happens during the construction period; however, it is an important factor that should be considered because it greatly influences the formation and the propagation of cracks that result from the other aging mechanisms;

• Shrinkage: resulting in concrete cracking;

• Creep: resulting in concrete cracking and loss of post-tensioning (affecting prestressed concrete of the reactor building); uneven creep throughout the wall section of the reactor building (in the presence of high in-plane prestressing):

resulting in the creation of micro cracking and affecting building's air tightness and concrete splitting (similar to the one at Cristal River NPP);

• Freeze-Thaw cycles (important for Canadian climate): resulting in concrete cracking and the propagation of cracks, delamination and spalling of concrete;

• Periodic integrated leakage-rate testing (high pressure air tightness safety testing at 124 kPa(g) every three years: resulting in the widening and the propagation of existing cracks during and after each performed test; the test frequency is imposed by the Nuclear Safety Authority and may not be considered as an actual aging mechanism. However, it is important to be mentioned because it influences the degradation of the concrete of the reactor building and reduces its air tightness;

• Overall and differential settlement of structures: resulting in concrete cracking and structural deformation (causing misalignment of turbo generator);

• Alkali-Aggregate Reaction: resulting in concrete cracking, concrete splitting (especially in pre-stressed concrete elements); structural deformation due to concrete swelling, and degradation of mechanical characteristics of concrete.

2. MATERIAL ASPECTS AND LABORATORY MATERIAL TESTING Hydro-Québec owns and operates more than 656 dams and 60 hydroelectric power plants. As of today, there are 36 hydraulic structures in these installations in which the AAR has been confirmed with only 7 are considered severely affected, 15 moderately affected, while a mild effect was found in the remaining 14 structures. G-2, the only 5

Nuclear Power Plant owned by Hydro-Quebec, is affected by AAR in concrete (severely affected if based on the free expansion rate of concrete but moderately if based on the in-situ measured expansions). Therefore, the AAR is the major concrete degradation mechanism at G-2.

In terms of concrete compressive strength, there are three different classes which were used in the design and the construction of the Canadian "CANDU 6" Nuclear power Plants (five NPPs). It can be assumed that similar classes were used in the construction of other NPPs around the world. The data below are extracted from the concrete test reports of G-2:

• Concrete compressive strength 20 to 25 MPa; ~ 260 kg/m3 cement; min 250kg/m3

• Concrete compressive strength 25 to 30 MPa; ~ 340 kg/m3 cement

• Concrete compressive strength 30 to 35 MPa; ~ 430 kg/m3 cement; max 450kg/m3 The aggregate used in the preparation of concrete mixes was limestone. In 1985, it was found that the aggregate in the region of Trois-Rivires - used in the construction of G is reactive. The average alkali content in the cement was estimated to be 0.820 %;

therefore, the alkali content in the concrete varied between 2.05 kg/m3 and 3.69 kg/m3.

Table 1 shows few examples for the classification of the degree of reactivity for concrete structures in the province of Quebec. Based on test results, it can be observed that the concrete used for the construction of G-2 NPP is more reactive when compared to the concrete used in the construction of any of the other AAR-affected structures in HQs hydroelectric power plants. For example, the maximum free expansion rate of concrete at Beauharnois hydroelectric power plant, as obtained from the in-situ measurements, is 95 m/m/y, at 100% Relative Humidity (RH). On the other hand, the maximum free expansion rate of concrete at G-2 NPP is estimated to be 150 m/m/y at 100% RH (may go up to 250 m/m/y when obtained from extracted concrete cores). It is known that lower concrete expansion rates are observed at lower relative humidity values. It is also known that concrete expansion is affected or slowed down, due to internal or external confinement. Most NPPs structures are heavily reinforced and/or post-tensioned and the relative humidity is usually low. In fact, the internal confinement and the lower relative humidity are the two main factors which may reduce concrete expansion rate considerably. As shown in Table 1, the free expansion rate of concrete at G-2 (150 m/m/y), is reduced to an average of 9 m/m/y in the walls of the containment building, under by-axial state of stress (in-situ internal confinement).

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Table 1 Classification of the degree of reactivity for concrete structures in Quebec From the experimental work that was conducted over the last forty-year period (in-situ and laboratory tests), particularly the last 25 years, HQ has acquired knowledge and gathered important information related to the following:

• Mechanical properties of concrete at G-2 NPP and hydroelectric power plants

• The expansion rate of concrete (small concrete samples - accelerated tests)

• Effect of RH and temperature on concrete expansion (small cylindrical samples)

• The results of uniaxial and multi-axial tests of AAR-affected concrete

• Effect of confinement on small cylindrical samples

• Effect of 1D confinement on concrete expansion (small cylindrical samples)

• Effect of 2D confinement on concrete expansion on 350 x 350 x 350 mm cubes

• Effect of 2D confinement on concrete expansion on 600 x 600 x 600 mm cubes

• The expansion rate of concrete on real structures as obtained from in-situ measurements. The effect of uniaxial and biaxial confinement on full scale structures.

At the beginning and following the time when the AAR was confirmed as the main problem in many concrete structures, initial research work was conducted and the objective was to understand the phenomenon: the chemical reaction and the consequent concrete swelling.

In fact, considerable research was carried out at that time, by researchers with expertise in material science, on concrete as a material affected by AAR. Today, the reactivity of existing concrete as well as the main factors contributing to an intensive Alkali-Silica Reaction are well defined. Standard testing procedures and methodologies to detect the reaction and to judge its intensity - based on the swelling of small samples - have been formalized in numerous national standards. It should be recognized that the main intention behind conducting the developed standardized tests was and still is to only 7

provide an answer to the question about whether the alkali-aggregate reaction is present or not in the examined concrete.

With the research that we have initially conducted on small samples, attempts were made to obtain other information such as: heterogeneity of the volumetric expansion, development of stresses in concrete due to the swelling of the silica gel, effect of confinement, etc. The objective of these attempts was to provide means for judging the behaviour of AAR-affected concrete in real structures; however, these attempts at HQ and, to the best of our knowledge, similar attempts elsewhere, have - so far - failed. The main reasons are: (a) the RH in each tested sample was 100 % with a uniform distribution of the RH in the entire sample (which is not the case for concrete in real structures), and (b) the samples were too small to provide adequate answers to the sought-after information listed above. In the authors opinion this is the most valuable lesson that HQ has extracted from many years of research (in particular testing small samples) on AAR-affected concrete. In fact, more meaningful conclusions, related to the behaviour of AAR-affected concrete in real structures, were obtained by bi-axial testing of 600 mm x 600 mm x 600 mm concrete blocks (the setup did not permit testing of larger cubic blocks) having non uniform distribution of RH (as is the case of real structures), and by conducting in-situ material testing and taking measurements of the deformed structures.

To test large concrete blocks in by-axial stress state; HQ, in collaboration with the University of Sherbrooke, has constructed an adequate press machine (Figure 2). The tested blocks were instrumented with 22 thermocouples and 9 strain extensometers. The locations where extensometers were located in the tested cubic blocks are shown in Figure 3. Two sides of each block were in constant contact with water reservoirs providing non-uniform distribution of temperature and humidity in the blocks (Figure 3).

The variation of the temperature and the humidity within the block was simulated by numerical analysis. The values were also obtained from the measurements taken by the installed instruments (Figure 4). The loading phases are presented in Figure 5. The measured deformations during a period of 1100 days (3 years) are presented in Figure 6.

Experimental Set-Up Experimental Set-Up Flatjack Load Cell Load cell Steel plate (115mm)

Steel rod (=75mm)

Data Acquisition Flatjack (Freyssinet)

System Concrete buffer Steel plate (120mm) 3.2 m Data Acq. System Water tank Horizontal Loading (CR10)

Concrete buffer Pressure accumulator Concrete slab Figure 2: Press machine for testing of large concrete blocks in by-axial stress state 8

Extensometer Location Extensometer Location Horizontal section Vertical section T4 G3 G2 G1 G4 T3 T1 G4 G5 T2 Figure 3: Locations of strain extensometers in a 600 mm cubic block Temperature Relative Humidity distribution (oC) distribution (%)

Figure 4: Temperature and Relative Humidity distribution in a 600 mm cubic block Loading Phases Days 0 100 200 300 400 500 600 700 800 900 1000 6.0 MPa 3.3 MPa Vertical Loading 0.1 MPa 1.4 MPa Horizontal Loading 0.0 MPa 0 100 200 300 400 500 600 700 800 900 1000 Figure 5: Loading Phases 9

Deformation of vertical extensometers

= 0.3 /day

= 0.3 /day

= 1.25 /day

= 0.875 /day (a) Vertical Deformation Deformation of horizontal extensometers

= 0.22 /day

= 0.45 /day = 0.4 /day

= 0.5 /day

= 1.33 /day

= 0.875 /day (b) Horizontal Deformation Figure 6: Deformations measured by extensometers The results presented in Figure 6 indicate that the volumetric expansion of the concrete block reduces even in case of a uniaxial confinement only (vertical load generating the stress of 3.3 MPa). As shown in Figure 6 (a), the extensometers T2 and G5, which were placed in the region of the concrete block having RH = 90% or more and had an initial expansion rate of = 1.25 /day in the direction of the applied load (3.3 MPa), slowed down, after the application of the vertical load (almost no expansion for about 200 days),

and then restarted again with an expansion rate of = 0.30 /day, which is 4 time slower. As for the extensometers G1 and T4 placed in the region of the concrete block having RH = 85% to 88% and RH = 82% to 85% respectively, they are not showing any expansion of concrete even after about 550 days following the application of the load.

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However, the expansion rate in the perpendicular direction has stayed almost the same (T3, G3) following the application of the vertical load (3.3 MPa), while the Poissons effect for the period of the load application (550 days) is noticeable as shown in Figure 6 (b).

The difference of the measured expansion rates in the water tanks direction (horizontal),

obtained from the extensometers T3 and G3, which are placed in the region of the concrete block having RH = 94% to 97%, and the measured expansion rates in the perpendicular horizontal direction obtained from the extensometers T1 and G4, placed in the region of the concrete block having RH = 85% to 88% and RH = 79% to 82%;

respectively, are due to the locations at which they were installed in the concrete block.

