POSIVA 2012-43

November 2013

POSIVA OY

Olki luoto

FI-27160 EURAJOKI, F INLAND

Phone (02) 8372 31 (nat. ) , (+358-2-) 8372 31 ( int. )

Fax (02) 8372 3809 (nat. ) , (+358-2-) 8372 3809 ( int. )

Matti Hakala

KMS Hakala Oy

Topias Siren, Kimmo Kemppainen

Posiva Oy

Rolf Christ iansson

Svensk Kärnbränslehanter ing AB

Derek Martin

University of Alberta

In Situ Stress Measurementwith the New LVDT-cell –

Method Description and Verification

ISBN 978-951-652-223-7ISSN 1239-3096

Tekijä(t) – Author(s)

Matti Hakala, KMS Hakala Oy Topias Siren & Kimmo Kemppainen, Posiva Oy Rolf Christiansson, SKB Derek Martin, University of Alberta

Toimeksiantaja(t) – Commissioned by

Posiva Oy

Nimeke – Title

IN SITU STRESS MEASUREMENT WITH THE NEW LVDT-CELL – METHOD DESCRIPTION AND VERIFICATION

Tiivistelmä – Abstract

Posiva Oy and SKB (Svensk Kärnbränslehantering AB) tested the suitability a new LVDT-cell (Linear Variable Differential Transducer cell) to measure the induced stresses in the vicinity of an excavated surface and further to use these results to interpret the in situ state of stress. It utilises the overcoring methodology, measuring the radial convergence of four diameters using eight LVDTs, and is similar in concept to the USBM-gauge. A 127 mm diameter pilot-hole is required and the overcore diameter is 200 mm. The minimum overcoring length is 350 mm, and hence a compact drill can be utilised. Extensive testing of the LVDT-cell shows it to be robust and suitable for use in an underground environment. Sensitivity tests also show that the cell can withstand a range of operating conditions and still provide acceptable results. The in situ stress at the measurement location can be solved by numerical inversion using the results of at least three overcoring measurements around the three-dimensional tunnel section. The large dimensions of the measurement tool and the ability to utilise multiple measurements at various locations in a tunnel section, provides flexibility in selecting an appropriate rock mass volume. Because the inversion technique relies on knowing the exact location of the measurements and the geometry profile of the tunnel, modern survey techniques such as Lidar or photogrammetric technology should be used. Checks using traditional surveying techniques should also be used to ensure adequate survey resolution, specially in case of sidecoring measurements. To evaluate the suitability of the LVDT-cell to provide the in situ state of stress, tests were carried out in the drill-and-blast TASS tunnel and TBM tunnel at the Äspö Hard Rock Laboratory in Sweden. The state of stress established using the LVDT-cell was in agreement with the state of stress established previously using traditional overcoring and hydraulic fracturing methods. In this study, the reliability of the numerical inversion solution for the TASS near tunnel profile measurements is lower than the inversion solution for the deeper measurements, which suggests that, in the case of a carefully blast designed drill-and-blast tunnel, the minimum measurement depth for establishing the state of stress should be approximately 500 mm, i.e., conducting the measurements outside the Excavation Damaged Zone. The TBM measurements showed that sidecoring could be used when high tangential stresses are encountered that may cause core discing when employing overcoring. The trials indicated that the field measurements could be efficiently carried out using the current configuration of the LVDT-cell. However, determining the Young’s Modulus and Poisson’s ratio using glued strain gauges caused considerable delays in processing the field measurements. It is recommended that a suitable method be developed to replace gluing strain gauges.

Avainsanat - Keywords

In situ, stress, measurement, rock, overcoring, sidecoring, LVDT, cell.

ISBN

ISBN 978-951-652-223-7

ISSN

ISSN 1239-3096 Sivumäärä – Number of pages

180 Kieli – Language

English

Posiva-raportti – Posiva Report Posiva Oy Olkiluoto FI-27160 EURAJOKI, FINLAND Puh. 02-8372 (31) – Int. Tel. +358 2 8372 (31)

Raportin tunnus – Report code

POSIVA 2012-43

Julkaisuaika – Date

November 2013

Tekijä(t) – Author(s)

Matti Hakala, KMS Hakala Oy Topias Siren & Kimmo Kemppainen, Posiva Oy Rolf Christiansson, SKB Derek Martin, University of Alberta

Toimeksiantaja(t) – Commissioned by

Posiva Oy

Nimeke – Title

IN SITU JÄNNITYSTILAMITTAUS LVDT-KENNOLLA – MENETELMÄN KUVAUS JA VERIFIOINTI

Tiivistelmä – Abstract

Posiva Oy ja SKB (Svensk Kärnbränslehantering AB) testasivat uuden LVDT-kennon soveltuvuutta louhitun pinnan ympärillä vaikuttavan sekundäärisen jännityskentän mittaukseen sekä mittaustulosten soveltuvuutta in situ jännitystilan tulkintaan. Mittaus perustuu irtikairaustekniikkaan USBM-kennon tapaan, jossa neljää halkaislinjan muutosta mitataan kahdeksalla radiaalisella LVDT-anturilla. Kenno asennetaan halkaisijaltaan 127 mm pilottireikään ja irtikairaus tapahtuu 200 mm terällä. Pienin irtikairaussyvyys on 350 mm, joten timanttikairaus voidaan tehdä kannettavalla kalustolla. Kattava testaus osoittaa LVDT-kennon olevan luotettava ja maanalaiseen ympäristöön soveltuva. Käyttöolosuhdetestit osoittavat kennon kestävän vaativiakin olosuhteita ja silti tuottavan hyväksyttäviä tuloksia Tunneliosuudella, jossa tila- ja mittausgeometria on kolmiulotteinen, vallitseva in situ jännityskenttä ratkaistaan vähintään kolmen irtikairausmittaustulokseen perustuen numeerisella inversiolla. Mittakennon suuri halkaisija ja mahdollisuus yhdistää jännitystilatulkinnassa eri rei’istä tehtyjä mittauksia, tuovat joustavuutta mittaukseen soveltuvan kalliomassan valintaan. Toisaalta, koska inversioratkaisu edellyttää tarkan sijaintitiedon mittausrei’istä ja tunneligeometriasta, tulee mittausalueen mallinnuksessa käyttää nykyaikaisia maanmittaustekniikoita kuten Lidaria tai fotogrammetriaa. Erityisesti sivukairaus (sidecoring) mittauksissa riittävä paikannustarkkuus tulee varmistaa perinteisillä mittaustekniikoilla. LVDT-kennomittausten soveltuvuus in situ jännityskentän määrittämiseksi arvioitiin tekemällä mittauksia Äspön kalliolaboratoriossa sekä poraus- ja räjäytysmenetelmällä louhitussa TASS-tunnelissa että täysprofiiliporatussa TBM-tunnelissa. LVDT-kennolla määritetty in situ jännityskenttä oli yhdenmukainen verrattuna aiemmin perinteisillä irtikairaus ja hydraulisen murtamisen menetelmillä määriteltyyn jännityskenttään. Työssä todettiin, että numeerisen inversioratkaisun luotettavuus TASS-tunnelin pinnassa tehdyille mittauksille on alhaisempi kuin syvemmällä tehdyille, mikä viittaa että varovasti panostetun, poraus-räjäytys menetelmällä louhitun, tunnelin minimimittaussyvyyden tulisi olla noin 500 mm, jolloin mittaukset tapahtuvat räjäytyksen vauriovyöhykkeen (EDZ) ulkopuolella. Mittaukset TBM-tunnelissa osoittavat, että sivukairausta voidaan käyttää korkeassa tangentiaalijännityksessä, joka voisi aiheuttaa kairasydännäytteen viipaloitumista normaalin irtikairauksen yhteydessä. Kokeet osoittavat, että kenttämittaukset voidaan suorittaa tehokkaasti nykyisellä LVDT-kennon kokoonpanolla. Kimmokertoimen ja Poissonin luvun määritys käyttäen liimattavia venymäliskoja aiheutti kuitenkin merkittäviä viiveitä kenttämittausten prosessointiin. Täten on suositeltavaa kehittää soveltuva menetelmä korvaamaan venymäliuskojen liimauksen.

Avainsanat - Keywords

In situ, jännitys, mittausmenetelmä, kallio, irtikairaus, viereenkairaus, LVDT, kenno.

ISBN

ISBN 978-951-652-223-7 ISSN

ISSN 1239-3096 Sivumäärä – Number of pages

180 Kieli – Language

Englanti

Posiva-raportti – Posiva Report Posiva Oy Olkiluoto FI-27160 EURAJOKI, FINLAND Puh. 02-8372 (31) – Int. Tel. +358 2 8372 (31)

Raportin tunnus – Report code

POSIVA 2012-43

Julkaisuaika – Date

Marraskuu 2013

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TABLE OF CONTENTS

ABSTRACT

TIIVISTELMÄ

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

2  DESCRIPTION OF THE NEW METHOD ................................................................ 5 

2.1  General background ......................................................................................... 5 

2.2  Sidecoring ........................................................................................................ 9 

2.3  Description of the Probe ................................................................................. 11 

2.4  Description of the biaxial testing ..................................................................... 12 

2.5  Description of the interpretation ..................................................................... 15 

3  PROBE SENSITIVITY TESTS .............................................................................. 19 

3.1  Thermal sensitivity .......................................................................................... 19 

3.2  Mounting system test ..................................................................................... 19 

3.3  Stability test .................................................................................................... 21 

3.4  Boulder test .................................................................................................... 21 

3.5  Overcoring heat sensitivity test ...................................................................... 23 

3.6  Summary of the sensitivity tests ..................................................................... 27 

4  ONKALO SHAFT STRAIN GAUGE AND LVDT MEASUREMENTS VERIFICATION TEST ........................................................................................... 29 

4.1  Site conditions at the -265 level ..................................................................... 29 

4.2  Description of the tests ................................................................................... 34 

4.3  Results ........................................................................................................... 36 

4.4  Summary of measurements in Olkiluoto gneiss ............................................. 41 

5  ÄSPÖ TESTING CAMPAIGN: VERIFICATION IN KNOWN STRESS STATE AND EFFECT OF THE EDZ .................................................................................. 43 

5.1  Site conditions at the -450 level ..................................................................... 43 

5.1.1  Description of the tests ............................................................................ 45 

5.1.2  TASS-tunnel Results ............................................................................... 52 

5.1.3  TBM-tunnel Results ................................................................................. 58 

5.1.4  Summary of Äspö measurements ........................................................... 64 

6  DISCUSSION AND CONCLUSIONS .................................................................... 67 

6.1  Sensitivity test ................................................................................................ 67 

6.2  Selection of test location ................................................................................ 67 

6.3  Biaxial testing ................................................................................................. 69 

6.4  Experiences with operating the probe ............................................................ 69 

6.5  ONKALO verification case .............................................................................. 70 

6.6  Äspö verification case .................................................................................... 70 

6.7  Analyses and uncertainties ............................................................................ 73 

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7  FUTURE WORK .................................................................................................... 75 

REFERENCES ............................................................................................................. 77 

APPENDICES ............................................................................................................... 79 

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1 INTRODUCTION

The nuclear waste industries in Finland and Sweden are in the final stages of designing their spent fuel repositories. The footprint area of the repository will reach several km2. One of the design requirements is to align the deposition tunnels with the orientation of the maximum horizontal stress component in order to reduce the potential for stress-induced damage around these openings. Hence, the final layout of the deposition tunnels will have to be adapted, at least in part, to the orientation of this maximum horizontal stress component. In addition to the orientation, the stress magnitudes are also required for the design of the underground openings. It is well known that traditional stress measurement techniques used in small diameter boreholes, e.g., overcoring and hydraulic fracturing, have limitations. These limitations are related to errors and data scatter in the techniques (Hudson and Cooling 1988). Experimental errors are handled by first identifying them and then improving the instrumentation or testing procedure to either remove them or quantify them so they can be removed analytically. Data scatter on the other hand results from natural variations in the rock mass, in which measurements are to be carried out, e.g., the natural variation of stresses due to local heterogeneities at various scales (Wiles & Kaiser 1990). This data scatter can be difficult to interpret if only a few measurements are carried out. Martin et al. (1990) showed that data scatter decreases as the volume associated with the measurement increases. Hence, a preferred stress measurement technique is one that involves a large rock mass volume and can be easily calibrated and monitored to minimise experimental errors. Stips Oy (Finland) has developed a LVDT-cell (Linear Variable Differential Transducer cell) to measure the stresses associated with the boundary of a tunnel, i.e. close to the surface. When combined with surveying of the tunnel geometry and numerical techniques, these measured stresses can be used to back calculate the in situ stress at measurement location. The in situ stress, as a second-order tensor quantity, is defined by six independent components, so the solution is based on three facts a) having at least six independent measures of stress induced convergences, b) that, in linear elasticity, the displacements or convergences caused by any known alteration in the stress state can be calculated if the geometry is known, and further that c) the in situ stress-induced displacements can be calculated by superimposing displacement components caused by each stress tensor component. The six components of the stress tensor can either be specified as three normal stress components and three shear stress components relative to a fixed set of axes, or as the three principal stress values (major, intermediate and minor) together with their orientations (expressed as the trend and plunge for each). In the work reported here, generally the principal stresses and their orientations will be presented. The motivation for developing the tool was the scale-related problems associated with foliation and heterogeneity of metamorphic rock and the gluing issues encountered with traditional overcoring methods. As a solution for these problems, the measurement volume was increased by using a 127 mm pilot hole and at least 200 mm of overcoring, with deformations being measured with spring-loaded Linear Variable Differential Transformers instead of glued strain gauges. To evaluate the reliability of the LVDT-

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cell, trials were carried out in tunnels at SKB’s Äspö Hard Rock Laboratory, located in South-Eastern Sweden. The state of stress in the trial area had been established earlier using various measurement methods, including traditional overcoring and hydraulic fracturing methods. This report describes the LVDT-cell, the workflow needed to determine the in situ stress, and compares the results from the field trials with the results from the existing traditional measurements.

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2 DESCRIPTION OF THE NEW METHOD

2.1 General background

The LVDT-cell is designed to measure excavation-induced stresses in the immediate vicinity of a tunnel surface. Being a two-dimensional overcoring device, similar to the USBM-gauge or IST-tool, it measures four diametrical convergences caused by full or partial stress release (Figure 2-1) (Hooker et al. 1974; Gray and See 2007). Partial stress release by side coring can be used instead of overcoring in high stress conditions when ring discing is likely to occur (see Chapter 2.2 ). In order to interpret the full 2-D in situ stress tensor, at least three, but preferably four to five measurements are needed. Measurements should be spatially distributed around the profile for the interpretation of the full stress tensor but, at the same time, the potential for core damage potential should also be considered since this is also spatially distributed around the tunnel. The distance between the 200 mm diameter overcored holes should be at least 1 m to prevent stress interference. The diameter of the pilot/installation hole is 127 mm and at least 200 mm for the overcoring. Because only short holes (<1m) are needed, a compact drill rig such as the DIMAS DDM-3HL or Husqvarna DM 406 H can be used (Figure 2-2). The orientation of the LVDTs is controlled manually; normally LVDT-sensor number one is orientated upwards using a small level or to the North or towards the tunnel end when the measurement hole is vertical.