It is to be noted that following the application of the horizontal load (generating the 1.4 MPa stress) in the same direction of the extensometers T1 and G4, the expansion in the same direction has almost stopped as shown in Figure 6 (b). However, the expansion in the perpendicular unconfined direction (water tanks direction) has continued at a very low rate of = 0.22 /day. The increase of the vertical load (generating stress of 6.0 MPa) combined with lateral confinement (generating the 1.4 MPa stress) has caused the expansion to almost stop in all three directions.

The same concrete behaviour was observed from the in-situ measurements in the confined (reinforced and prestressed) concrete elements of the NPP structures. During the construction (concrete casting) of the containment building, 139 extensometers were placed in the concrete. They were located on a vertical line making 203 degrees with the north direction and were placed in planes parallel to the post-tensioning cables near the external and the internal sides of the base slab, the walls, and the dome. The containment building was extensively analysed using elaborate numerical analysis software which was developed by HQ and which was used successfully in earlier studies to analyze AAR-affected structures of hydroelectric power plants. The results obtained by the numerical analysis were comparable to the results obtained from the instruments as shown in Figures 7 and 8.

Sensor XE-23 (XT-03-B) Sensor XE-24 (XT-04-B)

Vertical - internal Tangential - internal Level 4,115 m Level 4,115 m 40 120 30 100 20 80 Strain (um/m) 10 60 0 40 Strain (um/m)

-10 20

-20 0

-30 -20

-40 -40

-50 -60

-60 -80

-70 -100

-80 -120 85-01-01 87-01-01 88-12-31 90-12-31 92-12-30 94-12-30 96-12-29 98-12-29 00-12-28 85-01-01 87-01-01 88-12-31 90-12-31 92-12-30 94-12-30 96-12-29 98-12-29 00-12-28 Date Date Measurement FE estimate Linéaire (Measurement) Measurement FE estimate Linéaire (Measurement) m/m/day Rate= 0,0163 m/m/day Rate= 0,0093 Figure 7: Estimates of vertical and horizontal strain rates obtained by analyses vs.

measured values at the interior perimeter of the containment building wall (Elevation 4.115 m; 203° with the North) 11

Sensor XE-58 (XT-18-C)

Vertical - external Sensor XE-57 (XT-17-C)

Level 38,405 m Tangential - external 140 Level 38,405 m 120 0 100 -20 80 Strain (um/m)

-40 60 -60 Strain (um/m) 40 -80 20 -100 0 -120

-20 -140

-40 -160

-60 -180

-80 -200

-100 -220

-120 -240

-140 -260 85-01-01 87-01-01 88-12-31 90-12-31 92-12-30 94-12-30 96-12-29 98-12-29 00-12-28 85-01-01 87-01-01 88-12-31 90-12-31 92-12-30 94-12-30 96-12-29 98-12-29 00-12-28 Date Measurements FE estimate Linéaire (Measurements) Measurements FE estimate Linéaire (Measurement) Date Rate = 0,0187m/m/day Rate = 0,0143m/m/day Figure 8: Estimates of vertical and horizontal strain rates obtained by analyses vs.

measured values at the exterior perimeter of the containment building wall (Elevation 38.405 m; 203° with the North)

The minimum free expansion for the concrete at G-2 NPP was estimated to be 140m/m/y at RH of 100%. The average initial compressive stress in the concrete caused by post tensioning cables in the reactor building of G-2 NPP is presented in Table 2 and the effects of the internal confinement and the lower values of RH on concrete expansion rates, as observed at different locations in the plant, are presented in Table 3. As evident, pre-stressing of heavily reinforced concrete sections reduces the expansion of the confined building to a large extent. As shown in Table 3, the horizontal expansion rate varies from a minimum value of about = 0.6 m/m/y to a maximum value of about =

8.3 m/m/y, while the vertical expansion rate varies from about = 0.6 m/m/y to about

= 15.5 m/m/y. It is important to point out that an equivalent reduction of the expansion rate, or of the expansion, has been observed in the unrestrained direction - which is perpendicular to the wall - similar to that observed in both of the other prestressed (confined) directions. The same was concluded from the test results obtained using the 600 mm cubic blocks, as described earlier.

Table 2: Average Initial Compressive Stress in the concrete Caused by Post Tensioning Cables in the Reactor Building of Gentilly-2 Nuclear Power Plant Element of the Average Initial Direction Structure Compressive Stress (MPa)

Foundation mat Circular 2.297 Vertical 3.459 Wall Circular 5.938 Top Annular Beam Circular 3.349 Dome Spherical 10.713 12

Table 3: Strain rates at different elevations of the containment building wall as recorded by the extensometers (203° with the North).

Measured Measured Measured Measured Elevation (exterior) (interior) (exterior) (interior)

(m) Horizontal Horizontal Vertical Vertical (m/m/year) (m/m/year) (m/m/year) (m/m/year) 1.9 -0,4745 0,5840 0,5840 -0,6570 4.1 2,8835 5,9495 1,2775 3,3945 6.8 1,4600 2,7375 4,0515 2,8835 22.4 0,5840 0,4745 1,2410 5,5845 38.4 5,2195 6,5335 6,8255 4,5625 40.8 4,6720 8,2490 15,4760 5,1830 Based on the aforementioned test results for the large blocks as well as the behaviour of the confined containment building - obtained from instrumental data and other measurements - and from our experience, gained through the evaluation of numerous hydroelectric concrete structures affected by AAR, the following can be concluded:

• Humidity distribution plays an important role in determining the long-term, expansion rate and the accumulated total expansion of AAR-affected concrete structures.

• The influence of 1D, 2D or 3D confinement combined with the influence of humidity distribution have to be evaluated based on in-situ measurements of the real structure or based on laboratory tests conducted on concrete samples with sufficiently large dimensions.

• The results obtained from testing of small cylinders (D = 130 mm or 150 mm; H =

240 mm or 300 mm) having uniformly distributed 100% relative humidity in the entire sample, or from even larger samples, do not adequately represent the behaviour of AAR-affected concrete structures.

3. MATERIAL MODELING Hydro-Québec has instigated and participated in a number of laboratory testing and some material modeling work, which were carried out at the University of Sherbrook and

École Polytechnique. However, the numerical procedure that was developed and used by Hydro-Quebec is not intended for the simulation of material tests (concrete cylinders or blocks) and using it for this purpose is not adequate. In fact, within this framework, which represents a phenomenological approach, the tests on the cylinders and blocks are perceived as material tests that define the rate of free expansion under different conditions. These tests provide a valuable information that may be employed to identify the parameters entering the law of reaction kinetics; they cannot, however, be simulated as initial boundary-value problems. The latter requires a multi-scale approach (i.e.,

micro/meso-scale) that, even though conceptually attractive, cannot be employed in the context of analysis of large-scale structures.

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It is to be noted that Hydro-Quebec has attempted to use a mesoscale approaches that distinguishes between the concrete matrix and pores occupied by gel and water; however, even though, the analyses had some success on simulating the behavior of small structural elements, the attempts were not successful in the context of analysis of large-scale structures. Moreover, the type of testing that is required to define the mechanical properties of the concrete matrix - without voids - is not clear.

4. LAB. STRUCTURAL COMPONENTS TESTING - DESTRUCTIVE TESTING Hydro-Québec has sponsored and guided several testing programs that were mainly intended for finding answers to some questions related to AAR-affected concrete structures located in some of its hydroelectric power plants. In one of these programs, reinforced concrete beams - affected by AAR - were tested at the University of Sherbrook and in another program, real-size portions of the gate slots - affected by AAR - of Beauharnois and La Gabelle power houses were tested at McGill University.

5. STRUCTURAL MODELING 5.1 Introduction An adequate Aging Management Program (AMP) is required in order to assess the existing stress/deformation state in a nuclear power plant structure. The AMP is also required in order to identify the aging mechanisms which may impair the proper functioning of the plant. For these reasons, an AMP was carried out for the G-2 containment building. The primary objective of the Program was to ensure that the requirements of CSA Standard N-278 are satisfied and should continue to be satisfied as long as the building is in operation. The first step that was carried out in this AMP was the proper identification of all possible aging mechanisms.

Over the last fourteen years Hydro-Quebec has carried out extensive numerical studies for the containment structure of G-2 NPP and for the other concrete structures affected by AAR in the plant. These numerical simulations included the implementation of a transient thermal analysis for the period of construction and for the freeze-thaw cycles occurring during a fifty-year period. The numerical simulations included as well static analyses for simulating the continuing Alkali-Aggregate Reaction and dynamic analyses simulating seismic events. An in-depth study of the consequences of an assumed crash of a fast flying commercial aircraft upon the containment building was also carried out for assessing, in particular, the potential damage to the AAR-affected concrete structure. The study, which was prompted by the September 11, 2001 terrorist attacks in the USA, has also incorporated the assessment of the damage due to the subsequent fire that to follow the crash (due to about 92,000 litres spilled aircraft fuel).

Generally speaking, the currently available commercial finite element codes are not prepared to adequately address some complex problems involving AAR-related swelling.

In particular, most of these codes lack material models with constitutive relations that are 14

suitable for the description and the evolution of complex material properties related to AAR.

Based on Hydro-Quebec's experience in simulating the behaviour of Hydroelectric and NPP structures affected by AAR swelling, the essential requirements of the concrete/reinforced concrete constitutive model accounting for the chemo-mechanical interaction, which should be incorporated in advanced Finite Element (FE) codes, are as follow:

• Adequate description of the kinetics of the reaction

• General failure criterion, provision for the development of irreversible deformations, general criterion for the onset of macro-cracking in both compression and tension regimes

• Degradation law for strength and deformation characteristics

• Proper description of propagation of damage in both tension and compression regimes (viz. homogenization incorporating a characteristic dimension, XFEM or similar)

• Constitutive relation for the interface material relating the velocity discontinuity to the traction vector.

Hydro-Quebec has successfully developed constitutive relations for AAR-affected reinforced concrete that satisfy the aforementioned requirements and has incorporated them in the main algorithms of the commercially available finite element codes:

COSMOS/M and ABAQUS.

Pursuing an appropriate analytical procedure, that includes calibration steps, is of great importance in any nonlinear static or dynamic analysis as it is a basic requirement for obtaining reliable and accurate results. This concept was observed in the methodology applied by HQ for analyzing the structures of G-2 NPP. The procedure is outlined in the following:

Step 1:

• Evaluate material properties based on laboratory reports:

o results from tests on cylinders at 28 days o results from periodic tests conducted on extracted cores (from the time of construction until the present time)

• Evaluate the rate of free expansion o based on a non-confined direction (vertical) and the distribution of relative humidity o based on periodic reports: surveying results of the benchmarks installed in the structure and the instrumentation data (winter and summer readings, measurements for a period of 8 to 10 years minimum)

• Define the parameters used in the numerical formulation o The material properties required to define the model parameters are as follows (for COSMOS/M software):

Modulus of elasticity of concrete Poissons ratio 15

Density of concrete Initial compressive strength of concrete Strength parameter c10 Strength parameter c20 Strength parameter c30 Confinement stress at which the AAR expansion rate reduces Confinement stress at which the AAR expansion rate is zero Strain softening parameter Material constant that controls the rate of strain softening Material constant for brittle-ductile transition value of confinement stress Material constant that defines the form of residual strength envelope AAR initiation step number End of AAR step number Time increment corresponding to real time for AAR (EX: 1 time step = 1 month)

Maximum free expansion rate due to AAR Constant (Bo) in the exponential law for free expansion (this parameter is changed depending on total time length)

Degradation constant for c10 Degradation constant for c20 Degradation constant for c30 Degradation constant for Ex Correction parameter to allow the degradation of strength properties during the length of AAR Number of elements for detailed printout Frequency of writing output file

• Define the time step to be used Step 2:

• Run Preliminary Analyses: runs for 20, 30, 40... years, calibration of the input properties o Comparison of calculated and measured displacements at 20, 30, 40 years.

o Comparison of damaged regions (predicted versus observed), if possible o If the compared results are satisfactory, go to step 3, if not, make corrections in step1 and restart the analysis from the beginning (i.e., from the period right after construction) o Repeat the abovementioned steps until acceptable results are obtained Step 3:

• Run Final Structural Analyses for Present Time Evaluation: from the end of construction time until present time o Comparison of calculated and measured displacements at present.

16

o Comparison of damaged regions (predicted versus observed) at present o Seismic evaluation (nonlinear dynamic analysis of the cracked structure) based on the calculated Structural Safety Indicator Step 4:

• Run Final Structural Analyses for future time assessment: evaluation of the structural behaviour in the future; service life expectations or service life extension studies of the cracked infrastructure o The analysis to assume a continuous expansion of concrete for a period of 40 years starting from today. At the end of each period, seismic evaluation to be performed

6. NON-DESTRUCTIVE TESTING AND IN-SITU CONDITION ASSESSMENT 6.1 Introduction The effects resulting from some of the above-listed degradation mechanisms have been observed in many structures of G-2 NPP, including:

The containment building and the enclosed substructures The spent fuel pool and the transfer canal The turbo-generator foundation A brief discussion, of the in-situ leakage-rate testing and the in-situ condition assessment of the above-mentioned structures, is presented in the following sections.

6.2 Non-Destructive testing, Periodic Integrated Leakage-rate Testing For G-2 NPP, periodic integrated leakage-rate testing was performed every three years at 124 kPa(g). Generally speaking, these high pressure air tightness safety testing cause the widening and the propagation of existing cracks, during and after each performed test.

The testing frequency and the pressure magnitude are imposed by the Nuclear Safety Authority and these requirements are not the same in all countries. Therefore, the deterioration effects produced by these tests, in terms of concrete "aging" or cracking are different from one country to another. In fact, the tested building may not be affected due to confinement and, in this case, the test may not be considered as an aging mechanism.

However, the leakage-rate testing should be mentioned and considered as a concrete aging mechanism because, in many cases, it contributes to the degradation of the concrete as the air tightness of the building may be reduced after the test is repeated for certain number of times. In 2003, HQ analysed and monitored closely the cracking density and severity that existed before a pressure testing and that developed during and after the test in 15 panels (1m x1m) located at the outside surface of the containment building. These results are described in Tables 4 and 5.

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Table 4: Summary of high pressure test results - estimated total length of cracks

• Total length of Surface cracks (mm)

Test Area (TA) Elevation Position Area After (mm2) Before During

( June,11-12) (September, 15 )

(October, 7)

TA 11 Floor at +/- 3-6 Wall 985520 6571 6655 (1%) 6655(1%)

(zone 3) 10-6 5° TA 27,1 Floor at +/- 23-6 28-3 488188 2033 2033 (0%) 2033 (0%)

(zone 2) Cfor 330° TA 43 Wall Roof (zone 2) 991616 3507 3544(11%) 3544(11%)

84-4 175° 857659 Average over 13 areas of the Test (%) or 3647 3764 (3%) 3544(11%)

0.86 m2

• 11 percent increase in total crack lengths (sum of the lengths of all cracks in a panel)

Table 5: Summary of high pressure test results - estimated widths of cracks

• Total weighted average opening lno. of points l Surface width (mm)

Test Area (TA) Elevation Area Position (mm2) Before During After (June,11-12) (September, 15 ) (October, 7)

TA 11 Floor at +/- 3'-6 Wall 0.056 0.065 (16%) 0.058(4%)

985520 (zone 3) 10-'6 l77l l77l l77l 5° TA 27,1 Floor at +/- 23'-6 28-'3 0.458 0.705 (54%) 00502 (10%)

488188 (zone 2) Cfor l22l l22l l22l 330° TA 43 Roof Wall 0.102 0.206 (100%) 0.118(16%)

991616 (zone 2) 84-'4 l36l l36l l36l 175° 857659 Average over 13 areas of the 0.180 0.230 (42%) 0.177(6%)

or Test (%) l357l l357l l357l 0.86 m2

• (15%) percent of the crack width increase; l77l no. of points (locations) measured 18

The purpose of this test is to measure the volume of air that leaked from the containment envelope during the given time of the test. The volume of leaked air must be lower than that defined by the Canadian Nuclear Safety Commission (CNSC). The locations on the surface of the containment envelope, where leakage takes place, are detected experimentally by the application of liquid soap so that the formation of bubbles on the concrete surface can be observed. These locations can be also identified by performing elaborate numerical analysis. The cracking, caused by creep and/or concrete expansion due to AAR, affects the air tightness of the containment building. As the damage to the concrete increases with time; air leakage through the structure also increases. In the numerical modelling, the finite element technique is used to develop the distribution of damage in the containment envelope.

(a) (b)

(c) (d)

Figure 9: Containment building: (a) Areas of the outer surface where the majority of the air flow (m3/s) takes place, (b) point where the leakage was detected during the pressure test, (c) and (d) damage distribution (micro and macro cracking)

In our evaluation, a structural analysis was performed using constitutive formulation that was developed at HQ and was incorporated in COSMOS/M and in ABAQUS using one of the two software packages, and then a fluid flow modelling was performed using 19

ABAQUS. A similar mesh is used in both codes. The locations of the damaged (cracked) areas and the surfaces where air leakage takes place were determined from the results of numerical analysis and from the locations where the bubbles were observed during the leakage-rate test. These locations are shown in Figure 9.

The air leaked from the containment structure, as a function of the concrete damage factor , and the flow rate - with and without the Klinkenberg effect (m3/day) - are shown in Table 6. The presented analysis was our first attempt to numerically evaluate the volume of air leaking through the containment structure. The calculated flow rates, shown in Table 6, were calibrated by comparing them to the measured values combined with the volumes of air leaking through the penetrations points in the walls and through the spent fuel exchange room.

If the total leaked air volume exceeds 0.5% of the entire air volume (V = 245 m3/day) in the containment building; i.e. the level-1 criteria; then, corrective measures (repairs) should be undertaken. A loss of 2.5% of V (1225 m3/day) is considered the impairment criteria or level-3 criteria. Unfortunately, as the NPP was decommissioned, this work has never been finalised.

Table 6: Air leakage through the containment structure The flow rate without the The flow rate with the Klinkenberg effect (m3/day) Klinkenberg effect (m3/day) only through the through only through through all the elements the elements all the elements with 1 elements with 1 After 25 years of AAR 111 4 226 6 After 30 years of AAR 162 20 296 27 6.3 In-Situ Condition Assessment 6.3.1 Introduction to the Consequences of AAR in the Concrete Structures of a NPP It is understood that AAR is a chemical processes that involves the reaction of alkali ions in cement with silica mineral aggregates, and can cause degradation in concrete. The reaction forms an alkali-silica gel that expands, when it comes into contact with water, generating hydrostatic pressure causing concrete swelling. A similar reaction involves carbonate aggregates and alkalis. At early stage of the reaction and at the onset of concrete swelling, the damage typically manifests itself as small surface cracks in an irregular polygonal pattern. Extensive damage due to alkali-aggregate reaction, over time, consists of the formation and the propagation of two types of cracks: (a) dense polygonal micro cracking as a result of the internal pressure from the alkali-silica expansion, and (b) structural macro-cracking as a result of deformation due to concrete swelling (changes of 20

the geometry) that cannot be accommodated by the structural system. The latter type of cracks is more damaging from an engineering point of view.

As mentioned before, Alkali-Aggregate Reaction in the concrete is referred to as a "Primary Degradation Mechanism". The damage resulting from this mechanism is as follows: (a) cracking of concrete, which may or may not have an effect on the air-tightness of the containment building or on the leakage of water from the reactor vault, spent fuel exchange room, spent fuel storage pool, dousing tank etc.; and (b) degradation over time of the mechanical characteristics of the concrete that may have effects on the local or overall structural behaviour, and on the capacity of anchor bolts fixing the power plant equipment to the AAR affected concrete floors, walls, etc.