Figure 2-1. The waterproof LVDT-cell version II without battery and memory unit.

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Figure 2-2. LVDT-cell version I in the pilot hole just before overcoring in the Äspö TBM-tunnel. The drill bit is marked after every 50 mm for manual coring advance recording.

The cell installation depth depends on the excavation type and rock conditions. In raise- bored excavations or TBM-tunnels, the excavation damaged zone (EDZ) is minimal and the cell could be installed at a 50 mm depth, which is the practical minimum for the probe (Figure 2-3). In drill-and-blast tunnels, the depth of the EDZ will vary but our experience indicates an installation depth of 500 mm or greater should be beyond the depth of the EDZ for tunnels excavated using modern perimeter-control blasting techniques (see also Chapter 5.1.3). In drill and blast conditions, the coring has to be done in several phases because the normal core barrel length is 500 mm. Further, excavation surfaces should not be shotcreted in order to be able to locate and avoid natural fractures or poor quality rock close to the measurement hole.

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Figure 2-3. Suggested drilling and installation depths in the cases of drill-and-blast and full-face bored excavation.

Calibration readings of the LVDT-cell are obtained before and after the actual measurement (Figure 2-4). This calibration is carried out in a thick-walled aluminium cylinder, having exactly 1 mm difference between the diameters of both ends. The accepted difference in diameters should be between 0.96 and 1.04 mm, assuring linearity better than 4 %. After installation, the stability of the cell is checked at 10 minute intervals and vibrating (shaking) it after that to ensure the LVDTs are secure and stable. Stable convergence values before and after shaking are required in order for overcoring to proceed. After overcoring, the thermal effects from the coring-induced heat and flushing water must stabilise before the cell is recovered, normally one hour is sufficient. During measurement, an installation form is used to record the following items (example in Appendix 2): - measurement location - calibration logs - flushing water and rock temperature - initial cell position - overcoring advance and observations If LVDT measurement is carried out with an online cable connection, which is always preferred, the batteries are charged with an external power unit and the LVDT-reading values and convergence curves can be seen on the screen of the laptop used for data processing (Figure 2-5). The in situ state of stress can be solved by numerical inversion. Detailed and accurate geometrical data related to the tunnel shape, cell installation depth and drillhole locations, orientations and lengths are needed for the numerical three-dimensional model used to carry out the numerical inversion. The geometrical data relating to the

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tunnel and measurement area are best obtained using laser scanning technology or photogrammetry and completed/checked using traditional tunnel survey measurements (see Chapter 2.5). In addition to the geometrical data, the elastic constants of the rock needed for the numerical inversion are determined by testing the pilot cores in a biaxial chamber.

Figure 2-4. LVDT-cell calibration with precision lathed hole in aluminum body.

Figure 2-5. External power supply and rugged laptop for online monitoring and data acquisition.

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2.2 Sidecoring

In high stress conditions, overcoring becomes vulnerable for ring discing (Figure 2-6 and Figure 2-7). To overcome this problem, a partial stress release method termed ‘sidecoring’ was evaluated. This is possible because the applied inverse best fit in situ stress solution scheme does not require total stress release (see Chapter 2.5), i.e, as opposed to analytical solutions, numerical 3D simulations can be undertaken for the geometry before and after sidecoring. The first trial was a 2D numerical study, which was followed by field measurements in the Äspö HRL (Appendix 1). These results suggested a distance between the 127 mm pilot hole and the 200 mm sidecoring hole of between 40 mm and 50 mm. Other distances for the pilot and sidecoring holes are basically the same as for normal overcoring (Figure 2-8). As sidecoring produces the highest convergence and stress release in the direction defined by the centroids of both holes, special care should be paid in planning the locations of the sidecoring hole pair, so that all in situ stress components will be presented. After the measurements are completed, good quality surveying is needed to construct an accurate 3D geometrical model. This is best achieved using photogrammetry or laser scanning combined with manual survey checks. The key parameters that can influence the results are the relative orientation and distances between the two holes (sidecoring hole and the LVDT-cell hole). The measurement accuracy for the distance measurements when using the 50 mm wide sidecoring pillar must be better than 5 mm, for the relative position of the pilot and sidecoring holes. This measurement accuracy will result in less than 10 % interpretation error.

Figure 2-6. Ring discing of 127 mm / 200 mm Äspö diorite cylinder.

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Figure 2-7. Estimate for ring discing potential (left) and the used LVDT- measurement hole types and locations around the drill-and-blast tunnel (right); pilot holes for LVDT-cell (red) and sidecoring holes (green).

Figure 2-8. Pilot and sidecoring hole cylinders fitted to laser scanner points; shortest distances between the holes and corresponding measured values.

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2.3 Description of the Probe

The LVDT-cell described in this report is Version II. The major improvements of Version II are that it is waterproof and the batteries and memory are moved to the front of the cell to minimise the pilot hole length needed for installation. The total length of the LVDT-cell is about 450 mm, but the minimum pilot hole length needed is 350 mm, because the detachable battery and USB-memory unit are located outside the cell’s main body (Figure 2-9). The cell has an internal data logger but also has a parallel working online cable connection. The preferred cabling is connected through the rig drill rods, which requires a T-shape coupling for the flushing hose. Eight LVDT-sensors installed in a single plane measure the convergence of four diameters. The measurement range of the Solartron AXR/1/S analogue LVDT-sensor is ±1 mm and the accuracy is better than 1 m. The cell is mounted in the pilot hole with two expandable o-rings. The mounting is not rigid and it does not centralise the cell exactly, but it is steady and even one o-ring can support the load of the cell and overcored rock cylinder in a vertical hole. In addition to eight LVDT-readings, the temperatures of the cell’s aluminum frame and the pilot hole rock wall are also recorded. Normally, the logger reading interval is about one second and the resulting value is the mean of eight readings. The reading interval can be changed to about 0.2 s, by decreasing the number of readings used for averaging, or increased by using a delay loop.

Figure 2-9. Main components of the waterproof LVDT-cell, Version 2.

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2.4 Description of the biaxial testing

In order to establish the elastic Young's modulus and Poisson’s ratio of the rock, the 120 mm diameter pilot cores are biaxially tested in a biaxial chamber (a Hoek Cell). Two strain gauge rosettes, containing three strain gauges each, are glued in the middle of the sample before testing. The strain gauge rosettes are placed 90 from each other to check for anisotropic behaviour. An illustration of the strain gauges is shown in Figure 2-10.

Figure 2-10. The location and alignment of strain gauges in the 120 mm diameter sample.

The biaxial testing procedure has varied during each of the three measurement campaigns described in this report. The pressure sequence for all the testing, 0 MPa, 2 MPa, 4 MPa, 6 MPa, 10 MPa and 15 MPa, was used for the measurement taken in the year 2011 campaign. In years 2010 and 2012, one higher value of 20 MPa was added. Measurements are taken after holding the cell pressure constant for one minute, with both loading and unloading measurements being recorded. The peak pressure is held constant for three minutes to check for glue creep. The elastic properties are calculated using the unloading values only; however, the full loading cycle is examined to check the linearity of the response. In the first measurement campaign in 2010, the biaxial testing was carried out by Vesa Järvinen (Unisigma Oy) who specialise in strain gauge measurements. Unfortunately, Unisigma Oy was not available for the later campaigns and, in the year 2011 campaign, the biaxial testing was conducted 2–4 days after the measurements in the Äspö laboratory premises by Topias Siren (Posiva) and Daniel Ask (Vattenfall Swedpower Ab). In the year 2012 campaign, the measurements were undertaken afterwards at Posiva's premises by Topias Siren. The pressurisation system, (Enerpac hydraulic hand pump) and 123 mm biaxial chamber (Hoek Cell) with membranes for testing 120 mm samples, was the same in all

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campaigns. The equipment used during the year 2010 campaign is shown in Figure 2-11. The strain gauges were connected in every campaign with a three-wire, quarter-bridge strain gauge circuit.

Figure 2-11. The testing and pressurization system used during all campaigns photographed in 2010 by Vesa Järvinen.

The data recorder used in the 2010 biaxial testing was designed for long term recording of dynamic strain gauge signals. During the campaign in 2011, a Biolab dataTaker DT80 datalogger system was used with manually made bridge completion modules constructed from precision resistors (Figure 2-12) (Biolab 2009). In the year 2012 campaign, commercial bridge completion modules and an additional digital pressure meter were used (Figure 2-13). The year 2011 campaign was the only year when heat was applied to the strain gauge glue (Figure 2-12). The samples were heated in an industrial oven at 40 °C for 5 hours to speed up the glue curing/hardening. In all the other campaigns, the samples were stored in dry conditions for months before testing and no heating was used.

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Figure 2-12. Strain gauge treatment (above) and the dataTaker measurement equipment (below) with self-made bridge completion modules.

Figure 2-13. The testing equipment used in 2012 with commercial bridge completion modules and with additional digital pressure meter.

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2.5 Description of the interpretation

The in situ state of stress is solved by numerical inversion using the LVDT results of at least three optimally placed measurement locations around the excavation profile. Experience suggests that four to six measurements, in well-planned locations, increases the reliability of the results (Figure 2-14). The interpretation assumes a continuous, homogeneous, isotropic and linearly elastic material response (CHILE), but known transverse anisotropy or orthotropy material responses could also be applied. If at the measurement location only direct overcoring measurements are used, a three dimensional element model of the LVDT-cell installation geometry is adequate, because the total stress release results in equal and opposite radial convergence; this is the same assumption as applied for the USBM gauge (Hooker & Bickel 1974). However, when side coring is utilised, a model of the final hole geometry is also needed. The geometrical data for the measurement area are best obtained using laser scanning technology or photogrammetry and completed/checked using traditional tunnel survey measurements. The accuracy of these measurements is normally more than enough for the construction of a 3D-model for numerical simulations. No systematic studies on the effect of model accuracy has been done but some guidelines can be provided. In the case of a drill and blast tunnel, the LVDT-section is 0.5 m depth; based on this and taking an analogue from the Kirsch equation, any irregularity in the surface which is less than 70 mm has a less than 5 % effect on the tangential stress at a measurement location in a very heterogeneous stress field with x/y = 10. The locations and geometry of the LVDT-cell hole and the sidecoring hole must be known more accurately. Laser scanning or photogrammetry normally gives very accurate hole location on tunnel surface, but the true orientation could be more problematic (see Chapter 2.2). To overcome this, a second scanning can be done by using inward pointing strained tubes in all holes. All measurement holes can be included in the same model if the distance between the holes is greater than 1.0 m, which is five times the overcoring hole diameter. If modelling is undertaken with the BEM, 32 elements per circular annulus will provide less than 5 % error for the radial displacement—even in a very heterogenous stress state of x/y = 10. Appendix 4 provides a review of the modelling accuracy.

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Figure 2-14. Three overcoring pilots in tunnel walls, plus three sidecoring pilots and side holes (green) in the roof, and the simplified laser scanned profiles for the numerical model used in the inversion solution.

If the rock mass response satisfies the CHILE material conditions, i.e., no previous observation indicates that heterogeneity, foliation, contacts, fractures, faults or yield has a strong effect on the rock behaviour, a three-dimensional boundary element method is the most efficient numerical modelling tool providing results that are accurate to within 5 % (Figure 2-15). Well known geological features, such as different rock types, foliation, long fractures or lithological boundaries, can be taken into account within the normal limits of numerical 3D-modelling and linear elasticity. In such cases, other numerical approaches must be used. The inversion process requires each model to be run six times. For each run, one of the in situ stress tensor components is set equal to 1 MPa and the other components are set to zero. These unit stress tensor component calculations, give orthogonal displacements components at each LVDT-sensor head. In the case of linear elasticity, the LVDT-sensor head displacements, caused by any in situ stress state, i.e. (k×EE, l×NN, m×UU, n×EN, o×NU, p×UE), can be constructed by superimposing the multiplied displacement components caused by each stress tensor component (Equation 2-1):

ui(kEE, lNN, mUU, nEN, oNU, pUE ) = k×ui(EE=1) + l×ui(NN=1) + m×ui(UU=1) + n×ui(EN=1) + (2-1) o×ui(NU=1) + p×ui(UE=1),

where ui is orthogonal displacement component, i = East, North, Up ii is orthogonal stress tensor component, i = East, North, Up k,l,m,n,o,p are multipliers for stress tensor components and displacement components caused by each stress

tensor component.

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Based on superimposed LVDT-sensor head displacements and their initial locations, it is possible to calculate the convergences between opposite LVDT-sensor heads. The solution for the in situ stress state, i.e., for multipliers (k, l, m, n, o, p), is found iteratively. The accepted solution is the one that provides the best fit between the calculated convergence and the measured values (Appendix 4).

Figure 2-15. Examine3D boundary element method mesh for LVDT measurements in TBM-tunnel (left) and a close-up showing the orthogonal displacement components caused by the unit in situ stress tensor component at LVDT-head 1 location.

If only overcoring measurements are carried out close to the wall of a raise bored shaft or TBM-tunnel, then analytical codes for overcoring can be also used, e.g., codes based on analytical equations for total stress release, as presented by Leeman (1970). The shaft or tunnel is regarded as a large pilot-hole and the measured convergences can be transformed into ‘strain gauge’ strains by first solving the local stress, using equations 1, 2 and 3, presented in Chapter 4.3. For the calculations, elastic parameters are needed. Normally Young’s modulus and Poisson’s ratio obtained from biaxial testing of pilot cores are used (see Chapter 2.4). The validity of the use of intact rock parameters can be questioned; however, the use of parameter values modified for the rock mass can be criticised also because the cell is placed in intact rock. Depending on the amount of test data and the capabilities of the simulation code used, more sophisticated models can be used.

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3 PROBE SENSITIVITY TESTS

The LVDT-cell sensitivity was tested in order to check the possible sources of erroneous readings and to be sure about the functionality. The tests included thermal sensitivity of the cell electronics and response to thermal expansion of the overcored cylinder, the stability of themounting system, reading stability for stable in situ conditions, overcoring measurement in zero stress conditions, and finally more detailed tests of thermal expansion of the rock cylinder during overcoring. These testes are shortly described in this Chapter with more images of the results in Appendix 3.