Therefore, in the concrete structures of any NPP affected by AAR, almost all of the other recognised degradation mechanisms are influenced by the AAR degradation mechanism.

For example: the loss of pre-stressing depends on whether the concrete is affected by AAR. The same applies to the effects of creep; freeze/thaw cycles; degradation due to leakage-rate testing; thermal exposure following an accident; corrosion of reinforcing and prestressed steel; etc.

6.3.2 The change of post-tensioning due to combined effects of AAR and Creep The results of an elaborate study of the containment building of G-2 NPP for the assessment of the changes in the values of forces in the vertical and the horizontal post-tensioning cables are presented in this section. The changes from the original values that were developed until the time when the study was carried out in 2010 and the further anticipated changes which should take place until the year 2035, were assessed taking into consideration the combined AAR and creep effects.

The estimated changes of strain due to the combined AAR and Creep effects in the horizontal (tangential) and the vertical directions were obtained from a number of sampling points where instruments were installed. These points were placed in selected locations of the containment building wall in order to avoid the effects of boundaries at the extremities of the building. Deformation values were obtained numerically and then were used to evaluate the post-tensioning forces in the cables at the time of study (2010) and in future (2035). Estimates of the average changes of post-tensioning forces in the vertical and horizontal cables at the time of study (2010) and in the future (2035) are listed in Tables 7 to 10 The locations of the vertical cables are indicated in Figure1 and the changes in the vertical post-tensioning forces were calculated as an average of the values at five points, which were distributed along the height of the wall, i.e. each point was located at a different elevation.

21

Table 7: Estimates of the average changes of post-tensioning force in vertical cables at the time of study (year 2010)

Pint (%) (year 2010)

Average over entire height of Position Position Position Position Position Position Position Position mid-section of the wall 1 2 3 4 5 6 7 8 Creep Only (losses) -1,5 -1,57 -1,45 -1,46 -1,44 -1,54 -1,53 -1,58 Combined Creep (losses) + AAR (gains) 1,53 1,07 1,95 1,48 1,82 0,98 1,46 0,97 Table 8: Estimates of the average changes of post-tensioning force in horizontal cables at the time of study (year 2010)

Pint (%) at height from the base (year 2010)

Creep only 1,9 m 4,1m 6,8m 22,4m 38,4m 40,8m Ave. Position 1-5 (180 deg.) -0,50 -0,58 -0,66 -0,83 -0,75 -0,87 Ave. Position 3-7 (180 deg.) -0,48 -0,58 -0,70 -0,84 -0,75 -0,88 Ave. Position 5-1 (180 deg.) -0,50 -0,58 -0,65 -0,86 -0,73 -0,85 Ave. Position 7-3 (180 deg.) -0,53 -0,62 -0,69 -0,87 -0,76 -0,87 Average 360 deg. -0,52 -0,61 -0,70 -0,86 -0,77 -0,88 Pint (%) at height from the base (year 2010)

Combined Creep and AAR 1,9 m 4,1m 6,8m 22,4m 38,4m 40,8m Ave. Position 1-5 (180 deg.) 2,18 1,74 1,38 1,30 3,30 3,18 Ave. Position 3-7 (180 deg.) 2,16 1,63 1,16 1,22 3,33 3,20 Ave. Position 5-1 (180 deg.) 1,98 1,44 1,02 0,72 3,62 3,50 Ave. Position 7-3 (180 deg.) 2,03 1,46 1,00 0,90 3,28 3,16 Average 360 deg. 2,13 1,60 1,15 1,08 3,05 2,95 Table 9: Estimates of the average changes of post-tensioning force in vertical cables anticipated in 2035 Pin t (%) (year 2035)

Average over entire height of Position Position Position Position Position Position Position Position mid- section of the wall 1 2 3 4 5 6 7 8 Creep only (losses) -1,54 -1,60 -1,47 -1,49 -1,47 -1,58 -1,44 -1,61 Combined Creep (losses) + AAR (gains) 2.13 1.67 2.85 2.32 2.64 1.84 2.02 1.45 22

Table 10: Estimates of the average changes of post-tensioning force in horizontal cables anticipated in 2035 Pint (%) at height from the base (year 2035)

Creep only 1,98 m 4,1m 6,8m 22,4m 38,4m 40,8m Ave. Position 1-5 (180 deg.) -0,52 -0,60 -0,69 -0,85 -0,79 -0,91 Ave. Position 3-7 (180 deg.) -0,53 -0,64 -0,76 -0,89 -0,79 -0,91 Ave. Position 5-1 (180 deg.) -0,54 -0,63 -0,70 -0,90 -0,76 -0,89 Ave. Position 7-3 (180 deg.) -0,55 -0,63 -0,70 -0,87 -0,80 -0,91 Average 360 deg. -0,56 -0,66 -0,75 -0,91 -0,80 -0,92 Pint (%) at height from the base (year 2035)

Combined Creep and AAR 1,98 m 4,1m 6,8m 22,4m 38,4m 40,8m Ave. Position 1-5 (180 deg.) 3.68 2.95 2.45 2.19 4.22 3.85 Ave. Position 3-7 (180 deg.) 3.85 2.91 2.15 2.17 4.40 4.01 Ave. Position 5-1 (180 deg.) 3.47 2.57 1.92 1.25 4.65 4.29 Ave. Position 7-3 (180 deg.) 3.44 2.53 1.82 1.55 4.26 3.88 Average 360 deg. 3.72 2.81 2.12 1.89 4.07 3.71 The horizontal cables are semi-circular arches positioned at the middle of the wall and anchored at every second buttress as defined in the tables. The minus sign (-) indicates post tensioning loss and a plus sign (+) (or no sign) indicates post-tensioning gain.

In the presented tables two sets of results are shown: (i) the percent of post-tensioning losses in the cables if only creep effects are considered in the calculations and (ii) the percent of post-tensioning gains when both the creep and the AAR swelling effects are considered. From the presented results, it can be concluded that the effects of AAR in concrete (if the degradation of mechanical properties is not considered) are beneficial for the post-tensioning long term behaviour. It may be expected that in year 2035 the gradual increase of the initial post-tensioning force will be about 4% and 2.8% in the horizontal and the vertical cables, respectively. The variation of the post-tensioning losses/gains is influenced mainly by the non-uniformity of the walls (increased wall thickness at the closed openings initially used for entering the equipment or the existence of permanent wall openings).

The abovementioned results differ considerably from the results of the post-tensioning losses reported in the doctoral thesis authored by Granger in 1994 for studying

Électricité de France (EDF) NPPs. The difference in the results is due to two main factors: (i) the initial post-tensioning in the examined EDF NPPs produced compressive stresses in concrete between 8.5 to 9.3 MPa in the vertical direction and between 12.0 to 13.3 MPa in the horizontal direction. These values are almost two times the corresponding values for G-2 NPP and (ii) the AAR that provoked concrete swelling at G-2 is not present in the concrete of EDF NPPs. However, the post-tensioning losses obtained from the Pull-out Test of a sample beam at G-2 (corrected for two directional 23

stress fields and for lower compressive stresses in concrete) are of similar order of magnitude if compared to the losses reported by Granger.

To validate the numerical calculations, one of the post-tensioned sample beams, which were cast at the same time as the reactor building, was numerically simulated. The analysis results were compared to the corresponding results obtained from the Pull-out Test performed on the same sample beam. Under the combined effects of creep and AAR, the calculated long term post-tensioning losses were found to be about 19% to 21%,

which are similar values to those obtained from the Pull-out Test. However, these losses were found to be much higher than those calculated for and observed in the walls of the containment building.

An explanation for these differences can be given based on the behaviour of the AAR-affected confined concrete as observed from earlier tests conducted by HQ. From the tests conducted on the 600 mm blocks at the University of Sherbrooke, it was found that no expansion takes place in the direction of the post-tensioning force when the concrete is restrained in the same direction so that the imposed compressive stresses in the concrete is equal to or greater than 6.0 MPa (see chapter 2). The post-tensioning force in the beam induces compressive stress of 9.2 MPa in the concrete section. Therefore, it is only the creep effect that is present. As the inspectors who participated in the casting of the sample beams have recalled, the time of application of the full post-tensioning force was about 45 days (ti § 45 days; i.e., almost the same as the castings time). When only these parameters were considered in the numerical simulation (i.e. creep effects with ti =

45 days, average RH = 60% and AAR swelling effects are not considered), the long term post-tensioning losses were found to be about 17% , which is very similar to results of the Pull-out Test.

Reasonably accurate predictions of the long term deformation and post-tensioning losses/gains of the containment building can be made by numerical analysis when the numerical model is carefully calibrated. The obtained numerical results should be compared to the corresponding values obtained from in-situ instrumental measurements in an iterative procedure.

(a) (b)

Figure 10: Pull-out test of sample beam: (a) setup of the test and (b) test results 24

6.3.3 Micro cracking propagation throughout the prestressed concrete sections due to Creep and/or combined effects of AAR and Creep It is important to be noted that the numerical analysis of creep effects in the post-tensioned concrete of the G-2s containment building showed that due to elevated creep (based on high post-tensioning stresses) an extensive micro cracking propagates throughout the concrete sections. The ducts and the post-tensioning cables create a non-homogeneous concrete section. The stress fields due to the positioning of the cables are also non uniform. The RH throughout the thickness of the walls is not uniform; therefore, the Creep is not uniform as well. These factors are minor and may not be important if there is no requirement of air tightness of the containment building under internal pressure. More elaborate study revealed that the magnitudes of the slightly non-uniform compressive stresses in the concrete (induced by the 2D prestressing) which are affecting the creep should not be high. A parametric study has demonstrated that the post-tensioning stress in the concrete section should not be higher than 6.0 to 6.5 MPa in each direction in order to avoid the creation of micro cracking throughout the walls of G2 reactor building envelope under the combined effects of AAR and Creep. These values are expected to be slightly lower if only Creep effects are present.

The damage (propagation of the micro cracks) is not uniform in the structure. The main reason is the non-uniform expansion of concrete due to the uneven distribution of the humidity in the entire structure. Based on the magnitude of the evaluated humidity distribution, forty-two different rates of the concrete expansion were defined in the calibration process (matching the rates of expansions obtained from the built-in extensometers) and employed in the analysis. Moreover, the deterioration of the mechanical properties of concrete over time has great importance on the effect of this degradation mechanism. The results obtained from the parametric studies have demonstrated the significant effect that the predicted tensile strength of concrete has on the time of onset and the subsequent development of the structural cracking.

Example: assessing the value of the tensile strength of the concrete as being 13% higher than predicted at the end of 2035 (from ft2035 = 2.0 MPa to ft2035 = 2.25 MPa) will delay the damaging effect of cracking on the air tightness of the confinement for three (3) to seven (7) years based on the locations of the affected areas.

The results indicating the distribution pattern of the coefficient of deterioration due to self-weight, post-tensioning and the combined effects of AAR and creep are presented in Figures 11 and 12, where the zones of micro cracks ( 0.6) and macro cracks

( 1.0) are identified by different colors .

25

(a) (b)

Figure 11: Distribution of coefficient of deterioration due to self-weight, post-tensioning and after (a) 17 years (2002) ; (b) 32 years (2017) of combined AAR and creep effects ; Horizontal section at z=1.41m (a)

(b) (c) (d)

(e (f) (g Figure 12: Distribution of coefficient of deterioration due to self-weight, post-tensioning and after: a) 12, b) 17, c) 18, d) 20, e) 25, f) 32 and g) 35 years of combined AAR and Creep effects ; Horizontal section at z=1.41m. Detail of the wall opposite of the opening 6.3.4 Behaviour at the junction of the perimeter wall and the base slab caused by different expansion of the wall and the slab (due to different confinement)

The results of the analysis revealed that the early cracking, even without subjecting the containment building to internal pressure, in the areas close to the bottom of the walls was affected largely by the expansion of the base slab.

Table 11 shows strain rates at different locations of the base slab as measured by extensometers; together with the corresponding results obtained by FE analysis. (2010);

26

Table 11: Strain rates at different location of the base slab FE estimates Measurements FE estimates Measurements Radial Radial Tangential Tangential Location rate (m/m/day)

( rate (m/m/day) rate (m/m/day)

( rate (m/m/day) exterior 0,02007 N/A 0,03862 0,03173 3,3m interior 0,02121 0,0213 0,03880 0,0235 exterior 0,03705 0,0345 0,03920 0,0343 6,6m interior 0,03920 N/A 0,03919 0,026 exterior 0,04025 0,0333 0,03863 N/A center interior 0,03952 0,0379 0,03970 N/A Note : the locations of the measuring points are given as distance from the axis of the perimeter wall The expansion (growth) rate for the wall as obtained from the extensometers is: from =

0.0016 m/m/day ( = 0.6 m/m/y) close to the base slab, to a maximum value of =

0.0227 m/m/day ( = 8.3 m/m/y) near the ring girder, while the vertical expansion rate varies from = 0.0016 m/m/day ( = 0.6 m/m/y) to = 0.0425 m/m/day ( = 15.5 m/m/y). The difference between the rates of tangential expansions of the base slab (near the wall) of = 0.0235 m/m/day and the horizontal rate of expansion of the wall =

0.0016 m/m/day is of order of 15 times. The perimeter of the wall cannot increase by the same amount (having expansion rate of = 0.0016 m/m/day) as the base slab (having expansion rate of = 0.0235 m/m/day in tangential direction).

(a) (b)

(c) (d)

Figure 13: Distribution of damage factor at the connection of perimeter wall with base slab due to self-weight, post-tensioning and after: a) 20, b) 30, c) 40, d) 45 years of combined AAR and Creep effects (2035) 27

The difference between the rates of tangential expansions of the base slab (near the wall) of = 0.0235 m/m/day and the horizontal rate of expansion of the wall = 0.0016 m/m/day is of order of 15 times. The perimeter of the wall cannot increase by the same amount (having expansion rate of = 0.0016 m/m/day) as the base slab (having expansion rate of = 0.0235 m/m/day) in the tangential direction.

In fact as the wall cannot follow the radial expansion of the base slab, it deforms at the base. As a result, multiple vertical cracks of up to 1.5 meter in height and spaced 1.0 to 1.5 m apart were created on the entire perimeter starting from the base slab. The observations were confirmed by analysis and the cracks were detected at the same locations as shown in Figure 13. In a similar manner, multiple horizontal cracks were observed on the wall up to 4.0 m above the base as shown in Figure 14.

(a) (b)

(c) (d) (e)

Figure 14: Deformed shape: connection of perimeter wall with base slab due to self-weight, post-tensioning and after: a) 0, b) 20, c) 30 d) 40 and e) 45 years of combined AAR and Creep effects (2035) 6.3.5 Concrete splitting in the plane where the ducts with post-tensioning cables are located (due to Creep and/or combined effects of AAR and Creep)

The results reveal areas (regions) of the containment with relatively high tensile stresses perpendicular to the planes of the post tensioning cables. The continuous loss of tensile strength of the concrete as a result of AAR may provoke concrete splitting parallel to these planes as it was the case at the Montreal Olympic Stadium (constructed using the same concrete aggregate used at G-2). The splitting of concrete may occur also in a containment concrete structure not affected by AAR if the prestressing is relatively high.

The study of the G-2 containment building, excluding the AAR effects but applying 28

prestressing forces in horizontal and vertical direction that are producing higher (double) compressive stresses in the concrete walls, indicated very high probability of concrete splitting. (Question to the ASCET participants: Is this probable cause of the concrete splitting at the Cristal River NPP in Florida). Before the decommissioning of G-2, Hydro-Quebec have started an elaborate study to evaluate: (a) the Ultimate Pressure Capacity of containment building, and (b) to indicate all the regions where a concrete splitting will probably appear. It was also planned to proceed with in-situ testing (hammering and drilling the wall) to detect the splitting and to confirm the validity of the numerical analysis. Study of this type was done in the evaluation of the safety of the Montréal's Olympic Stadium where the areas of concrete splitting were located and the corrective measures were taken.

The preliminary results of the analysis intended to locate the areas of probable concrete splitting are presented in Figure 15 below.

Figure 15: Preliminary results of an analysis intended to locate the areas of probable concrete splitting The tensile stresses of the concrete in the areas where splitting may occur, are calculated between 0.5 MPa and 1.0 MPa. for the year 2010, and between 1.0 MPa to 1.5 MPa for the year 2020. The reduction of the tensile concrete strength, as a result of the degradation of concrete properties due to AAR, was obtained, based on laboratory tests on extracted cores, as 2.5 MPa for 2010 and conservatively estimated to be 2.0 MPa for 2035. Based on the previous experience, once numerically located and confirmed by in-situ tests, the regions where the splitting occurred are relatively easy to be repaired.

The numerical results of the G-2 containment building , by assuming 25% increase of present compressive concrete stresses due to prestressing, predicted occurrence of many regions where concrete splitting.

29

The studies on possibility of concrete splitting in prestressed containment buildings not affected by AAR, demonstrated that prestressing induced compressive stresses in the concrete of 8.0 to 10 MPa may cause concrete splitting.

6.3.6 Effects of AAR on non-prestressed reinforced concrete in contact with water at G-2 NPP The non-prestressed reinforced concrete structures such as: Spent Fuel Storage Pool, Spent Fuel Exchange Room, Reactor Vault, etc. are more affected by the concrete swelling (growth) than prestressed structures. The concrete growth rate varies between 60 and 100 m/m/y. As a consequence of the elevated growth extensive cracking and large deformations are observed (Fig. 16). Hence, water leakage, distortion of the bases of the equipment serving to manipulate (exchange or store) the spent fuel rods, jamming of doors etc. are disturbing the normal operation of the NPP. The degradation of the concrete strength in some cases is affecting the capacity of steel anchors attaching the equipment and the pipes to the concrete floors and walls.

Piscine de stockage Piscine de stockage Piscine de stockage Piscine de stockage Figure 16: Observed damage at the Spent Fuel Storage Pool The spent fuel pool consists of a reinforced concrete structure with outer plan dimensions of approximately 25.32 m x 14.12 m and a height of 9.14 m. The wall thickness of the reinforced concrete structure varies with the location of the wall but is generally 1.22 m to 1.68 m at the buttresses. The spent fuel pool is founded on a layer of crushed gravel underlain by rock. The evolution of damage in concrete due to AAR was examined in detail and documented in studies conducted specifically in relation to the spent fuel pool at G-2.

The damage evolution in the concrete component of the reinforced concrete accounts, through a very rigorous elasto-plasticity formulation (see chapter 3), for the development of micro-cracks and macro-crack evolution during AAR. Furthermore the modelling also accounts for the time-dependent evolution of damage due to AAR during 29 years, following 7 days of heating to 1000C caused by failure of the cooling system. The first step of the analysis included evaluation of structural integrity and safety of AAR affected pool due to all loading conditions (static and seismic). The study was further extended to assess the water 30

leakage. Since attention is focused on the estimation of leakage through the spent fuel pool, the objective of the exercise was to ascertain from the AAR studies the regions of the structure that have undergone both micro-cracking or macro-cracking. For example, Figure 17 illustrates the induced damage through the contours of the Stress Intensity Factor (). The stress intensity factors provide an estimate of the intensity of AAR damage that can be used to estimate the level of permeability alteration in the concrete. Figure 18 illustrates the Flow rate contours and Fluid velocity vectors obtained from the unconfined flow seepage analysis through the storage pool.

(a) (b)

Figure 17: Stress Intensity Factor representing damage or areas of micro and macro cracking: (a) after 29 years of AAR and (b) 29 years of AAR and water temperature increase effects (T=100 oC).