3.1 Thermal sensitivity

The first version of the LVDT-cell was subjected to two thermal and six mechanical disturbances in order to evaluate the reliability of the readings. In the first thermal test, the cell was installed in the overcored rock cylinder and the outer surface of the rock was heated with a heat gun so that the temperature of the inner surface increased by 11 °C, which is more than normally observed during a measurement (Figure 3-1). The applied heat resulted a rapid 25–35 m expansion when the outer surface of the rock cylinder expands, followed by a 5 m larger contraction when heat is transferred through the cylinder (Figure 3-1). After a one hour cooling period, all convergences were approximately 5 m less than before heating. The total temperature recovery takes over 10 hours, but the changes in measured diameters are small after two to three hours. In the second thermal test, the LVDT-section of the cell was inside an overcored rock cylinder and only the aluminum body, including the cell electronics, was heated so that the temperature increased by 10 °C degrees. This resulted in a maximum 15 m change in measured diameters. Again, the cooling period was not long enough to monitor reversal (Figures A4.3 and A4.4).

3.2 Mounting system test

In the mechanical tests, the reliability of the o-ring mounting system was studied. In these tests, the cell was installed either inside the overcored rock or inside a steel cylinder. First, the cell was shaken from a mounting screw by hand, then the cylinders with the cell were rolled on the floor with high and low speed; finally, the cylinder with the cell was set on a table and the table was impacted with a rubber hammer (Figures A4.5 to A3-16). All these tests showed that the o-ring mounting is not rigid, allowing an oscillation close to the mounting position, but none of the tests resulted in more than over 10 m permanent change in the measured convergence. This type of mechanical disturbance is normally seen as a peak or immediate jump in the measurements and can be removed from the final results.

20

Figure 3-1. Set-up for LVDT-cell heating test 1.

Figure 3-2. Heating test 1, diametric deformations when the rock cylinder around the LVDT-cell is heated.

21

3.3 Stability test

The effect of drilling-induced thermal expansion was evaluated during field trials in the Äspö TASS tunnel with the new cell (Version II). After overcoring, the cell was left in the pilot hole for 12 hours in order to allow the temperature to cool down to its initial value. At the same time, the stability of the cell readings was monitored. The cell proved to be very stable for the 12-hour monitoring period and the maximum thermally induced convergence was less than 4m (Figure 3-3).

Figure 3-3. Twelve hour stability monitoring of LVDT-cell in the Äspö TASS R1-deep measurement.

3.4 Boulder test

A LVDT overcoring measurement was carried out in a rock boulder free from external loads. (Figure 3-4). The assumed result is close to a zero stress state since Martino et al. 1997 showed at the URL in Canada that in similar hard rock the residual stresses of rock are minor (< 1 MPa). Before the test, the boulder was stored in ONKALO Investigation Niche 1 at the -140 m level for two years in very stable conditions. This test resulted in convergences ranging from 8–11m, which corresponds to about 3 MPa hydrostatic stress in the measurement plane (Figure 3-5). During the tests, the cell temperature increased by 2 °C. The sensor for rock temperature was not working, but normally it is few degrees higher than measured in the probe. The data recording was interrupted before the rock had totally cooled down.

22

Before this test, the Borre probe (Sjöberg & Klasson, 2003) and CSIRO-HI (Worotnicki 1993.) cells were also tested in this same boulder. Both devices resulted in principal stresses between 2 MPa compression and 3.5 MPa tension. The most probable explanation for these results is thermal expansion of the overcored rock cylinder caused by drilling-induced heat and flushing water.

Figure 3-4. Boulder used for several in situ stress measurement tests.

Figure 3-5. LVDT-cell overcoring results in a boulder free from external loads.

23

3.5 Overcoring heat sensitivity test

Based on the experiences from the boulder test (Section 3.4) the drilling induced heat was found to be a potential source of error for the LVDT measurements. Key questions related to the magnitude, duration and correlation with rock type. The test campaign was performed in the ONKALO POSE investigation niche. The campaign consisted of seven overcoring experiments, where temperatures were measured in the rock during overcoring with 28 temperature sensors (Rantanen 2013). The heat profile of the pilot hole wall was measured with a device consisting of 20 digital band-gap sensors, distributed evenly both in the axial and radial directions (Figure 3-6). In addition to the pilot hole wall temperature measurements, eight miniature NTC-thermistors with fast thermal response were installed in 14 mm drillholes, located so that the overcoring drill bit would intersect the sensors at 200 mm depth (Figure 3-7). The idea was to measure the temperature immediately in front of the advancing drill bit. The sampling rate of the digital sensors was 1/3 Hz. The analogue sensors had a sampling rate of 2 Hz in the first two measurements, and 20 Hz in the last measurement.

Figure 3-6. Instrumentation system to monitor overcoring induced heat in conditions similar to the LVDT measurement.

24

Figure 3-7. Pilot hole temperature monitoring tool, two inclined drill bit temperature measurement probes and the drill rig.

Measurements were made in two different rock types, veined gneiss (VGN) and pegmatite granite (PGR). In previous LVDT measurements, it had been established that these two rock types were quite different in terms of drill bit penetration rate. PGR is much harder to core drill and was thus expected to produce considerably more heat during overcoring. The rock types also have slightly different thermal properties. Pegmatite has a 22 % higher diffusivity and a 34 % higher thermal expansion coefficient; other differences are less than 12 %. Two other factors were also included in the testing campaign. The first was to test how the residual heat generated during pilot-hole coring affected actual overcoring temperatures. The second was the overcoring depth, because LVDT measurements are generally made at two different depths, depending on the excavation method. The following tests were carried out (hole ID in parentheses):

1. Overcoring in VNG at 50 mm depth, 12 h cooling time after pilot-hole coring (ONK-SH125)

2. Overcoring in VNG at 50 mm depth, immediately after pilot-hole coring (ONK-SH124)

3. Overcoring in VNG at 500 mm depth (ONK-SH124) – Test failed 4. Overcoring in PGR at 50 mm depth (ONK-SH126) 5. Overcoring in PGR at 300 mm depth (ONK-SH126) 6. Overcoring in VGN at 50 mm depth (ONK-SH138) – Test failed 7. Overcoring in VGN at 500 mm depth (ONK-SH138) – Test terminated

25

The most significant observations from the testing campaign are as follows: Temperature change in good overcoring conditions is not very high. The lower

temperature occurs in the softer rock types, which are faster to core drill. Maximum ΔT on pilot hole wall varied from 3.7 °C (test no. 1, VGN) to 14.5 °C (test no. 5, PGR).

The maximum temperature reached in the pilot hole increases with depth. This is most likely due to both gradual heating of the drill bit and longer flushing water circulation as the drill bit advances deeper. It is also worth noting that a temperature plot ΔT(t) for any single set in Figure 3-8 can be modelled as a log-normal function.

The temperature increase in the rock cylinder is radially symmetric. The thermal expansion in the cylinder is therefore axially symmetric. In a mathematical sense, this also reduces the overcoring induced heat in the pilot hole wall to three dimensions: depth, temperature and time.

Flushing water temperature and circulation is a major factor. It was found that differences in flushing water pressure and circulation had quite a large and instantaneous effect on the measured temperatures.

The cooling process is quite slow. In these tests, a temperature of 1 °C above ambient was not reached until about 6 hours after overcoring.

The rock temperature on the outside of the drill bit seems to be higher than inside, Figure 3-9. As the cooling flushing water is injected from inside the drill bit, it provides considerably more cooling to the rock cylinder inside the drill bit than to the rock mass around the drill bit. This might yield an interesting asymmetric temperature distribution during sidecoring measurements.

Numerical simulations based on measured temperature distributions were done to estimate the radial displacement at the location of the LVDT-sensors (Rantanen 2013). Due to the axially symmetrical temperature field, a 2D axi-symmetrical model was used for the simulations. Results from the simulations suggest that a 3° C increase in temperature at the location of the LVDT-sensors (as in test no.2) results in a maximum of about 1.5 µm radial expansion. This result is for fast overcoring in the veined gneiss. The maximum temperature increase was measured in tests 4 and 5, when overcoring was performed in the pegmatitic granite. The high amount of quartz in this rock type makes it hard to core drill, thus creating more heat and slowing of the drilling rate. According to the numerical simulations, a 10 °C temperature increase measured in tests 4 and 5 results in a maximum of 8 µm radial expansion in the rock cylinder at the location of the LVDT-sensors. These displacements apply only to the maximum temperature.

Based on the temperature measurements and the numerical simulations, it is suggested to leave the LVDT-cell to measure displacements for at least four hours after the end of overcoring, and preferably for six hours. This will ensure that temperature effects do not affect the measurements.

It should be noted though that none of the various axially variable heating power profiles used in the simulations were able to completely mimic the measured behaviour. Simulations also involve other error sources and suffer from certain simplifications. It is thus advised to perform complementary laboratory tests to obtain an actual response, as measured by the LVDT-cell itself.

26

Figure 3-8. Mean axial temperatures along the pilot hole wall versus time in test no. 2, overcoring took place during the first 24 minutes.

Figure 3-9. Temperatures close to the drill bit in test no. 2 for the first 40 minutes.

27

3.6 Summary of the sensitivity tests

All the field trials and sensitivity tests indicated that the LVDT-cell is a robust tool that should provide reliable and consistent results in most field operating conditions. However, as with all field tools containing electronic and sensitive equipment, reasonable care is required. In particular, the temperature changes of the rock and the cell have clear effects on the results and should be minimised. This requires controlled flushing water temperature, suitable drill bits and adequate time for the cell to adjust to any temperature changes. If adequate time after a measurement is not provided, this can produce about a +10 µm error in the measured convergences—which corresponds to about +3 MPa error for all stress components in the measuring plane, assuming a deformation modulus and thermal parameters typical for the Olkiluoto gneiss.

28

29

4 ONKALO SHAFT STRAIN GAUGE AND LVDT MEASUREMENTS VERIFICATION TEST

4.1 Site conditions at the -265 level

The ONKALO is a research facility and in the future it will be a part of the nuclear waste repository. The ONKALO is located within the Olkiluoto Island in Satakunta, on the south-western coast of Finland. LVDT measurements were performed in several phases in the ONKALO access tunnel and a shaft. The shaft measurements were undertaken at several depth levels, with the -265 m level being studied in more detail. Generally, Olkiluoto is located in the middle of the 1.9–1.8 Ga Palaeoproterozoic bedrock, which consists of pelitic (in the South) and psammitic (in the North) migmatites. In the middle of the area, there are Mesoproterozoic rapakivi granites (1.58–1.48 Ga), sandstone (1.4–1.3 Ga) and olivine diabases (1.27–1.25 Ga). The bedrock at Olkiluoto (Figure 4-1) is mainly described as variations of migmatites (supracrustal high-grade metamorphic rocks) which can be divided into four major classes by the mineral composition, texture and migmatite structure: 1) migmatitic gneisses; 2) tonalitic-granodioritic-granitic gneisses (TGG gneisses); 3) other gneisses occurring as inclusions in the migmatitic gneisses, including mica gneisses, quartz gneisses and mafic gneisses; and 4) pegmatitic granites (Kärki & Paulamäki 2006). The migmatitic gneisses can be further sub-divided into several types on the basis of their migmatite structures and degree of migmatisation. All lithologies have gone through a polyphase ductile deformation during the tectonic evolution; five of these phases can be found. The geology is described in detail in the Olkiluoto Site Decription (Posiva 2013). The fracturing in the rock at Olkiluoto is rather regularly distributed as three sets. One set is gently or moderately dipping to the south-southeast, which is parallel to the ductile foliation (Posiva 2012). Two sets are steeply (sub-vertical) dipping fracture sets, striking to the north-south and east-west. The fractures are filled with calcite, pyrite, kaolinite, chlorite and illite. The first two infillings are more common in the surface conditions. In most of the cases, the fracture profile is irregular or semi-rough and undulating (Posiva 2012). The fracture system includes also hydrothermal veins, which are mainly planar or undulating. The vein infillings are spherulitic iron-rich chlorite, sphalerite, galena and pyrite. Those can be detected visually. Defining the rock mass quality is done in several phases, using Q (Quality) and GSI (Geological Strength Index) classifications. According to the tunnel mapping data, the rock mass quality in ONKALO is better than good (GSI-classification) below the brittle deformation zone (OL-BFZ019) (Figure 4-2) (Posiva 2013).

30

Figure 4-1. Lithological map of Olkiluoto.

Figure 4-2. The rock mass quality (Geological Strength Index, GSI) in the ONKALO tunnel (Posiva 2013). The LVDT measurement location is indicated by the black arrow.

z=+10m

z=‐437m

‐265 m level measurement location 

31

LVDT measurements were performed at the -265 m level in the exhaust air shaft. The lithology varies at the location, and the rocks can be described as migmatitic gneisses. The grain size varies from a few millimetres to some centimetres (Figure 4-3). The amount of leucosome varies from 30 to 90 percent. From a geological point of view, the rock does not fulfil the CHILE (Continuous, Homogeneous, Isotropic, Linearly Elastic) assumptions, but a final decision is made after the biaxial testing.

Figure 4-3. Core box image from drillhole OL-KR38 (Rautio, 2005).

During the summer of 2005, Posiva diamond-drilled the shaft pilot hole OL-KR38 (Rautio 2005). The drilling was done by using a NQ3- triple tube core barrel, which produces 50.5 mm core samples. The deviation of the drillhole was measured frequently, to ensure the drillhole remained within 3 m of the shaft perimeter. The drilling depth at the LVDT measurement location is around 275 m. The rock type was

32

described as fine or medium grained migmatite (diatexitic gneiss), where it occurs as pegmatite stripes (Figure 4-4). Some fine grained gneiss inclusions were also found. The fracture frequency at that depth is sparse with only random fractures. The rock quality designation (RQD) is 100 % at this depth level. The foliation direction and type varies, which is the norm at Olkiluoto. The geological ductile deformation logging described the foliation as irregular (Milnes et al. 2006) around the depth of the LVDT measurement location. At drilling depths 262–263 m and 281–282 m, the foliation is moderately banded and dipping gently to the south-east to east direction. The LVDT-pilot holes were drilled from 265 m depth in the exhaust air shaft walls. The drilling depth was selected after visual inspection in the shaft, when the basic rule was to avoid natural fractures if possible. Also, all drilling locations avoided the thicker pegmatitic veins and fine grained gneiss inclusion. All pilot holes were drilled in diatexitic gneiss (Figure 4-8), where grain size and the amount of leucosome varied. Pilot hole samples from the hole ONK-SH68 are not available. Samples ONK-SH69, 71, 72 and 73 are coarse grained diatexitic gneisses, where leucosome can be seen in several locations as like veins, contacts are diffusive (not sharp). The sample from the hole ONK-SH70 is more medium or fine grained than that previously mentioned. This hole was drilled partly in fine grained mica or quartz gneiss (from Figure 4-4 to Figure 4-8).

Figure 4-4. The pilot core samples from drillhole ONK-SH69 (R2), the sawing is done from 5 cm depth, i.e. at the plane of the LVDT-gauges.