(a) (b)

Figure 18: (a) Flow rate contours (m3/s) and (b) Fluid velocity vectors; at different points on the surfaces of the pool after damage induced by AAR (29 years) and water temperature increase effects (T=100 oC),

The computational modelling is performed to accomplish the following objectives:

(i) To establish the steady flow rate through the storage pool structure under the differential total head of 7.7724 m when the permeability alterations are only due to AAR effects and the water temperature is 25 0C.

(ii) The crack width in the computations of permeability in regions with 1 , at the water temperature of 25 0C, is adjusted such that the leakage from the storage pool corresponds to 240 litres/day. This corresponds to an effective crack width of 0.018 mm and the number of damaged elements is 1442.

(iii) The effect of thermal damage in increasing the number of cracked elements is considered. The thermal damage results in an increase in the number of damaged elements to 4333. Maintaining the effective crack width at 0.018 mm 31

for all elements where 1 , the leakage through the damaged structure is calculated at water temperature T=100 0C .

(iv) Keeping the number of damaged elements at 4333 and the crack width in the damaged elements with 1 at 0.018 mm, the temperature is decreased to 25 0

C and the leakage through the damaged structure is calculated. The decrease in the temperature will result in the increase of the viscosity of water and a decrease in the flow rate. The estimated flow rate is the leakage after a failure and successful repair of the cooling system.

(v) Possible thermally-induced alterations in the crack widths are not taken into consideration.

The computational results of the study are summarized in Table 12.

Table 12: Estimated rate of leakage through the walls and the base slab of the pool Viscosity Density of Leakage rate Type of Damage Number of Cracked Temperature (oC)

Water w (Litres Leaky Surfaces (Pa.s) 3 /day)

Elements (kg/m )

AAR (29 yrs) 1442 25 0.000891 997.1 240* all surfaces AAR(29 yrs) + all surfaces 4333 100 0.000282 958.4 1870 Thermal damage AAR(29 yrs) base slab only 4333 100 0.000282 958.4 1430

+Thermal damage AAR(29 yrs) all surfaces 4333 25 0.000891 997.1 620

+Thermal damage AAR(29 yrs) base slab only 4333 25 0.000891 997.1 470

+Thermal damage

• The currently measured total rate of leakage, In order to evaluate structural safety under combination of various static and seismic loads the similar studies were undertaken to analyse the Spent Fuel Exchange Room and the Reactor Vault. The examined problems of the aging of AAR affected concrete of these two structures will not be presented here. The report is already exceeding the limitation of the number of pages set by the ASCET Organizing Committee.

6.3.7 Effects of AAR on the Turbo-Generator foundation at G-2 NPP The Turbo-Generators foundation is non-prestressed reinforced concrete structure with outer plan dimensions of approximately 70 m x 20 m and a height of 20 m. The foundation was constructed in 1975 and immediately after its construction settlement of the structure was detected. For normal operation some adjustments of the turbo-generator base plates was required. Several years after the construction the rock masses bellow the foundation stabilised, therefore, the settlements stopped in 1988-89. The foundation as all the other structures at G-2 NPP is affected by the concrete swelling (growth). Based on the measurements obtained from the Invar wires, installed after the effects related to AAR were detected, the concrete growth rate wearies between 40 and 60 m/m/y. As a consequence of the elevated growth extensive cracking and large deformation of the 32

structure was observed, it is present today and is estimated to continue in the future. The extent of the damaged areas of the concrete foundation after 14 years of settlement and AAR growth is shown in Fig. 19. The concrete growth induces uneven (differential) displacements (much larger than the prescribed tolerances) which do not permit normal operation without frequent adjustments. Each intervention requires period of time for adjustments so for 670 MW installed capacity generator it represents a considerable production lost for Hydro-Quebec. To optimise the intervention time for the future displacements of the foundation at the points of the turbo-generators supports, an elaborate numerical analysis was performed and the model was calibrated using the results of the installed instruments. Based on the predicted future displacements (Figure

20) a proposal for the adjustment of the turbo-generator supports was submitted to General Electric. In addition, a dynamic analysis of the cracked foundation was performed. The purpose of the analysis was to evaluate the possibility of eventual future vibration of the unit supported by cracked (deteriorated) foundation. The vibration of hydroelectric generating units was observed in some of the AAR affected powerhouses.

Figure 19: After 14 years of AAR; Turbo-generators - foundation L§70 m; b § 20 m; h § 20 m Figure 20: Gentilly-2 Turbo-generators - foundation; Comparison and prediction of vertical displacements until 2033 33

6.3.8 Structural integrity under seismic excitation of AAR affected containment The structural behaviour of the containment building already weaken by respectively 25 or 50 years of progressive AAR was examined under the seismic excitation. Two separate analyses considering slightly different tensile strengths of the concrete, at the end of the planned life extension at 2035, were undertaken. In the seismic analysis the predicted tensile strengths of concrete at 2035 were ft2035 = 2.0 MPa and 2.25 MPa.

The structure was analyzed using dynamic procedure with the AAR constitutive model containing inelastic constitutive relations defining the mechanical behaviour of plain concrete affected by swelling due to AAR and exposed to the other prescribed loadings.

The reinforcement was modeled using two different approaches: (a) in discrete manner and (b) as volume fractions in a homogenized approach. The effect of the confining pressure was included directly in the evolution law of the program.

The time histories used for the non-linear seismic analyses were developed by Atkinson (Hydro Quebec, 2009-b) employing six (06) ground motion time history records to match the target spectrum for the Gentilly 2 NPP site (1/10 000 p.a.), on rock, for the 84% confidence level. Three (03) sets are obtained from seismological simulations and three (03) other sets are obtained from spectral matching technique by modifying the frequency content of historical recorded ground motions.

The effects of an eventual earthquake in the region of G2 on the propagation of the damaged (cracked) concrete regions at present (2010) are shown in figures 21 and 22. It can be seen that during an earthquake, even though the structural integrity is not affected, the cracked regions extend at the junction of the wall and the base slab. The probability of loss of air tightness after an earthquake (2011) is very high. This probability will be even higher in the coming years (years 2011-2035) since the concrete cracking will gradually increase. This is confirmed by the results obtained from the analysis simulating expansion of the concrete followed by an earthquake in 2035 (Figures 23 and 24).

(a) (b)

Figure 21: Distribution of coefficient of deterioration due to self-weight, post-tensioning and after 25 years of combined AAR and creep effects (a) before and (b) after earthquake. (Analysis of current structural state; 2010); Interior view.

34

(a) (b)

Figure 22: Distribution of coefficient of deterioration due to self-weight, post-tensioning and after 25 years of combined AAR and creep effects (a) before and (b) after earthquake (Analysis of current structural state ; 2010); Exterior view.

(a) (b)

Figure 23: Distribution of coefficient of deterioration due to self-weight, post-tensioning and after 50 years of combined AAR and creep effects; (a) before and (b) after earthquake (Analysis of current structural state ; 2035); Interior view.

(a) (b)

Figure 24: Distribution of coefficient of deterioration due to self-weight, post-tensioning and after 50 years of combined AAR and creep effects; (a) before and (b) after earthquake (Analysis of current structural state ; 2035); Exterior view.

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6.3.9 Initial cracking - Construction of the containment structure For the construction of the containment building (April 1974 to June 1976) the concrete prepared with 450kg/m3 type 10-Portland cement having cumulative heat liberated in seven days of 335 KJ/kg was used. The construction of the 0.914 m (3 ft) thick and 42.22 m (138.5 ft) high walls was done in 15 days using slip-form technique. The construction sequences and the ambient temperatures during concrete poring, as well as the variation of the mechanical properties of the concrete for the same period of time were properly planned. Any sudden change of the ambient conditions (period of low temperature) were not predicted.

The temperature due to hydration process was calculated and instrumented from 60°C to 75°C at the middle of the wall thickness. The hydration heat distribution is shown in figure 25.

Twenty days following the final concrete pore the imprisoned temperature in the wall was between +40°C and +50°C. At the same time (first two weeks of November) the recorded ambient temperature at G2 was -18.2°C. This unpredicted low temperatures formed a temperature gradient between the surfaces and the core of the wall of T = 58 to 68°C. Therefore, crack appeared on the concrete surfaces (Fig. 26).

Figure 25: Heat distribution (end of the wall construction) 36

The (micro) cracks having width of 0.5 to 1.0 mm on the surface, confirmed from the extracted samples, have propagated between 50 and 100 mm in the wall. This cracking (on both surfaces) was observed with the inspection done following the post-tensioning of the cables. The observed cracking did not affect the ultimate strength however, to insure air-tightness of the containment structure the precautionary corrective intervention took place after the construction was finalized. The internal surface was covered by epoxy coating and the external surface with watertight flexible membrane. Therefore, the concrete aging started right from the beginning. It was observed later that these initial cracks defined the cracking network on the walls and these cracks were the ones that propagated even further (150 to 200 mm) into the concrete wall. The analysis performed to identify the air tightness of the containment building confirmed that initial cracking played an important role in the formation and the propagation of cracks caused from other sources and contributed to the more advanced than expected deterioration (localised concrete spalling) at the outside surface of the containment.

The initial thermal cracking (during and right after construction) may not be considered as a concrete aging mechanism, however, it plays, based on our experience, an important role in the premature concrete aging of containment building located in cold regions.

(a) (b)

Figure 26: Regions of initial surface cracking (a), Penetration of surface cracks (b) 6.3.10 Freeze-Thaw cycles - Concrete cracking and their propagation, delamination and concrete spalling The thermal stresses under winter-summer thermal cycles (i.e., seasonal temperature fluctuations), is a degradation mechanisms of the exposed concrete resulting in formation (at the surface) and propagation (within the concrete element) of cracks. The propagation of the existing (cracking initiated during construction period) as well as newly formed surface cracks was examined by: (a) numerically simulating 50 to 75 thermal cycles typical for the extremely cold Canadian winters, and by (b) precisely monitoring the cracked surfaces (Figure 27). The observation of the cracking was done on 1.00m x 1.00m panels on 15 locations determined from the results of the numerical analysis. The time curves representing the ambient temperature history on the outside face, the inside faces and across the concrete sections were evaluated from the climatic data for the 37

region provided by Environment Canada. The results showing the cracked concrete areas after twenty thermal cycles, including one pressure test, are given in Fig. 28.

Figure 27: In-situ monitoring of the cracked surfaces The delamination and concrete spalling as a consequence of the thermal cycles was observed on the annular beam and the sides of the buttresses where the heads of post-tensioning cables are located.

Results of the nonlinear analysis for aging mechanisms cracking due to freeze-thaw cycles, 50th winter Exterior view Cross-sectional view 70 Interior view Enlarged cross-sectional view Figure 28: cracked concrete areas after twenty thermal cycles, including one pressure test 38

7. STRUCTURAL ACCEPTANCE CRITERIA FOR STRUCTURES WITH PATHOLOGIES/DEGRADATION MECHANISMS No Contribution

8. STRUCTURAL REPAIR 8.1 Introduction According to our experience at G-2 NPP, the structures of a NPP affected by AAR do not necessarily require major repairs. Generally speaking, the structural integrity of these structures is not compromised due to the AAR effects. However, in a NPP with AAR-affected concrete, it is the uninterrupted production which is often challenged. Some of the problems and inconveniences which may arise in these circumstances are as follows:

Areas, where extensive micro/macro cracking is observed, are main cause of water or air leakage. Cracking in areas like the Spent Fuel Storage Pool, Spent Fuel Exchange Room and the tunnel connecting them, is likely to cause leakage. The appropriate repair in this case is usually made by lining the affected areas by a flexible membrane. As for the air tightness, it can be restored by adding layer of NORMAC on the concrete surfaces on the same side of the high air pressure. These interventions are relatively easy to do and their cost is relatively low. However, the cost of the production loss during repairs may be substantial.

Structural deformation as a result of concrete swelling: adjustments will likely be needed; particularly, for certain equipment: (a) the turbo-generator: the addition or the removal of steel shims with different thicknesses under the base plates represents an easy and fast-to-do solution; however, it may require careful planning to provide margin for further levelling due to future expansion; (b) accommodating concrete deformation by making adjustments to the anchorage and the base plates of important equipment to ensure accurate leveling or plumbness; (c) adjustments for the main pipe supports attached to the concrete; (d) grinding of the concrete floors at the containment building entrance; (e) grinding of deformed door frames and other wall and slabs openings; (f) adjustments of the pipes which are anchored to the Reactor Vault in order to insure safe operation of the CALANDRIA.

Degradation of the mechanical properties of the AAR-affected concrete: this problem requires careful evaluation and - when necessary - additional reinforcement may be necessary. The replacement of some concrete anchor bolts may be needed in order to compensate the reduction in concrete shear strength.

Splitting of concrete in the areas of high two-directional prestressing: this problem may require pinning the concrete wall or element using concrete anchors.

A more detailed description of the different repair strategies should be included in the final report.

39

8.2 Gentilly Spent Fuel Pool Lining Project The Spent Fuel Storage Pool of G-2 NPP is one of the concrete structures of the plant which are the most affected by the AAR. This is due to the fact that it is constructed from non-prestressed reinforced concrete as well as it is in a direct and continuous contact with water. The limited confinement and the high percentage of relative humidity are the main reasons for the elevated expansion rates observed in the storage pool. In fact, the concrete expansion rate in the different structural elements of the pool varies between 60 and 100 m/m/year. As a result of the elevated expansion rates, signs of the deterioration of the concrete structure were clearly observed. The deterioration manifested itself in the form of extensive cracking, large deformations, water leakage, etc.

In order to reduce the damage caused by the AAR in the G-2 Spent Fuel Pool, a project for lining the pool with a liner waterproofing system was carried out in the recent years.

The project consisted of covering the four walls of the pool, including the corners between the walls themselves and the corners between the walls and the bottom slab by a flexible EPDM geomembrane that was mounted to the walls of the pool using a fastening system made of stainless steel components, profiles, clamps and anchors. The project also included the installation of a water leakage detection system as well as a drainage system for water collection.

Figure 29: The EPDM geomembrane was mounted to the walls of the Spent Fuel Storage Pool using a fastening system made of stainless steel components 40

9. CONCLUSIONS AND RECOMMENDATIONS FOR FUTURE WORK The conclusions presented in this report are based on the experimental and theoretical work that was carried out by Hydro-Quebec during the last forty years during the course of conducting various structural evaluations for many of its existing hydroelectric power plants as well as for the various studies and assessments that were carried out by HQ for Gentilly-2 NPP. The following is a summary of these conclusions:

• In order to ensure the safe operation of a nuclear power plant, it is essential that the effects of age-related degradation mechanisms on NPPs structures, systems and components, to be assessed and managed during current and subsequent life cycles.

• With the exception of the alkali-aggregate reaction in concrete, many of the known degradation mechanisms discussed in the technical literature are not present at G-2 NPP and at the other hydroelectric power plants owned by HQ. However, some of these mechanisms were encountered at certain existing plants but none of them was found to have a significant effect on the structural integrity or the normal operation of these power plants. It is believed that this is the case for the majority of NPPs as well for the following reasons:

o The concrete, used in the construction of NPPs structures, particularly containment buildings, is of high quality. The water/cement ratio for the concrete used in these structures is usually around 0.42 which is sufficient to ensure the durability of the concrete.

o Despite that the concrete is affected by the AAR at G-2 NPP, the compressive strength of concrete cores extracted from the storage pool and the confinement building, twenty years after the plant was commissioned, was found to be 22%

and 40%; respectively, higher than the corresponding values obtained from the standard tests after 28 days.

o Regardless of the quality, mechanical characteristics and the durability of concrete used in the construction of NPPs, the following degradation mechanisms are often present:

The initial surface cracking, due to hydration temperature (in case of inappropriate concreting or curing), is only 10 to 15 mm in depth. The same cracks propagate further due to the freeze/thaw cycles. However, these cracks do not have significant effect.

Shrinkage contributes to the extension of the cracks created by the initial hydration. It also produces some additional cracks at an early stage. However, these cracks do not have significant effect as well.

Cracking caused by concrete creep may influence the air tightness of the containment building. This is in the case when prestressing is high enough to produce compressive stresses in excess of 8 MPa.

• The longevity of the NPPs structures is the result of:

o NPPs are exceptionally demanding; therefore, a strict adherence to the quality control plan during the design and construction of the plant is required in order to ensure a high quality of the NPPs structures.

o The Canadian Nuclear Safety Commission (CNSC) and the United States Nuclear Regulatory Commission (U.S. NRC) monitor closely the performance of all 41

NPPs. They react promptly in case that any abnormality is observed. As a result of the continuous efforts made by the International Atomic Energy Agency (IAEA) and the similar organisations, it may be assumed that the safety requirements imposed by the Nuclear Regulatory Authorities in other countries are similar to those in force in the U.S. and Canada.

• The experience gained from the evaluation of G-2 NPP; has been enriched by the numerous exchanges we made with other NPPs owners (mainly CANDU and EDFs plants). Based on this experience, it can be concluded that while many degradation mechanism are affecting different types of structures such as bridges, hydroelectric powerhouses and dams, tunnels, etc.; however, these mechanisms are not present or not affecting to the same extent NPPs. Therefore, the notion that NPPs are safely operating for the initial period of 30 to 40 years - and before any re-evaluation is required - should be considered as the default status.

• From the studies on AAR affected G-2 NPPs structures the following may be concluded:

• AAR in concrete is a primary degradation mechanism.

• The damage resulting from this degradation mechanism is:

o Cracking of the concrete o Structural deformation due to concrete growth o Degradation over time of the mechanical characteristics of the concrete

• Almost all recognised degradation mechanisms for concrete structures of most NPPs affected by AAR are influenced by the AAR degradation effects.

• Laboratory and in-situ material testing is essential in order to define properly the mechanical characteristic of the affected concrete. Tests on small samples may adequately provide this information.

• However, the tests on small samples cannot adequately provide certain information related to real structures such as: heterogeneity of the volumetric expansion, development of stresses in concrete due to the swelling of the silica gel, effect of confinement, etc.

• More meaningful conclusions, related to the behaviour of AAR-affected concrete in real structures, are obtained by multi-axial testing of relatively large concrete blocks and by conducting in-situ material testing.

• Regardless of how elaborate these tests are; however, they cannot provide answers to the most important question which is as follows: at what level of safety and operational reliability this AAR-affected structure stands? Unfortunately and after 40 years of research, the approach of conducting identical material tests as a main tool of investigation is still the most common. In fact, from structural engineering point of view, the question can be answered mainly by performing adequate structural analysis.

• In an existing and operating NPP, many areas are inaccessible, while only a restricted access is available to many other locations in the plant. Conducting an analytical study maybe the only feasible evaluation method to assess the state of degradation and the safety level at a NPP. In this case, an adequate elaborate numerical analysis is 42

essential to evaluate the state of the NPP and to provide a high level of confidence in the obtained results and the arrived at conclusions.

• Many of the commercially available software packages can be used to perform an adequate and an elaborate structural analysis for concrete structures which are not affected by AAR. To the contrary, it is not the case when it comes to analyzing AAR-affected concrete structures. This is due to the fact that these codes are lacking material models with adequate constitutive formulation that is particularly suitable for simulating the behaviour of AAR-affected concrete structures. In many cases, the software can be used to arrive to some reasonable displacements solutions, however; the associated results may not be representative to the state of the analyzed structure in terms of the severity of damage, state of stress, crack locations and predicted behaviour. Therefore, many of these models should be re-examined carefully.

• Based on HQ's experience and practice, the essential requirements of the concrete/reinforced concrete constitutive model accounting for the chemo-mechanical interaction in AAR-affected concrete, are as listed below. Advanced Finite Element (FE) codes should incorporate material models that satisfy these requirements in order to provide an adequate tool to analyze AAR-affected structures:

o Adequate description of the kinetics of the reaction o General failure criterion, provision for the development of irreversible deformations, general criterion for the onset of macro-cracking in both compression and tension regimes o Degradation law for strength and deformation characteristics o Proper description of propagation of damage in both tension and compression regimes (viz. homogenization incorporating a characteristic dimension, XFEM or similar) o Constitutive relation for the interface material relating the velocity discontinuity to the traction vector.