Figure 4-5. The pilot core samples from drillhole ONK-SH70 (R3), the sawing is done from 5 cm depth, i.e. at the plane of the LVDT-gauges.

33

Figure 4-6. The pilot core samples from drillhole ONK-SH71 (R4), the sawing is done from 5 cm depth, i.e. at the plane of the LVDT-gauges.

Figure 4-7. The pilot core samples from drillhole ONK-SH72 (R5), the sawing is done from 5 cm depth, i.e. at the plane of the LVDT-gauges.

Figure 4-8. The pilot core samples from drillhole ONK-SH73 (R6). The sawing is done from 5 cm depth, i.e. at the plane of the LVDT-gauges.

During the overcoring, ring discing occurred at two locations (samples R1 and R6 Figure 4-9 ONK-SH68 and ONK-SH73). Ring discing caused steps in the diametrical deformation results and finally interrupted the drilling when the sample broke inside the core barrel. The main rock type for these samples was diatexitic gneiss and the leucosome was coarse grained. It appears that the discing starts from the leucosome rich areas. The disc thickness varies between 20 and 40 mm; the reason for this is a combination of lithology, tensile strength of the rock and stress magnitude.

34

Figure 4-9. Overcoring samples from drillhole ONK-SH68 (R1) and ONK-SH73 (R6). Pervasive ring discing in both samples, with ring thickness varying between 20 to 40 mm.

4.2 Description of the tests

As a first in situ stress measurement case, the LVDT-cell was verified against long strain gauge measurements at the -265 m level in the ONKALO exhaust air shaft (KU2). Both measurements were made in the same six locations around the shaft perimeter (Figure 4-10). Firstly, the strain gauge rosettes were installed on the diamond-ground shaft wall surface (Figure 4-11). Each rosette consists of three 50 mm strain gauges: tangential (horizontal), 45 inclined, and vertical. The rosettes were released from the shaft surface secondary stress state by 127 mm diamond coring. All the strain gauges were monitored during the drilling. The LVDT-cell was installed in this 127 mm pilot hole so that the LVDT-section was 50 mm from the shaft wall (Figure 4-12). The LVDT-cell was released from the ‘double’ secondary stress state caused by the shaft and pilot hole geometry by 200 m diameter overcoring. The length of the overcoring was 250 mm.

35

Figure 4-10. Location of strain gauge and LVDT-cell measurements at the -265 m level in the ONKALO exhaust air shaft.

Figure 4-11. Glued strain gauges on the ONKALO exhaust air shaft hole wall (upper left); release of the strain gauge rosette by 127 mm diamond drilling (upper right); and biaxial testing (lower left).

36

Figure 4-12. LVDT-cell installed in the pilot hole before overcoring (left); released cylinder after overcoring (right).

The Young’s modulus was defined from the unloading part of the biaxial test on 120 mm pilot cores assuming CHILE plane stress conditions. All three existing strain gauges at the core end were used. Because the end of the core with the strain gauges has to be outside the biaxial cell loading rubber, this test set-up will result in errors ranging from 0 % to +43 % depending on the unloaded length of 0 to 5 mm. This error estimate is based on an axisymmetric finite element Phase2 simulation. The Poisson’s ratio cannot be interpreted from this strain gauge configuration and 0.25 was assumed. The comparison of the interpreted values with the ONKALO area uniaxial test results from the same depth level suggest that the error is about -10 %, although a positive error was expected. Therefore, the assumed Poisson’s ratio is too high. The error in Young’s modulus and the missing Poisson’s ratio value does not affect these verification tests because the same values are used for both methods.

4.3 Results

The detailed measurement data for the strain gauge measurements and biaxial testing are presented in Appendix 5. The LVDT measurement data are in Appendix 6. The biaxial test results showed a linear unloading, R2 = 0.99–1.0, but moderate variation between the samples exist (Figure 4-13). The mean Young’s modulus value is 50.5 GPa and variation between individual rosettes is from 39.4 GPa to 64.9 GPa (Table 4-1).

Although there is no clear foliation at the measurement level, nor in the individual samples, a possible correlation between strain gauge orientation and the large scale foliation orientation (dip 35°, dip direction 105°) was tested (Figure 4-15). This test supported the lack of consistent deformation anisotropy because the assumed correlation was observed only in the case of sample R6, i.e., that the modulus would have the lowest value perpendicular to the foliation (Hakala et al. 2005). Based on this, it can be concluded that the variation is mainly caused by heterogeneity of the rock.

37

Table 4-1. ONKALO exhaust air shaft at level -265 m, Young’s modulus from biaxial test unloading. Values for individual gauges and the mean.

Measurement Tangential Inclined Axial Mean   [GPa]  [GPa] [GPa] [GPa] 

R1  72.6  62.1 59.8 64.9 R2  48.1  49.4 47.5 48.4 R3  52.8  51.7 52.0 52.2 R4  51.8  51.9 56.9 53.5 R5  42.4  46.2 44.9 44.5 R6  49.0  37.5 31.8 39.4 

Figure 4-13. ONKALO exhaust air shaft at level -265m, biaxial test unloading of all tangential (upper left), axial (upper right) and inclined (lower left) strain gauges.

For the strain gauge and LVDT measurements, two different principal stress interpretations were made, based on the final stable strain or convergence values. The first solution was for the measurement location of the principal stresses in the shaft wall, assuming that plane strain conditions and the local mean modulus values (Equations 4-1 and 4-2 and Figure 4-14). For the LVDT convergence measurements, a numerical best

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38

fit solution for the Kirsch equations was applied (Equation 4-3). The second solution is for the in situ state of stress at the measurement level. This solution was carried out using the overcoring stress measurement interpretation code CSIRA (Amadei 2000), assuming the shaft as a large vertical pilot hole. For both methods, linear elasticity and the measurement level mean modulus were assumed. The strain gauge results can be applied directly but, for the LVDT measurement, the first interpreted local shaft wall stress state was transformed into local strains in tangential-, axial-, and 45° inclined directions. This transformation was done using the local modulus.

 

  

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(4‐3)

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Figure 4-14. Naming of strain gauges and definition of principal axis.

39

Figure 4-15. Measured Young’s modulus for each strain gauge and the angle between gauge vector and mean foliation plane (red plane through pilot holes dip 35°, dd 105°)

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Local principal stresses For strain gauge measurements, the solution is always exact because there are three unknowns and three measurement values; on the other hand, no information on the reliability of the local solution is achieved. In the case of the LVDT measurement, there are four measurements allowing error calculation. The cumulative error compared to the sum of measured convergences is less than 2 % for measurements R1 to R5 and 6 % for measurement R6, indicating good quality measurement (Appendix 4). There was clear ring discing observations on measurements R1 and R6. The comparison of the local principal stress solution shows excellent agreement for measurements R1 and R2, good for R5 and R6, and moderate for R3 and R4 (Table 4-2). The orientations of the local principal stresses are in excellent or good agreement. Further, the local magnitudes and orientations around the shaft are reasonable for slightly a W–SW dipping in situ stress field.

Table 4-2. Strain gauge and LVDT-cell measurement solutions for local principal stresses in the shaft wall.

Location  Method  Sigma_1  Sigma_3  Rot_sigma1 * 

    [MPa]  [MPa]  [°] 

R1  SG  56.1  30.8  ‐11   LVDT  56.0  31.0  ‐13 

         

R2  SG  33.1  14.4  ‐7   LVDT  33.5  13.5  ‐4 

         

R3  SG  26.4  20.0  ‐5   LVDT  24.5  12.5  ‐9 

         

R4  SG  34.4  22.0  +12   LVDT  30.5  15.5  +13 

         

R5  SG  25.3  13.1  +10   LVDT  29.5  13.0  +10 

         

R6  SG  36.7  19.0  +10   LVDT  39.5  20.5  +13 

*) clock wise from horizontal plane when looking towards shaft wall 

 

In situ stresses For the CSIRA solution of in situ stress, the local stresses of the LVDT measurement have to be transformed into strains. For this transformation, local and average modulus values were tested (Table 4-3). Also, one solution without ring discing measurements R1 and R2 were tested. For the strain gauge measurement, only one solution was made.

41

The resulting principal stress magnitudes and bearings are in good agreement, but the dip of maximum compression is 18° higher for the strain gauge measurements (Table 4-3, Figure 4-16). The intermediate and minor principal stresses have almost equal magnitudes which make the orientations sensitive. The maximum compression is SW–S, which agrees with the ring discing observations in the LVDT measurements R1 and R6. This is also consistent with the local stress solutions.

Figure 4-16. In situ principal stress orientations and magnitudes based on strain gauge and LVDT measurements at ONKALO exhaust air shaft level -265 m. Strain gauge orientations have black indicators.

4.4 Summary of measurements in Olkiluoto gneiss

The result of this first trial was encouraging. The magnitudes of the in situ principal stresses are close to the strain gauge measurement and the internal errors for both methods are about the same. The major difference is in the plunge of the maximum principal stress, but which one is correct is unknown. The strain gauge solution is more sensitive with only one inclined strain gauge at each measurement. The comparison of the intermediate and minor principal stress orientation is meaningless because the magnitudes are equal.

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43

5 ÄSPÖ TESTING CAMPAIGN: VERIFICATION IN KNOWN STRESS STATE AND EFFECT OF THE EDZ

5.1 Site conditions at the -450 level

The Äspö Hard Rock Laboratory (Äspö HRL) is located in south-eastern Sweden. The underground laboratory was constructed to create opportunities for research, development and demonstration for geological disposal of nuclear waste in an in situ environment. The bedrock is dominated by intrusive rocks with a quartz monzodioritic, granodioritic or granitic composition, i.e., with a variable content of quartz, which belong to the 1.86–1.65 Ga so-called Transscandinavian Igneous Belt (TIB), Kornfält et al. (1997). Different rock types have also been distinguished on the basis of their variable grain size and texture, including both porphyritic and even-grained rocks. The dominant rock type at Äspö is the porphyritic so-called Äspö diorite that varies in composition between quartz monzodiorite and granodiorite. The bedrock in the Äspö region is dominated by well-preserved intrusive rocks. However, a faint to weak foliation is present. In many cases, it is difficult to decide whether this foliation represents a flow foliation formed during the igneous evolution, or whether it is a later overprinting structure. The gross structural framework was formed when the bedrock still responded to deformation in the ductile regime and discrete, low-temperature, brittle-ductile to ductile shear zones form the most prominent ductile structures in the region. The majority of the deformation zones in the region are dominated and characterised by polyphase brittle reactivation. In general, the patterns of fracture orientations and relative fracture intensities in the bedrock are consistent with the orientations of the regional deformation zones. The most dominant fracture set is steeply dipping and trending NW–SE, see Figure 5-1. This orientation is also close to the orientation of the major horizontal stress.

Figure 5-1. Poles to all water-conducting features (WCF) in the Äspö HRL /Mazurek et al., 1996/. These fractures are frequently traceable across the full tunnel section.

44

Stress measurements have been carried out in different campaigns and with different methods at the Äspö HRL. The first compilation was for the ZEDEX (Zone of Excavation Disturbance EXperiment) project (Emsley et al. 1997). The results from previous stress measurements included two surface boreholes together with six tunnel boreholes drilled from the tunnels. Measurements were performed using hydraulic fracturing and overcoring (the latter with the Borre Probe and CSIRO Cell). The stress measurements were carried out from the 300 m level down to the 600 m level with a concentration around 420 m to 480 m. The results from the stress measurements were estimated to give a well assembled picture for the vertical and the minimum horizontal stress, with a tendency for higher stresses and larger dispersion with depth. However, the maximum stress was found to have a large dispersion and no clear depth dependence. Based on these early measurements, the expected vertical stress at the level 450 m is 14–21 MPa, maximum horizontal stress between 16 and 26 MPa, together with a minimum stress between 9 and 14 MPa. The orientation of the maximum horizontal stress was expected to be around 120–150° (local Äspö96 co-ordinate system). Christiansson and Jansson (2003) reported stress measurements in two orthogonal boreholes at -450 m using the AECL Deep Doorstopper (DDGS), the Borre triaxial overcoring cell and hydraulic fracturing (Figure 5-2). The triaxial measurements with the Borre probe were the only method that provided an estimate of the principal stress magnitudes and orientations, see Table 5-1 and Figure 5-2. Orientations are given in the local Äspö96 co-ordinate system. Table 5-1. Summary of Borre probe results from borehole KF0093A01. (Äspö96)

Measuring point No.

Borehole depth (m)

σ1 [MPa]

σ1 Trend/pl.

σ2 [MPa]

σ2 Trend/pl.

σ3 [MPa]

σ3 Trend/pl.

1 32.14 32.5 307/38 13.8 096/48 8.7 204/16 2 32.70 36.0 310/38 17.7 114/51 8.9 214/08 4 35.38 23.2 308/10 14.2 044/30 6.9 204/58

460 m level average

29.8 310/31 14.8 088/52 9.4 206/21

Figure 5-2. Principal stress directions for test points in borehole KF0093A01 and lower hemisphere schematic plot. North refers to local north (Äspö96).

45

Christiansson and Jansson (2003) concluded the following stress state for the target volume at about –455 m. The maximum horizontal stress is 24 ±5 MPa, (most likely within the upper range)

trending 136° (in RT90, 124° in local Äspö96 co-ordinate system). The minimum horizontal stress is between 10 and 13 MPa, which is similar to the

magnitude of the vertical stress. The measured vertical stress ranged between 15 and 20 MPa, while the theoretical

vertical stress is 11.8 MPa assuming a gradient of 0.026 MPa/m. The scattering is attributed to localised nearby vertical fracturing.

Andersson et al. (2004) conducted back analyses of tunnel deformation measurements approximtely 50 m from the location of the orthogonal boreholes (tunnel behind the shaft in Figure 5-3). The tunnel was excavated perpendicular to the major horizontal stress according to the measurement results from the two orthogonal holes. The analyses assumed plane strain (2-D) conditions and the tunnel is perpendicular to the major horizontal stress. He found the maximum horizontal stress to be 27 MPa, trending at 310° and plunging at 7°. Ask (2005) re-evaluated the Äspö HRL database and concluded two stress regimes at the 420–450 m level. The minor deformation zone NE2 was used to explain the difference in the state of stress between the ZEDEX area at the 420 m level (near borehole KK0045G01 in Figure 5-3) and the results from the 450 m level. Nonetheless, it is believed that the results from the two orthogonal boreholes provide the most reliable stress model for the 450 m level and these data will be used for comparison with the results of the measurements using the LVDT probe.

5.1.1 Description of the tests

TASS-Tunnel LVDT measurements in the TASS-tunnel (T = tunnel, AS = ÄSPÖ, S = tunnel) were carried out in August 2011 between chainages 60 m and 65 m (at the -450m level). TASS is a 4.5 m by 5.5 m horse-shoe-shaped tunnel excavated carefully with the drill and blast method. The tunnel bearing is 218° and plunge is 0.6° upwards (Figure 5-3). Because the tunnel was excavated approximately perpendicular to the major horizontal stress,s two overcoring measurements were made on both walls and three side corings in the roof (Figure 5-4). Hole locations were designed to measure stresses both in the vertical and horizontal planes and to avoid natural fractures. Overcoring holes R1 and R6 measured the in situ stress component along the tunnel axis and the excavation disturbed vertical stress component. The side coring holes R2, R3 and R4 measured the excavation disturbed horizontal stresses in the roof and shoulders. LVDT measurements were planned to study the effect of the excavation damaged zone (EDZ); therefore, two sets of measurements were made: the first at 10–15 cm depth, probably in the excavation damaged zone (EDZ); and the second at 460–660 mm depth in undamaged rock. To define the elastic parameters, all undamaged pilot hole cores were tested in the biaxial cell.

46

Figure 5-3. View of the Äspö HRL from the SSW. The location for the test of different measurement methods in two orthogonal boreholes is north of the F tunnel and the location of the TASS-tunnel (S-tunnel in the Figure) is at the -450 level.

Figure 5-4. Pre-selected locations of LVDT overcoring (yellow) and side coring (orange) hole locations, ends of the drill-and-blast round with visible excavation damage are marked with gray shading (left). The surveyed hole locations on the TASS-tunnel profile and depths for overcorings, sidecorings and LVDT-sections (right).

47

TBM-Tunnel LVDT measurements in the TBM tunnel were done in three phases. All measurements are located close to chainage 3330 m at the -440 m level (Figure 5-3). The tunnel bearing is 248° and plunge 8° downwards. The tunnel was excavated using a full face tunnel boring machine (Figure 5-7). Seven overcoring measurements were first attempted in October 2010 (Figure 5-7). Three measurements were successful; in two attempts, ring discing occurred; in one hole the unsealed LVDT-cell version I did not function properly and in one hole the presence of a natural fracture prevented overcoring. The ring discing occurred immediately when the dip of the radially aligned measurement hole was over 45° from the sub-horizontal tunnel centreline plane (Figure 5-8). In two other holes, dip -13° and 27°, incipient ring discing was observed 25 minutes after overcoring. The second campaign was done in August 2011 using the LVDT Version II and coincident with the TASS tunnel measurements. These measurements were intended to complete the first campaign using sidecoring holes in the roof in the region having the highest tangential stresses (Figure 5-9). In this second campaign, four successful side coring measurements were completed. All these measurements were carried out close to the TBM tunnel-surface in the depth range of 60 cm to 240 mm. In the third campaign, four deeper measurements were done by continuing the previous holes (Figure 5-10); two of those were overcoring and two sidecoring. The depths for these measurements were between 350 mm and 780 mm.

Figure 5-5. The Äspö TBM-tunnel with the overall measurement locations indicated.

1

2

3

48

Figure 5-6. The measurement locations in the Äspö TBM-tunnel (see Figure 5-7).

Figure 5-7. Exact locations of the seven first LVDT measurement attempts in the Äspö TBM-tunnel.

1

2

3

49

Figure 5-8. Well developed immediate ring discing caused by high stresses perpendicular to the core axis in overcored cylinder R6 (left) and incipient ring discing in R7 observed after overcoring (right).

Figure 5-9. TBM-tunnel LVDT measurement holes used in the first and second campaign. Depths for overcorings and sidecorings and LVDT-sections. Pilot holes are red, sidecoring holes yellow, interrupted overcoring attempts green and unused pilots with a dashed line.

50

Figure 5-10. TBM-tunnel LVDT measurement holes used in the third campaign. Installation depths for the LVDT-section.

Biaxial testing During the campaigns in 2010, 2011 and 2012, the biaxial samples in Table 5-2 were tested. The individual stress-strain Figures and other information are presented in the Appendices for each individual sample. The samples are from various depths measured from the tunnel surface and judged to belong to the following three domains, which are used in the Figures and text: (1) surface (0–400 mm), (2) semi-surface (400–700 mm) or (3) deep (>700 mm). In the TBM tunnel tests, because of the expected shallow depth of the EDZ, all biaxial tests were expected to be of a high quality. In the TASS tunnel, the EDZ was expected to impact the tests in the surface domain. The quality of the testing was judged using three categories: (1) high confidence, (2) low confidence, and (3) failed. Measurements where the strain gauge readings were stable during the testing, show linear response and return to zero after the measurement are reliable and therefore judged as high quality testing. An example of high quality testing results from the biaxial sample from hole 1 in the TBM-tunnel (ÄSPÖ-TBM-PL3335 R1) at the depth of 835 mm from the tunnel wall is shown in Figure 5-11.

51

Table 5-2. Tested biaxial samples in the 2010, 2011 and 2012 campaigns.

Work name  Hole ID  Year  Sample type  Testing depth  Quality 

ÄSPÖ‐TBM‐PL3335 R1  2012  deep  830 mm  High 

ÄSPÖ‐TBM‐PL3335 R2  2012  semi‐surface  690 mm  Low 

ÄSPÖ‐TBM‐PL3335 R2  2012  deep  920 mm  High 

ÄSPÖ‐TBM‐PL3334 R3  2012  deep  850 mm  Failed 

ÄSPÖ‐TBM‐PL3334 R3B  2012  surface  130 mm  Failed 

ÄSPÖ‐TBM‐PL3334 R3B  2012  semi‐surface  410 mm  Low 

ÄSPÖ‐TBM‐PL3336 R9C  2012  surface  150 mm   Failed 

ÄSPÖ‐TBM‐PL3336 R9D  2012  surface  140 mm  Low 

ÄSPÖ‐TBM‐PL3336 R9D  2012  semi‐surface  420 mm  Failed 

ÄSPÖ‐TBM‐PL3335 R10    2012  semi‐surface  660 mm  Low 

ÄSPÖ‐TBM‐PL3335 R10C  2012  surface  140 mm   Failed 

ÄSPÖ‐TBM‐PL3335 R2  KA3334C02  2011  surface  300 mm  High 

ÄSPÖ‐TBM‐PL3332 R6  KA3331A02  2011  surface  300 mm  High 

ÄSPÖ‐TBM‐PL3335 R9  KA3335I04  2011  surface  150 mm  High 

ÄSPÖ‐TBM‐PL3335 R10  KA3335I02  2011  surface  250 mm  High 

ÄSPÖ‐TBM‐PL3334 R3  2010  surface  120 mm  High 

ÄSPÖ‐TBM‐PL3334 R4  2010  surface  120 mm  High 

ÄSPÖ‐TBM‐PL3332 R7  2010  surface  120 mm  High 

ÄSPÖ‐TASS‐PL64 R1  KS0063B01  2011  surface (EDZ)  180 mm  High 

ÄSPÖ‐TASS‐PL64 R1  KS0062I01  2011  surface (EDZ)  200 mm  High 

ÄSPÖ‐TASS‐PL62 R2  KS0062I01  2011  deep  750 mm  High 

ÄSPÖ‐TASS‐PL62 R2  KS0061H02  2011  deep  800 mm  High 

ÄSPÖ‐TASS‐PL61 R3  KS0061H04  2011  surface (EDZ)  270 mm  High 

ÄSPÖ‐TASS‐PL61 R4  KS0061H04  2011  deep  750 mm  High 

ÄSPÖ‐TASS‐PL63 R6  KS00063A01 2011  surface (EDZ)  130 mm  Failed 

ÄSPÖ‐TASS‐PL63 R6  KS00063A01 2011  semi‐surface  430 mm  High  

Interestingly, in most of the samples the strain gauge in the 45° direction often does not give results that would be expected with CHILE conditions, which would be the mean of the tangential and axial strain gauge values. For the high quality tests, the measurements were repeated multiple times with the same results. The interpreted results for individual strain gauges are also very close to each other 69.3 GPa and 67.8 GPa, giving clear indications of low anisotropy and low deviation. These consistent values give high confidence for these results. The tests where one or more strain gauges showed non-linear behaviour were interpreted as low confidence results. These results often showed big differences between different strain gauge rosette values; however, the mean value of the results could be still acceptable if anisotropic behaviour is present. If the interpretation were impossible or the tests were not able to be completed, the test was stated as ‘failed’.

In the third campaign, existing holes 1-3 and 10 were extended by overcoring and sidecoring. Also, holes 3B, 9C, 9D and 10C were drilled but used only for biaxial testing (Figure 5-12).

52

Figure 5-11. The high quality results of the biaxial sample from hole 1 in TBM-tunnel (ÄSPÖ-TBM-PL3335 R1).

Figure 5-12. Overview of the TBM tunnel measurements during the 2012 campaign.

5.1.2 TASS-tunnel Results

An overview of the TASS tunnel measurements with the measurement locations are shown in Figure 5-6. LVDT measurements and biaxial testing are documented in detail in Appendices 7 and 9.

3 – OverCoring

Looking along tunnel axisx=East, y=North, z=Up

1 – OverCoring

2 – SideCoring

10 – SideCoring

9C & 9D biax samples

10C biax sample

3B biax sample

53

Biaxial test results All successful biaxial test results are shown in Table 5-3. The failed tests are not shown in the Table. For individual rosettes, the biaxial results are calculated (set A and set B in Figure 5-13). The mean deviation of Young's modulus in individual rosettes in the near excavation surface samples is 17.4 GPa compared to samples deeper in the hole where the mean deviation is 6.7 GPa. In the R6 surface sample, the variation in Young's modulus between the rosettes is 56.1 GPa, while the difference in other samples between rosettes varies from 1.8 GPa to 22.8 GPa. If the hole R6 results from the surface samples are ignored, the mean and standard deviation of Young's modulus for individual rosettes in the surface samples are 57 GPa/10.4 GPa and for individual rosettes in deeper rock samples are 59 GPa/6.7 GPa. A similar approach was used to establish the Poisson’s ratio (Figure 5-13). A Young’s modulus of 58 GPa and a Poisson’s ratio of 0.23 were selected for the inversion modelling. Table 5-3. All successful biaxial test results in the TASS tunnel.

      Rosette 1 Rosette 2 All data 

Sample ID  Depth  ν E (GPa) ν E (GPa) ν  E (GPa)

ÄSPÖ‐TASS‐PL64 R1  edz  0.34  68.49  0.25  55.69  0.29  61.43 

ÄSPÖ‐TASS‐PL64 R1  deep  0.25  54.70  0.27  67.30  0.26  60.40 

ÄSPÖ‐TASS‐PL62 R2  edz  0.25  53.85  0.20  67.31  0.23  59.83 

ÄSPÖ‐TASS‐PL62 R2  deep  0.23  53.79  0.19  59.29  0.21  56.41 

ÄSPÖ‐TASS‐PL61 R3  deep  0.17  48.94  0.26  67.67  0.21  56.80 

ÄSPÖ‐TASS‐PL61 R4  edz  0.21  41.03  0.19  63.78  0.20  49.94 

ÄSPÖ‐TASS‐PL63 R6  semi‐surface  0.25  62.85  0.20  61.09  0.22  61.96 

54

Table 5-4. Results for the TASS tunnel biaxial testing in different domains. The mean deviation is calculated using mean values of rosettes for each sample.

Results of all data (high confidence) σv E (GPa) σE (GPa)  Samples 

Surface 0.24 0.04 57.1 6.2 3 Semi-surface 0.22 ‐  62.0 ‐  1 Deep 0.23 0.03 57.9 2.2 3 Surface + semi-surface 0.24 0.04 58.3 5.6 4 Semi-surface + deep 0.23 0.02 58.9 2.7 4 All depths 0.23 0.03 58.1 4.2 7

Figure 5-13. Young's modulus and Poisson’s ratio for individual rosettes (set A and set B) for surface and deep LVDT measurement pilot hole cores.

Interpretation of the in situ state of stress All LVDT measurements gave stable responses and no data were rejected (Appendix 7). The measured stress release induced changes in diametrical distances varied in the overcoring tests between -5m and 65m and in the side-coring tests between -45m and 125m. Except for the overcoring measurement R6, the differences in the surface and deep measurements were less than expected (Figure 5-14). The inversion modelling was carried out using two models. The first model was for the full-length pilot holes and the second for the full-length pairs of side coring holes, i.e., phases 1 and 3 in Figure 5-15. Leaving out the mid-length overcoring and sidecoring, the Phase 2 results had very small errors in the modelled convergence. For the tunnel surface, a model with mean point to point distance of 0.5 m was constructed based on detailed laser scanning model (Figure 5-16). The position and orientation of the test holes was measured with three-dimensional photogrammetry and tunnel survey, and the relative distances between the side coring holes were manually checked. For the best fit in situ stress solution, the LVDT-sensor head displacements were imported to Excel and the solving was done using macros. For both surface and deep measurements, 20 convergences are used to solve for the six unknown stress tensor components.

0

20

40

60

80

100

R1 R2 R3 R4 R6

Young's M

odulus (GPa)

Sample

Surface A

Surface B

Deep A

Deep B

0

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

R1 R2 R3 R4 R6

Poisson's ration ()

Sample

Surface A

Surface B

Deep A

Deep B

55

Figure 5-14. Comparison of surface (dashed lines) and deep (solid lines) measurement responses/convergences in the LVDT sidecoring measurement R2.

1 – Pilot hole geometry  2 – ignored phase    3 – Final geometry 

Figure 5-15. Calculation phases for Äspö TASS-tunnel inversion calculation.

56

Figure 5-16. Truncated Examine3D element mesh for pilot hole phase 1; the tunnel mesh is based on simplified laser scanned profiles.

Based on the deep measurements, the major in situ stress component is almost horizontal trending 134°–314° with a magnitude of 24.6 MPa (Figure 5-17). The intermediate and minimum principal stress components are not horizontally/vertically aligned, but rotated 30° around the major horizontal stress axis. The value for the vertical component is 10.5 MPa—which is about 15 % less than the weight of the overburden (Table 5-5 and Table 5-6). For the tunnel surface measurements, the trend of the maximum in situ principal stress deviates 10 from the deep measurements (325°–145°), but the magnitude is clearly lower, being 18.9 MPa. The two other principal stress components are about the same as found in the deep measurement, but without the rotation around the maximum. The mean unexplained convergence, i.e., internal error of the solution, is 37% for the surface measurements and 27 % for the deep measurements. However, the correlation between the measured convergence and the best fit solution indicates that the solution for the surface measurements is questionable because the slope of the linear best-fit estimate is 0.67 instead of the assumed 1.0, as it is for the solution for the deep measurements (Figure 5-18). The selection of the convergence values, final versus stop of overcoring, has a notable effect on the resulting major and minor principal stress magnitudes. These differences are most likely related to the coring-induced heat because, in all overcoring cases, the convergence increases after the coring is stopped but, in the sidecoring cases, the measured convergence decreases.