• Periodic integrated Leakage-rate Tests conducted at G-2 NPP have had negative effects on the containment building as they caused the cracks to propagate. Therefore, the frequency of these tests should be re-evaluated even though the test is considered non-destructive.

• The conclusions of the in-situ condition assessment (chapter 6.3) can be summarised as follow:

o The long term post-tensioning losses in the cables can be predicted with reasonable accuracy by numerical analysis if only creep effects are considered.

o The long term post-tensioning gains in the cables can be predicted with reasonable accuracy by numerical analysis and from the in-situ tests if the combined actions of creep and the AAR swelling effects are considered.

o Micro cracks may propagate throughout the concrete envelope of the containment building due to creep effect, elevated prestressing or the combined effects of creep and AAR swelling.

43

o Local disorders, such as cracking and loss of air tightness, may occur due to uneven expansion of structural components (wall and base slab junction; wall and dome junction).

o High prestressing - whether alone or combined with the AAR effects - may cause concrete splitting in the planes which are parallel to the wall surfaces.

• The following recommendations should be considered for future work related to NPPs affected by AAR:

o Concrete laboratory testing should include tests of large cubes representing the concrete of a real NPP structure. The samples should be reinforced with two families of reinforcement. The reinforcement should provide internal confinement so that the blocks should represent the reinforced concrete of the majority of NPP structures. The outside confinement (prestressing) may be imposed by testing the block with 2D press. The humidity in the block should not be uniform (RH100%). The test should not be accelerated, 38oC material test used to confirm presence of AAR. The testing time (these are not accelerated tests) should be for 10 years as a minimum. To have an adequate numerical analysis the rate of expansion should be defined more precisely in function of the combined effect of relative humidity and confinement (internal or external).

• Fortunately, 139 extensometers were installed in the base slab, the wall and the dome of the containment building at G-2 NPP. This has allowed us to have adequate calibration of our numerical models. Providing adequate instrumentation should be considered in any new NPP. However, excessive instrumentation should be avoided.

o The same type of tests should be performed on non-AAR affected concrete in order to confirm by testing the results obtained numerically. This should be of help to see if the non-uniform creep (uniaxial and biaxial) is the reason for concrete splitting and air leakage o The need for repeated in-situ and non-destructive testing activities should be re-evaluated from the point of view of a structural engineer.

44

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3/4 Pande, N. G. and Shin, H-S. 2003, Implementation of a Homogenized Reinforced Concrete Model for COSMOS, Centre for Civil and Computational Engineering University of Wales Swansea, UK, May 2003.

3/4 Pietruszczak, S., and Oulapour M. 1999. Assessment of dynamic stability of foundations on saturated sandy soils J. of Geotechnical and Geoenvironmental Eng. , ASCE, 125, 576-582.

3/4 Pietruszczak S., Gocevski V. 2002. On rehabilitation of concrete structures affected by alkali-silica reaction, Intern. Journ. Comp. Civil & Struct. Engng, vol. 1, No.3.

3/4 Mitchell, D., Cook, D. W., 2003. Lift-Off Test and Recovery of a Portion of Duct from Post-tensioned Beam (Gentilly 2), A Report to Hydro-Quebec, Civil Engineering and Applied Mechanics, McGill University, Montreal.

3/4 Gocevski, V. 2006. Constitutive Model for Evaluation of Nuclear Containment Structures, International Symposium, Fontevraud-6, SFEN French Nuclear Energy Society, Paper A006-T10, 18-22 September 2006, Fontevraud France.

3/4 Gocevski, V. 2005. Numerical Model for Simulation of Aging Mechanisms of Nuclear Containment Structures, International Symposium, Concreep-7, SFEN French Nuclear Energy Society, Paper A006-T10, 18-22 September 2005, Nantes, France.

3/4 Gocevski, V. 2004. Centrale nucléaire de Gentilly-2; Effets de l'écrasement d'un avion commercial - Bâtiment du Réacteur, Volume 1 du Rapport Final, Hydro-Québec TANIA-12242-001, juin 2004.

3/4 Gocevski, V. 2003. Centrale nucléaire de Gentilly-2; Analyse du comportement du Bâtiment du Réacteur, Volume 1 du Rapport Final, Hydro-Québec TAYAA-12242-001, avril 2003.

3/4 Gocevski, V. 2010. Gentilly 2 NPP - concrete aging effects on long term pre-stress losses and propagation of concrete cracking due to pressure testing, International Symposium, Fontevraud-7, SFEN French Nuclear Energy Society, Paper A158-T10, 26-30 September 2010, Avignon, France.

3/4 Gocevski, V. 1998. "Power Plant Repair", Concrete Engineering International, Vol. 2, NO. 2, pp. 11-14, 3/4 Gocevski, V. 1989. << Elasto-plastic model for cemented and pure sand deposits >>,

Computers and Geotechnics, vol. 7, pp. 155-187.

3/4 Gocevski, V. 2010. "Effects of Alkali-Aggregate Reaction in Concrete on Long Term Pre-stressing Losses", 13th International Symposium of the Serbian Association of Structural Engineers, 22-26 september 2010, Zlatibor, Republic of Serbia.

3/4 Gocevski, V. 2010. Gentilly 2 NPP - concrete aging effects on long term pre-stress losses and propagation of concrete cracking due to pressure testing, International Symposium, Fontevraud-7, SFEN French Nuclear Energy Society, Paper A158-T10, 26-30 September 2010, Avignon, France.

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3/4 Gocevski, V. 2010. "Effects of Alkali-Aggregate Reaction in Concrete on Long Term Pre-stressing Losses", 13th International Symposium of the Serbian Association of Structural Engineers, 22-26 september 2010, Zlatibor, Republic of Serbia.

3/4 Gocevski, V., 2009, "Solution to Complex Structural Problems through Advanced Numerical Methods", 11th National and 15th International Sciantific Meeting, iNDiS 2009, Novi Sad, Republic of Serbija, November 25-27, 2009.

3/4 Gocevski, V., 2009, "Solution to Complex Structural Problems through Advanced Numerical Methods", 11th National and 15th International Sciantific Meeting, iNDiS 2009, Novi Sad, Republic of Serbija, November 25-27, 2009.

3/4 Gocevski, V., 2008, "Evaluation of AAR Affected Powerplant Exposed to Turbine Vibration or Earthquake", 15th International Seminar on Hydropower Plants from November 26th to November 28th 2008, Vienna, Austria 3/4 Gocevski V, 2008, "Ageing of Concrete Considered in the Safety Evaluation of a Nuclear 3/4 Containment Structures", OECD-NEA Workshop on Aging Management of Thick Walled Concrete Structures, Prague, Czech Republic, 1-3 October 2008 3/4 Gocevski, V. 2006. Constitutive Model for Evaluation of Nuclear Containment Structures, International Symposium, Fontevraud-6, SFEN French Nuclear Energy Society, Paper A006-T10, 18-22 September 2006, Fontevraud France.

3/4 Gocevski, V. 2005. Numerical Model for Simulation of Aging Mechanisms of Nuclear Containment Structures, International Symposium, Concreep-7, SFEN French Nuclear Energy Society, Paper A006-T10, 18-22 September 2005, Nantes, France.

3/4 Gocevski, V., 2004, "Constitutive Model for Reinforced Concrete Applied in the Analysis of the Gentilly-2 Reactor Building", Sixth International Conference on Simulation Methods in Nuclear Engineering, 2004 October 13-15, Montréal, Québec, Canada 3/4 Gocevski V., Pietruszczak S. 2004, Numerical simulation of aging mechanisms of heavily reinforced concrete structures, in: Extreme loadings, aging and durability of concrete structures, McGill University Publ., 2004.

3/4 Gocevski V. 2003, Deterioration of a gravity dam due to concrete swelling; Evaluation of present and prediction of future seismic stability, Commission Internationale Des Grands Barrages; Vingt et unieme Congres des Grands Barrages, Q83., R48, Montreal, June 2003.

3/4 Nour A, Cherfaoui A, Gocevski V, Léger P. (2012). CANDU 6 Nuclear Power Plant:

Reactor building floor response spectra considering seismic wave incoherency.

Proceedings of the 15th World Conference on Earthquake Engineering, 24-29 September 2012, Lisboa, Portugal, Paper No 0078.

3/4 Cherfaoui A, Nour A, Gocevski V, Léger P. (2012). Seismic fragility analyses of a CANDU 6 reactor building nuclear power plant. Proceedings of the 15th World Conference on Earthquake Engineering, 24-29 September 2012, Lisboa, Portugal, Paper No 1597.

3/4 Nour A, Gocevski V, Léger P. (2010). Gentilly 2 Nuclear Power Plant: Reactor building floor response spectra using ground motion time histories compatible with median UHS.

Proceedings of the 14th European Conference on Earthquake Engineering. August 30-Septembre 03, 2010 Ohrid, Macedonia, Paper No 1385.

3/4 Tashkov L, Krstevska L, Manova K, Nour A, Gocevski V, Garevski M. (2010).

Evaluation of floor response spectra based on ambient vibration measurements.

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Proceedings of the 9th U.S. National and 10th Canadian Conference on Earthquake Engineering. July 25-29, 2010, Toronto, Ontario, Canada, Paper No 1296.

3/4 Gocevski V., Pietruszczak S., Seismic analysis of damaged hydraulic structures, 2nd Intern. Symp. on Structural Engng., INDIS 2000, Novi Sad 2000, vol.1, pp. 123-132.

3/4 Gocevski V., Pietruszczak S., The effects of slot-cutting in concrete dams affected by alkali-aggregate reaction, 11th Intern. Symp.on Alkali-Aggregate Reaction, Quebec 2000, pp.1303 1312.

3/4 Gocevski V., Remedial work of water intake structure, Symposium 2000, YASE, Vrnjacka Banja,Yugoslavia. November 2000.

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