57

Figure 5-17. Best fit solution for in situ principal stresses based on deep and surface measurements. Two solutions are given based on the overcoring stop time convergence values and the final convergence values.

Table 5-5. Best fit solution for in situ principal stresses based on deep and surface measurements. Two solutions are given based on the overcoring stop time convergence values and the final convergence values. The stress magnitudes are in MPa and trends are in the co-ordinate system RT90.

Table 5-6. Best fit solution for horizontal and vertical in situ stress components based on deep and surface measurements. Two solutions are given based on overcoring stop time convergence values and final convergence values. The stress orientations are in co-ordinate system RT90.

Case 1 trend,° plunge,° 2 trend,° plunge,° 3 trend,° plunge,°

Deep, Final values 24.6 134 7 13.6 40 29 9.2 238 60

Deep, OC Stop values 23.3 137 8 13.3 43 28 8.5 242 61

Surface, Final values 18.9 325 7 13.4 55 5 9.8 180 81

Surface, OC Stop values 17.6 333 10 13.1 243 1 9.2 149 80

Case H trend, ° h trend, ° V

Deep, Final values 24.3 136 12.6 226 10.5

Deep, OC Stop values 23.0 139 12.2 229 9.8

Surface, Final values 18.8 144 13.4 234 9.9

Surface, OC Stop values 17.3 153 13.1 243 9.5

58

Figure 5-18. Correlation between measured and the best fit solution convergence values of surface (EDZ) and deep measurements in the Äspö TASS-tunnel.

5.1.3 TBM-tunnel Results

An overview of the TBM tunnel measurements is shown in Figure 5-20. LVDT measurements and biaxial test results for the TBM tunnel are documented detail in Appendices 8 and 10. Biaxial test results In the year 2010 campaign, three samples were biaxially tested and they were of high quality. In the year 2011, four samples from the TBM tunnel were tested with one sample where two strain gauges from one rosette failed and that sample was judged of low quality with only one functioning strain gauge. In the year 2012 campaign, multiple tests failed—mainly due to the poor quality of the samples. Most of the samples had 1–2 mm shallow grooves from the diamond coring which made the surface uneven and caused the strain gauges to loosen from the rock surface. Also, it was noticed that, of the two different glues used in the 2012 biaxial testing, the lower viscous glue version CC-33A failed several times compared to the tests using the more viscous CC-35 glue. Both glues were recommended for the strain gauges used in the testing. All successful biaxial test results for the TBM tunnel samples are shown in Table 5-7. The biaxial results were calculated for individual rosettes (set A and set B in Figure 5-20). The individual successful results from rosette 2 in tests R3B and R10 are included in the Figure. The results from R3B are combined with R3 in Figure 5-20. If the individual failed rosettes are ignored, the mean deviation of the Young's modulus in samples near the excavation surface is 8.3 GPa compared to samples deeper in the

y = 1.01xR² = 0.84

y = 0.67xR² = 0.82

‐60

‐40

‐20

0

20

40

60

80

100

‐60 ‐40 ‐20 0 20 40 60 80 100

Calculated  Convergence (microstrain)

Measred Convergence (microstrain)

Deep, final

Surface, final

Measured Convergence (microstrain)

59

hole where the mean deviation is 5.7 GPa. The differences in high quality tests between individual rosettes vary from 1.2 GPa to 12.6 GPa. A similar approach was used to establish the Poisson’s ratio and the differences in high quality tests between individual rosettes vary from 0.1 to 0.4 (Figure 5-19).

Figure 5-19. Overview of the TBM tunnel measurements.

Table 5-7. All successful biaxial test results in the TBM tunnel. High quality results are marked in green; failed results are marked with dark red; and low quality tests in light red.

      Rosette 1 Rosette 2 All data 

Sample ID  Depth ν E (GPa) ν E (GPa)  ν  E (GPa)

ÄSPÖ‐TBM‐PL3335 R1  deep  0.25  69.27  0.26  67.76  0.26  68.77 

ÄSPÖ‐TBM‐PL3335 R2  semi‐surface  0.49  52.70  0.31  77.32  0.45  60.73 

ÄSPÖ‐TBM‐PL3335 R2  deep  0.26  59.75  0.22  65.80  0.25  60.78 

ÄSPÖ‐TBM‐PL3334 R3B  semi‐surface  1.01  90.88  0.39  81.55  0.82  74.17 

ÄSPÖ‐TBM‐PL3336 R9D  surface  0.23  56.53  0.73  37.89  0.55  38.87 

ÄSPÖ‐TBM‐PL3335 R10  semi‐surface  0.47  109.35  0.28  76.81  0.38  91.10 

ÄSPÖ‐TBM‐PL3335 R2  surface  0.24  64.29  0.20  68.25  0.22  66.21 

ÄSPÖ‐TBM‐PL3332 R6  surface  0.25  54.74  0.27  67.30  0.26  60.37 

ÄSPÖ‐TBM‐PL3335 R9  surface  0.21  59.30  0.19  46.70  0.20  52.25 

ÄSPÖ‐TBM‐PL3335 R10  surface  0.22  55.38  0.20  57.30  0.21  56.32 

ÄSPÖ‐TBM‐PL3334 R3  surface  0.20  79.33  0.22  75.39  0.21  77.31 

ÄSPÖ‐TBM‐PL3334 R4  surface  0.17  68.59  0.22  67.38  0.19  67.98 

ÄSPÖ‐TBM‐PL3332 R7  surface  0.16  66.49  0.20  67.69  0.18  67.09 

3 – OverCoring‐PILOT 102 cm‐OC    35…99 cm

Looking along tunnel axisx=East, y=North, z=Up

1 – OverCoring‐PILOT  105 cm‐OC     64 cm

2 – SideCoring‐PILOT  106 cm‐OC     103 cm

10 – SideCoring‐PILOT 90 cm‐OC      74 cm

9D biax samples‐PILOT 55 cm

3B ‐biax sample‐PILOT 55 cm

9C biax samples‐PILOT 31 cm

10C biax sample‐PILOT 27 cm

7 – OverCoring‐PILOT  35 cm‐OC     57 cm

4 – OverCoring‐PILOT  59  cm‐OC     35 cm

6 – SideCoring‐PILOT  54 cm‐OC     50 cm

9 – SideCoring‐PILOT  45 cm‐OC     40 cm

5 – OverCoring‐PILOT 58 cm‐OC    12,5 cm

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Similar to the biaxial results in the TASS tunnel, the values of two rosettes are summed for further analyses (Table 5-8). The mean value and standard deviation of the Young's moduli in the surface samples is 63.9 GPa/8.3 GPa and for deeper rock samples 64.8 GPa/5.7 GPa. For the Poisson's ratios, the respective values are 0.21/0.02 and 0.26/0.01. The results are presented in Table 5-8. A Young’s modulus of 64.5 GPa and a Poisson’s ratio of 0.22 were selected for the inversion modelling.

Figure 5-20. Young's modulus and Poisson’s ratio for individual rosettes (set A and set B) for surface and deep LVDT measurement pilot hole cores.

0

10

20

30

40

50

60

70

80

90

R1 R2 R3 R4 R6 R7 R9 R10

Young's modulus (GPa)

Surface A

Surface B

Deep A 

Deep B

0.1

0.15

0.2

0.25

0.3

0.35

0.4

0.45

R1 R2 R3 R4 R6 R7 R9 R10

Poisson's ratio ()

Surface A

Surface B

Deep A 

Deep B

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Table 5-8. Results for TBM tunnel biaxial testing in different domains.

  Results of all data Results of high quality data only 

v  E (GPa) Samples  v σv E (GPa) σE (GPa)  Samples 

Surface  0.21  63.0 8 0.20 0.02 63.9 8.3  7 

Semi‐surface  0.33  78.6 3 ‐ ‐ ‐ ‐  0 

Deep  0.25  64.8 2 0.26 0.01 64.8 5.7  2 

Surface + semi‐surface  0.24  67.2 11 0.21 0.02 63.9 8.3  7 

Semi‐surface + deep 0.30  73.0 5 0.26 0.01 64.8 5.7  2 

All depths  0.24  66.9 14 0.22 0.03 64.1 7.5  9 

Interpretation of the in situ state of stress All completed LVDT measurements gave stable responses and no data were rejected (Appendix 8). The measured stress release induced changes in diametrical distances varied in the overcoring tests between 2m and 171m and in the side-coring tests between -64m and 94m. Convergences in near surface measurements are somewhat higher than in deep measurements, except in sidecoring measurement R3 where the ratio has the opposite trend. The inversion modelling was originally carried out after each measurement campaign, but results in this report are shown for all measurements and separately for surface and deep measurements. For this final solution, the LVDT-head displacements were collected from the three separate simulations. As in the case of the TASS-tunnel modelling, the first model is for full length pilots and overcoring attempts if any and the second model is for final length side coring holes. The locations of pilot and sidecoring holes were interpreted from the photogrammetric digital elevation model made after the measurements and for the TBM-tunnel a fitted perfectly circular cross-section was assumed. The relative distances between the side coring holes was manually checked and corrected before final modelling. For the best fit in situ stress solution, the LVDT-sensor head displacements were imported to Excel and the solving was implemented with macros. For the surface measurements and deep measurements, 28 and 16 convergences respectively were used to solve the six unknown stress tensor components. An example of the Examine3D TBM-model used for the analyses is shown in Figure 2-15. Based on all the measurements, the major in situ stress component is sub-horizontal, trending 140° and plunging 18° with a magnitude of 25.5 MPa (Figure 5-21, Table 5-9). The intermediate and minimum principal stress components are rotated 39° around the maximum. The value for the vertical component is 18.4 MPa, which is about 58 % greater than the weight of overburden (Table 5-10). Principal stress magnitudes for deep and surface measurements are within 1.9 MPa if compared to the solution for all measurements. The biggest difference is that, for the deep measurements, the orientations for the major and intermediate principal stresses are changed, and also give the highest magnitudes.

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The mean unexplained convergence, i.e., internal error of the solution, is 20 % for all measurements, 15 % for surface and 13 % for the deep measurements. The solution constrained to the vertical and horizontal planes gives the highest error of 39 %. The correlation between the measured convergence and the best fit solution indicates a reliable solution because the slope of the linear best-fit estimate is close to unity and the correlation coefficient is also close to unity (Figure 5-22).

Figure 5-21. Best fit solution for in situ principal stresses based on all (black symbols), all constrained to horizontal/vertical, surface and deep measurements in Äspö TBM-tunnel. The stress magnitudes are in MPa and trends are in the co-ordinate system RT90.

Table 5-9. Best fit solution for in situ principal stresses based on all, surface and deep measurements in the Äspö TBM-tunnel. The stress magnitudes are in MPa and trends are in the co-ordinate system RT90.

Solution  1  trend,°  plunge,° 2  trend,°  plunge,° 3  trend,°  plunge,°

All  25.5  140  18  22.5  34  39  13.9  249  46 

Surface  24.7  137  26  21.1  25  38  12.9  253  40 

Deep  27.4  39  36  23.5  132  3  15.0  226  54 

All H/V  26.7  117  0  21.0  27  0  19.4  0  90 

                 

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Table 5-10. Best fit solution for horizontal and vertical in situ stress components based on all, surface and deep measurements in Äspö TBM-tunnel. The stress magnitudes are in MPa and the trends are in the co-ordinate system RT90.

Solution  H  trend,° h  trend,°  V 

All  25.0  152  18.5  242  18.4 Surface  23.6  154  16.7  244  18.4 Deep  23.7  161  23.0  251  19.3 All H/V  26.7  117  21.0  207  19.4 

 All  

 Surface 

 Deep 

 All H/V 

Figure 5-22. Correlation between measured and the best fit solution convergence values of all, surface and deep measurements in the Äspö TBM-tunnel.

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5.1.4 Summary of Äspö measurements

The Young's moduli from the TBM tunnel as a function of angle from the vertical axis (0–90°) shows a dependency with the angle. A linear line can be fitted to accommodate the results starting from 52 MPa when vertical to 74 MPa when horizontal.

The Poisson's ratio in the TBM tunnel varies from 0.20 at the tunnel surface to 0.26 at a depth of 90 cm. However Poisson's ratio in the TASS tunnel does not show this correlation with depth but varies from 0.20 in the vertical direction to 0.25 in the horizontal direction—which the data from the TBM tunnel does not exhibit. This suggests that the excavation damage zone around the TASS tunnel is deeper around the perimeter and in the TBM is only is present on the surface, decreasing quickly with the depth.

From the LVDT measurements made in the TASS- and TBM tunnel, only the surface measurements in the TASS tunnel turned out to be less reliable than the others. This supports the idea of the more extensive excavation damaged zone (EDZ) caused by the drill and blast excavation method. From the biaxial LVDT measurements done in the TASS- and TBM tunnel, only the surface measurements in the TASS tunnel turned out to be less reliable than the others. In overall the TASS tunnel biaxial results have higher internal error compared to other biaxial measurements. This supports the assumption that an excavation damaged zone (EDZ) caused by D&B excavation method exists in the TASS tunnel surface.

Figure 5-23. High confidence results of Young's modulus and Poisson's ratio as a function of angle from the vertical (0–90°).

Figure 5-24. High confidence results of Young's modulus and Poisson's ratio as a function of distance from the tunnel surface.

0

10

20

30

40

50

60

70

80

90

0 15 30 45 60 75 90

Young's modulus (GPa)

Angle from vertical axis (°)

TASA TBM

TASS

0

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0 15 30 45 60 75 90

Poisson's ratio

Angle from vertical axis (°)

TASA TBM

TASS

0

10

20

30

40

50

60

70

80

90

0 20 40 60 80 100

Young's modulus (GPa)

Depth from tunnel surface (cm)

TASA TBM

TASS

0

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0 20 40 60 80 100

Poisson's rartio

Depth from tunnel surface (cm)

TASA

TASS

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As a final phase, a combined solution using the deep TASS and all TBM measurements was made. The resulting in situ state of stress does not diverge much from the local solutions, but the reliability is slightly lower; the slope between measured and calculated convergences is 0.92 and correlation coefficient R2 = 0.87 (Figure 5-25 and Table 5-11 and Table 5-12).

Figure 5-25. Best fit solution for in situ principal stresses based on all TBM tunnel and TASS tunnel deep measurements (black symbols) and for both tunnels separately. The stress magnitudes are in MPa and the trends are in the co-ordinate system RT90.

Table 5-11. Best fit solution for in situ principal stresses based on the Äspö TBM tunnel and TASS tunnel deep LVDT measurements. The stress magnitudes are in MPa and trends are in the co-ordinate system RT90.

Solution  1  trend,°  plunge,° 2  trend,°  plunge,° 3  trend,°  plunge,°

TBM & TASS  24.9 131  23  19.1  20  40  12.9 243  41 

TBM All  25.5 140  18  22.5  34  39  13.9 249  46 

TASS Deep  24.6 134  7  13.6  40  28  9.2  238  60 

                 

Table 5-12. Best fit solution for horizontal and vertical in situ stress components based on the Äspö TBM tunnel and TASS tunnel deep LVDT measurements. The stress magnitudes are in MPa and the trends are in the co-ordinate system RT90.

Solution  H  trend,° h  trend,°  V 

TBM & TASS  23.7  140  15.8  230  17.3 TBM All  25.0  152  18.5  242  18.4 TASS Deep  24.3  136  12.6  226  10.5 

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6 DISCUSSION AND CONCLUSIONS

6.1 Sensitivity test

All the field trials and sensitivity tests indicate that the LVDT-cell is a robust tool that can provide reliable and consistent results in most field operating conditions, providing the CHILE assumptions used for the analyses are not violated, i.e, the rock is not discontinuous, strongly heterogeneous or foliated and/or does not exhibit time dependent permanent deformations. The drilling heat-induced thermal expansion of the rock cylinder can produce stress errors of 3 MPa in the measurement plane. This effect can be minimised by using temperature controlled flushing water; a water temperature at a few degrees below the in situ temperature of the rock is suggested. Further, sufficiently long-term monitoring should be undertaken after stress release to obtain reliable convergence values. Total thermo-mechanical recovery can take up to over six hours but, if the inner surface temperature of the overcored rock cylinder does not change more than 10 °C, then the error caused by thermal expansion should be after two hours about a half, corresponding to about 5m deformation or stresses of 1.5 MPa in the measurement plane (Chapter 3).

6.2 Selection of test location

Careful selection of the test location and thorough thinking of the whole process significantly decreases the risk of problems when carrying out the measurements. Designing overcoring and sidecoring The experiences with LVDT measurements below -360 m in the ONKALO are that ring discing will cause problems in tunnel roof holes and shaft wall holes perpendicular to the major horizontal stress component. If there is the possibility for ring discing in high stress conditions, overcoring with a smaller diameter drill, drill bits of 35mm and 65 mm in diameter, can be applied for scoping purposes before the actual stress measurements. After the scoping or using previous knowledge about the major principal stress direction, the overcoring and sidecoring can be planned. Overcoring is usually possible, at least in the direction of the major principal stress because then the stress concentrations around the testing hole are less. Sidecoring should be used to avoid ring discing problems. The sidecoring locations should be optimally situated to provide enough data for all directions. In Figure 6-1, a hypothetical situation from the ONKALO shaft measurement campaign at level -420 is presented where the scoping calculations show ring discing to the North but not to the NW. After the scoping calculations, it is suggested that the overcoring be oriented to West and East but sidecoring should be located to the SE and SW.

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Figure 6-1. Illustration on how the overcoring and sidecoring locations are decided based on the scoping holes for high stress conditions in the vertical shaft.

Number of measurements The number of measurements has to be designed so that influence of all six stress tensor components can be measured. Three measurements, covering at least 90° around the tunnel or shaft axis, is the minimum number of tests when using overcoring in good conditions (e.g. raise bored shaft). When only sidecoring is used, a minimum of 5–6 tests are required. Using both overcoring and sidecoring, a minimum number of measurements is for example 1–2 overcorings and 3–4 sidecorings (Figure 6-1). Effect of fractures, EDZ and anisotropy The measurements should be made in unfractured rock domains avoiding the ends of the blast rounds, heterogeneity and structures in the nearby rock mass. Special attention should be given to avoid fractures that will intersect the overcoring beyond the rock surface. Although instant stress release caused by a fracture does not fundamentally harm the measurement, it can cause instability of the cell mounting or deformations which are hard to interpret. Therefore fractures should be avoided as much as possible and the best way to do this is to carefully plan the measurements by selecting the locations in situ from a clean rock surface. If a fracture intersects the planned overcoring hole during the pilot coring stage, the measurement can still be conducted— either by deepening the hole or switching to sidecoring measurement. Although locations can be decided using detailed geological mapping (see Figure 5-4) the measurements on shotcreted surface are not recommended. In tunnels excavated using the drill and blast (D&B) method, the excavation damage zone (EDZ) especially at the end of the blast round should be avoided because there is high fracture intensity and at the uneven surface the stresses are possibly already released. The measurements

OvercoringOvercoringmajorprincipalstressdirection

scoping holeswith ring disking

horizontal

intact scoping holes

Overcoringand sidecoring locationsare decided basedon the scoping holes

69

in the TASS tunnel suggest that the variation of results is higher in the first 0.5 m (in the EDZ) and that the measurements should be conducted between 0.5–1.0 m. Although no measurements have been conducted with the device in systematic anisotropy, the LVDT measurement and interpretation system are applicable for ‘standard’ anisotropy. This means that special care should be paid in testing the pilot cores and, if the core orientations are not favourable, to define all elastic constants; additional laboratory test of sub-cores may be needed. The anisotropic solution also sets requirements for the numerical code used.

6.3 Biaxial testing

Biaxial testing using strain gauges and the Hoek Cell is a highly demanding task. Usually it is also the bottleneck in the LVDT measurements as the biaxial testing is time consuming to conduct. An experienced and specialised professional is recommended for strain gauge measurements. Two rosettes with long strain gauges should be glued to each sample to avoid problems with heterogeneity in the rock material. The strain gauge glue should be compatible with the strain gauges and delivered by the same manufacturer. The variation in Young's modulus around the TBM tunnel can be caused by various factors. One assumption is that the samples in the roof are damaged during the stress release in overcoring and also there is the unlikely possibility of rock stress damage in the roof although the state of stress does not seem high enough. Overall, the number of samples is too low to draw conclusions. Experience to date suggests biaxial results deviate significantly in heterogeneous rock . To compensate for this heterogeneity, it is recommended to use mean values for Young's modulus and Poisson's ratio for the inversion modelling.

6.4 Experiences with operating the probe

Experiences with operating the probe indicate that typically 1–2 tests can be done per measurement day, depending on the measurement conditions (narrow shaft or quiet tunnel). Typically it takes a week to conduct LVDT measurements in one tunnel section with 5–6 measurement points. However, campaigns can be conducted faster, for example in the ÄSPÖ TBM tunnel the measurements had to be made during the night or at the weekend. The time for testing was reduced by using two drills at the same time. The measurement team should consist of 1–2 skilled technical persons with a professional driller. The detailed reporting of the current measurement depth during the overcoring is vital and requires constant access to the drill. The overcoring depth can be marked on the drill bit before starting the measurement to ease the depth measurement (see white markings on the drill bit in Figure 2-2). Maintaining a detailed log during the measurement is also necessary. One of the most important parts of the measurement is surveying the hole locations, especially when sidecoring measurements are used. Surveying accuracy is critical when using sidecoring holes and the accuracy of the pillar width at the bottom of the holes

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should be at least 5 mm. Adjustable steel or plastic pipes and laser scanning can be used to help when measuring the angles between holes.

6.5 ONKALO verification case

The ONKALO verification test demonstrated that the LVDT-cell operates reliably in field conditions. The orientation for the maximum principal stress is in good agreement with ring discing observations. Further, magnitudes of the in situ principal stresses are very close to those obtained using the LVDT measurement and the internal errors for both methods are about the same. The correctness of absolute stress magnitude values could not be evaluated because the accuracy the ONKALO area stress model is not that good.

6.6 Äspö verification case

The two tests with the LVDT-cell carried out at the Äspö HRL 450 m level are compared to each other and to older results. The calculated principal stresses are compared to the results from the Borre probe in two vertical boreholes and one horizontal borehole, Table 6-1 and Table 6-2. The locations of the boreholes, related to the LVDT measurement are shown in Figure 5-3.

Figure 6-2. Location of boreholes were overcoring measurements with the Borre probe has been carried out.

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Table 6-1. Calculated principal stresses from LVDT measurements in the two tunnels and the Borre probe results from the horizontal of the two orthogonal holes.

Table 6-2. Calculated horizontal and vertical stress components.

Measurement  H  trend,° h  trend,° V 

LVDT             TBM, all  25.0 152  18.5  242  18.4   TASS, deep  24.3 136  12.6  226  10.5   Mean  23.7 140  15.8  230  17.3  Borre (mean)   KF0093A01  25.7 137  10.2  227  18.0   KK0045G01  25.9 139  13.5  229  15.5   KA3579G  34.1 153  13.3  243  17.6   Mean  28.3 145.1  12.5  235.1  17.1   Christiansson & Jansson, 2003   24±5 136  10‐13 226  15‐20

Visual comparison of the results in Table 6-1 and Table 6-2 is subjective. However, the estimated magnitude and orientation of the major principal stress is within the estimated range by Christiansson & Jansson (2003), except for overcoring results in borehole KA3579G. There are uncertainties in both methods when estimating the magnitude and orientation of the intermediate and minor principal stresses. It is notable that the results of impression packer tests during hydraulic fracturing in the horizontal hole KF0093A01 gave both vertically and gently dipping induced fractures. The borehole was drilled in the direction of the major horizontal stress. This indicates that the difference in stress magnitude between σ2 and σ3 is minor. A significant uncertainty in stress measurements with both overcoring with the Borre probe and using the LVDT-cell is the Young’s modulus. The range of the estimated Young’s modulus for the successful Borre tests in Table 6-1 and Table 6-2 is between 50 and 73 GPa. This range influences the calculated stress magnitudes. The same applies for the LVDT results. The mean Young’s modulus used for the TASS tunnel is lower than the value for the TBM tunnel and the calculated stress magnitudes are also

Measurement  1  trend,°  plunge,° 2  trend,° plunge,° 3  trend,°  plunge,°

LVDT                     TBM, all  25.5  140  18  22.5 34  39  13.9 249  46 

  TASS, deep  24.6  134  7  13.6 40  28  9.2  238  60 

  Mean  24.9  131  23  19.1 20  40  12.9 243  41 

 

Borre (mean)                  

  KF0093A01  29.8  322  31  14.8 100  52  9.4  218  21 

  KK0045G01  27.4  134  20  17.1 24  43  10.4 242  40 

  KA3579G  34.2  153  3  17.7 257  80  13.1 62  10 

  Mean  28.4  326  4  17.5 70  73  12.1 234  17 

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lower for the TASS tunnel. It is reasonable that the drill and blast tunnel exhibits lower Young’s modulus compared to the TBM tunnel, assuming more damage and microcracks in the drill and blast tunnel. But, on the other hand, the scatter in the Young’s modulus measurements from the TBM tunnel is larger (standard deviation 7.5 GPa compared to 4.3 in the TASS tunnel). Instead of comparing the individual principal stress components of the stress tensor, it is often more informative to compare the mean stress values, (1+2+3)/3). Figure 6-3 compares the mean stress from the LVDT measurements to the means stress from the overcoring measurements in three boreholes that span the volume containing the LVDT measurements. Based on this comparison, the state of stress from the LVDT measurements is comparable to that determined from the overcoring data. When the components of the tensor are examined, it can also be concluded that the estimated state of stress determined using the new LVDT-cell is within the range proposed by Christiansson & Jansson 2003 based on the traditional overcoring and hydraulic fracturing techniques.

Figure 6-3. Comparison of the mean stress from individual overcoring measurement in three different boreholes with the LVDT results.

Table 6-3. Comparison of test results to earlier best estimate of state of stress at the 450 m level.

σH

MPa σH trend (RT90) σh

MPa σv

MPa

Christiansson & Jansson (2003) 

 24 ±5 

 136 

 10 ‐ 13

 15‐20

This study Deep, > 0.5 m 

 24‐25

 136°‐152° 

 12‐18 

 10‐18

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6.7 Analyses and uncertainties

The CHILE (Continuous, Homogeneous, Isotropic (or transversely isotropic) and Linear Elastic) assumptions used for the LVDT analyses are identical to the CHILE assumptions used in traditional overcoring and hydraulic fracturing stress measurement techniques. Consequently, so long as these assumptions are not violated, the solutions from any technique will be acceptable. The major advantage of the LVDT-Probe methodology is the scale of the measurements. It has been shown by Martin et al. (1990) that one large-scale measurement can provide a reasonable estimate and level of confidence in the in situ stress state, while many small scale measurements are required to provide the same level of confidence. Figure 6-4 shows the variation in the mean stress (1+2+ )/3 as a function of the number of measurements used in the LVDT numerical inversion. These results show that, regardless of the number of measurements, the methodology provides a good estimate of the stress magnitudes.

Figure 6-4. Mean stress as a function of LVDT measurement used. Data are from Äspö TBM-tunnel measurements and the used measurement combinations so that they could produce reasonable estimation of in situ stress.

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7 FUTURE WORK

The spatial variation of the Young's modulus and Poisson's ratio around TBM tunnels could be investigated in more detail in order to understand whether the use of mean values in the inversion calculation is a reasonable approach. This can be done by conducting Ground Penetrating Radar measurements around the tunnel perimeter and repeating some failed deep biaxial tests. The biaxial testing procedure can be upgraded by manufacturing a miniature LVDT-cell and using it in inside the biaxial sample during the biaxial measurement. In this way, the problems with strain gauge gluing can be avoided. This development work is undergoing.

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REFERENCES

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Andersson C.J., Martin C.D. and Christiansson R. 2004. SKB’s Pillar Stability Experiment, Sweden. ARMA/NARMS Houston, 2004. Paper 04-503.

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Emsley S., 1997. ZEDEX: a study of damage and disturbance from tunnel excavation by blasting and tunnel boring, SKB technical report 97–30, Sweden. 198 p.

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Hooker, V.E, & Bickel, D.L. 1974. Overcoring equipment and techniques used in rock stress determination. USBR IC-8618.

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Martin, C. D., Read, R. S., and Chandler, N. A. (1990). Does scale influence in situ stress measurements?– Some findings at the Underground Research Laboratory. In da Cunha, A. P., editor, Proc. First Int. Workshop on Scale Effects in Rock Masses, Loen, Norway, pp. 307–316. A.A. Balkema, Rotterdam.

Martino, J. B., P. M. Thompson, N. A. Chandler, and R. S. Read (1997): The in situ stress program at AECL's Underground Research Laboratory, 15 years of research (1982–1997). - 100 pp., NWMD, Ontario Hydro, Canada.

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Wiles, T. D. and Kaiser, P. K. (1990). A new approach for the statistical treatment of stress tensors. In Herget, G., Arjang, B., B ́etournay, M., Gyenge, M., Vongpaisal, S., and Yu, Y., editors, Proc. Stresses in Underground Structures, Ottawa, pages 62–76, Ottawa, Canada. Canadian Government Publishing Centre.

Worotnicki, G. 1993. CSIRO triaxial stress measurement cell. In Comprehensive Rock Engineering — Principles, Practice & Projects (ed. Hudson J.A.), Vol 3, pp. 329–394. Oxford: Pergamon Press.

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APPENDICES

List of appendices Appendix 1: Sidecoring study. Appendix 2: Example of LVDT measurement log sheet. Appendix 3: Review of Examine3D approach and Solve algorithm for LVDT-cell Appendix 4: LVDT Sensitivity study. Appendix 5: Strain gauge measurements in ONKALO exhaust air shaft level -265 and biaxial testing of the pilot cores. Appendix 6: LVDT-cell measurements in ONLKALO exhaust air shaft level -265 Appendix 7: LVDT-cell measurements in Äspö TASS-tunnel Appendix 8: LVDT-cell measurements in Äspö TBM-tunnel Appendix 9: Biaxial test results in Äspö TASS-tunnel Appendix 10: Biaxial test results in Äspö TBM-tunnel

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Appendix 1

Sidecoring study

82

The two-dimensional elastic Phase2 simulation was done to study the possibility of interpreting the in situ state of stress based on sidecoring. In sidecoring, a bigger release hole is drilled close to the pilot hole. This will result in only partial stress release but also in a measurable convergence in the pilot hole. (Figures A1.1, A1.2 and A1.3). The simulation consists of two phases; pilot hole drilling and sidecoring. Dimensions for the pilot hole and sidecoring holes were 127 mm and 200 mm; the pillar width was 36 mm. The convergences of the pilot hole caused by the sidecoring drilling were recorded from the model (Figure A1.3) and the in situ stress was back calculated by using a best fit inverse solution equal to the one used in normal LVDT measurement interpretation (Figure A1.44). The applied horizontal in situ stress was 40 MPa and the vertical was 10 MPa. The elastic parameters were E = 55 GPa and = 0.25. The resulting in situ stress state was xx = 39.0 MPa and yy = 8.1 MPa with mean internal error of 3 %. The forced solution for the applied in situ stress gave the slightly higher mean error of 5 %. The estimate for the xx in situ stress component, which is in the direction of the side coring hole, is clearly better (-2.5 %) than for the yy (-19 %) component. Also, the estimate for the shear component xy was zero, as it should be. The simple example shows that sidecoring i.e., partial stress release, can be used to estimate or fulfil the information on the in situ state of stress. Conclusions of this study were:

in locations where core discing takes place during the overcoring, sidecoring could be used instead;

a sidecoring hole should be located so that it shadows the LVDT-installation in the direction of the maximum secondary stress on the excavation surface (Figure A1.5);

in the case of a Ø127 mm installation hole and Ø200 mm sidecoring hole, the pillar width could be around 40 mm;

over- and side-coring results can be used together in the best fit solution; but overcoring is the primary and preferred stress release method.

83

Figure A1.1. Maximum principal stress contours before (above) and after (below) sidecoring.

84

Figure A1.2. Minimum principal stress contours before (above) and after (below) sidecoring.

85

Figure A1.3. Displacement magnitudes caused by the 200 mm sidecoring, values in mm.

Figure A1.4. Best fit inverse solution for measured and calculated convergences (centre) and the internal solution errors for best fit solution and known in situ stress state.

86

Figure A1.5. Preferred location for the sidecoring hole related to the orientation of the maximum in situ compression.

87

Appendix 2

Example of LVDT measurement log sheet

(3 pages)

88

* Translations: kivi = rock, vesi = water

89

* Translations: huuhteluvesi = out-coming flushing water, menovesi = in-coming flushing water

90

* Translations: välikalibrointi = calibration between two measurements

91

Appendix 3

Review of Examine3D approach and Solve algorithm used

for LVDT-cell intepretation

Christer Andersson Vattenfall Power Consultant AB

March 2010

(5 pages)

92

1. Examine3D model Although the review of the Examine3D focuses on the LVDT measurements at the EDZ niche, the review is general and can be regarded as valid also for the convergence measurements in the various shafts. The meshing of the measurement holes and the tunnel close to the holes (Figure 1 and Figure 2) was analyzed and the uneven surface of the tunnel induces stress effects that potentially may affect results. If so, the mesh around the tunnel in the section containing the LVDTs could benefit from reduced element sizes. To verify the mesh, the elements at the measuring hole for LVDT#3, Figure 3, were reduced in size and a recalculation was made focusing on observing effects on the distance between the opposite LVDT heads. The exercise shows that for the Sxx unit stress tensor there is no effect and the original mesh is thus well designed. The resulting stress field between the LVDT’s 2 and 3 measurement holes were thereafter checked for the unit stress tensor Sxx and the resulting stress tensor (Figure 4). The objective was to check if the model had to be run in several stages, which is necessary if two holes are within the influence radius of each other. For the unit stress tensor, such effects cannot be observed. For the “real” stress tensor there is a very small effect, which is likely negligible.

Figure 1. Overview of the stress field close to the tunnel.

93

Figure 2. Overview of the stress field close to the tunnel and on the tunnel surface.

Figure 3. Close-up on the meshing of the measurement holes.

94

Figure 4. Stress field between LVDTs #2 and 3. Results from “real” stress tensor.

2. Excel SOLVE function The comparisons made for the deformations are based on the fact that the expansions measured in the hollow cylinder by the LVDTs are exactly the same but in opposite direction to the deformations recorded in the Examine 3D model. This is the conventional methodology for e.g. overcoring, but such tests are undertaken in more uniform stress. Hence, the assumption may to some extent be questioned. The SOLVE function in Excel has been checked by rotating the unit stress tensor 45 degrees clockwise horizontally. When comparing with the original unit stress tensor used the resulting magnitudes and orientations remain essentially the same. The stress tensor in Table 1 has been used in the Examine 3D model to determine the resulting displacements for LVDT #3 and to compare those with the actual measured displacements. The results presented in Table 2 indicate that the difference between the measured values and the calculated ones are, on average, 0.0135 mm or 15 % in absolute terms. The difference between the measured and calculated results is a combined effect of the stress field used and the, geometrical accuracy of the model and the quality of the meshing. All of these parameters need to be carefully managed to achieve good results.

95

Table 1. Stress tensor derived by Hakala.

Magnitude Dip Dip dir.

1 31.8 1 351

2 19.9 21 260

3 13.5 66 84

3. Discussion The Examin3D application presumes ideal linear elastic conditions and for these conditions, the back calculation of the stress tensor through the unit tensor and measured convergences gives results that seem accurate and credible. In more realistic conditions, however, there are multiple issues that affect the results. The most important involve:

EDZ, fractures in the rock mass, inhomogeneity in the rock mass, Young’s modulus of rock mass, in situ stress, model geometry and meshing, sensitivity of SOLVE function.

The EDZ is very important for the behavior of the rock close to a drill and blast tunnel. In the model, LVDT’s are located approximately 300 mm in the normal direction to the excavation surface. It is likely that the EDZ reaches this far in the rock mass and affects the deformation properties in a skin around the tunnel. Naturally occurring fractures affect the deformation properties of the rock, similarly to inhomogeneities. Because the measurements are local, LVDT’s can be individually affected by these parameters. It is therefore important to check the deformation pattern of each LVTD individually and two opposite LVDT’s in combination with the objective of finding irregularities. If this is the case, the measurement should be disregarded. The above discussed parameters affect the stiffness of the rock mass. The majority of the strain that causes the displacements in the numerical model is mobilised close to the measurement hole. This is also the part of the rock mass around the tunnel that is affected by the EDZ. The actual Young’s modulus around the measurement holes cannot be determined. Therefore, the stress tensor should also be calculated for reasonably high and low values of Young’s modulus, thus providing a range of possible tensors. An assessment of the total error in the determination of the stress field using this method should be done. The major errors relates to the accuracy in the back calculation of the convergences (exemplified in Table 2), and the error in the SOLVE function when trying to match the displacements resulting from the stress field with the measured displacements. Both of these errors are approximately 15 % of the measured convergence.

96

The model geometry needs to be very similar to the real excavation geometry because the convergence measurements are made so close to the tunnel. The secondary effects of the tunnel on the stress field are therefore an important factor. In the present case, the model is constructed from a laser scanning of tunnel sections and the resemblance between the model geometry and reality is therefore likely to be good enough. In summary, it can be said that the approach to calculate the in situ stress field is adequate. It is though necessary to validate the assumptions made regarding the elastic conditions of the rock mass close to the excavation before these results can be regarded as a best estimate of the in situ stress at the studied location. Table 2. Comparison of resulting displacement in the ENU system using the stress tensors in Table 1. Young’s modulus E as 55Gpa. The positive sign in the difference column means that the calculation indicates larger displacements than those measured.

LVDT head Measured convergence

(µm)

Calculated convergence

(µm)

Difference: Measured – Calculated

convergence (µm)

Difference in % of measured

1: 1-5 87  112  ‐25  ‐29 1: 2-6 53  81  ‐28  ‐52 1: 3-7 48  18  30  62 1: 4-8 59  50  8  14 2: 1-5 104  114  ‐10  ‐10 2: 2-6 105  99  6  6 2: 3-7 137  138  ‐1  ‐1 2: 4-8 143  153  ‐10  ‐7 3: 1-5 126  101  25  20 3: 2-6 131  110  21  16 3: 3-7 149  171  ‐22  ‐15 3: 4-8 170  162  7  4 4: 1-5 105  108  ‐3  ‐3 4: 2-6 102  102  1  1 4: 3-7 163  140  23  14 4: 4-8 151  146  5  3 5: 1-5 79  102  ‐23  ‐30 5: 2-6 72  74  ‐2  ‐2 5: 3-7 78  84  ‐6  ‐7 5: 4-8 90  112  ‐22  ‐24 8: 1-5 126  104  23  18 6: 2-8 81  71  10  12 8: 3-7 37  40  ‐3  ‐9 8: 4-8 84  72  12  15 

97

Appendix 4

LVDT Sensitivity study

98

Figure A4.1. Heating test 1: Diametrical deformations when rock cylinder is heated.

Figure A4.2. Heating test 1: Radial deformations when rock cylinder is heated.

99

Figure A4.3. Heating test 2: Diametrical deformations when LVDT-cell body is heated.

Figure A4.4. Heating test 2: Radial deformations when LVDT-cell body is heated.

100

Figure A4.5. Shake test 1: Diametrical deformations when LVDT-cell inside rock cylinder.

Figure A4.6. Shake test 1: Radial deformations when LVDT-cell inside rock cylinder.

101

Figure A4.7. Shake test 2: Diametric deformations when LVDT-cell inside steel cylinder.

Figure A4.8. Shake test 2: Radial deformations when LVDT-cell inside steel cylinder.

102

Figure A4.9. Roll test 1: Diametrical deformations when LVDT-cell inside rock cylinder.

Figure A4.10. Roll test 1: Radial deformations when LVDT-cell inside rock cylinder.

103

Figure A4.11. Roll test 2: Diametrical deformations when LVDT-cell inside steel cylinder.

Figure A.3.12. Roll test 2: Radial deformations when LVDT-cell inside steel cylinder.

104

Figure A4.13. Impact test 1: Diametrical deformations when LVDT-cell inside rock cylinder.

Figure A4.14. Impact test 1: Radial deformations when LVDT-cell inside rock cylinder.

105

Figure A4.15. Impact test 2: Diametrical deformations when LVDT-cell inside steel cylinder.

Figure A4.16. Impact test 2: Radial deformations when LVDT-cell inside steel cylinder.

106

107

Appendix 5

Unisigma P09POS11_Rep_01

Strain gauge measurements in ONKALO exhaust air shaft level -265 and biaxial testing of the pilot cores

108

109

110

111

112

113

114

115

116

117

118

119

120

121

122

123

124

125

126

127

128

129

130

131

132

133

134

135

136

137

Appendix 6

LVDT-cell measurements in ONLKALO exhaust air shaft level -265

138

Figure A6-1. LVDT measurement in ONKALO exhaust air shaft level -265 m,

Measurement R1.

Figure A6-2. LVDT measurement in ONKALO exhaust air shaft level -265 m,

Measurement R2.

139

Figure A6-3. LVDT measurement in ONKALO exhaust air shaft level -265 m,

Measurement R3.

Figure A6-4. LVDT measurement in ONKALO exhaust air shaft level -265 m,

Measurement R4.

140

Figure A6-5. LVDT measurement in ONKALO exhaust air shaft level -265 m,

Measurement R5.

Figure A6-6. LVDT measurement in ONKALO exhaust air shaft level -265 m,

Measurement R6.

141

Table A6-1. LVDT measurement in ONKALO exhaust air shaft level -265 m, Summary of local stress solution.

Measurement location R1 R2 R3 R4 R5 R6 Measured convergences vertical, m 3-7 159 141 92 104 131 187 cw 45*, m 4-8 153 87 62 48 56 119 horizontal, m 1-5 51 11 22 28 20 23 cw 135*, m 2-6 79 68 44 89 100 154 *) clockwise angle when looking toward shaft wall Resulting stress state 1, MPa 56.5 33.5 24.5 30.5 29.5 39.5 3, MPa 32 13.5 12.5 15.5 13 20.5 alfa* (degrees) -14.5 -4.5 -9.5 13.5 10 13 *) cw from vertical Error between measured and resulting convergences vertical 3-7 1 % 1 % -1 % 3 % 0 % 1 % cw 45 * 4-8 -11 % 0 % 9 % -1 % -1 % -37 % horizontal 1-5 9 % 13 % -2 % 2 % 2 % 57 % cw 135 * 2-6 0 % -1 % 1 % -1 % -4 % 0 % *) clockwise angle when looking toward shaft wall Cumulative error between measured and calculated convergences 2 % 0 % 1 % 1 % 1 % 6 %

142

143

Appendix 7

LVDT-cell measurements in Äspö TASS-tunnel

144

Figure A7.1. Äspö TASS-tunnel, LVDT overcoring in hole R1 at depth of 120 mm.

Figure A7.2. Äspö TASS-tunnel, LVDT overcoring in hole R1 at depth of 630 mm.

145

Figure A7.3. Äspö TASS-tunnel, LVDT sidecoring in hole R2 at depth of 100 mm.

Figure A7.4. Äspö TASS-tunnel, LVDT sidecoring in hole R2 at depth of 500 mm.

146

Figure A7.5. Äspö TASS-tunnel, LVDT sidecoring in hole R3 at depth of 100 mm.

Figure A7.6. Äspö TASS-tunnel, LVDT sidecoring in hole R3 at depth of 630 mm.

147

Figure A7.7. Äspö TASS-tunnel, LVDT sidecoring in hole R4 at depth of 150 mm.

Figure A7.8. Äspö TASS-tunnel, LVDT sidecoring in hole R4 at depth of 460 mm.

148

Figure A7.9. Äspö TASS-tunnel, LVDT overcoring in hole R5 at depth of 100 mm.

Figure A7.10. Äspö TASS-tunnel, LVDT overcoring in hole R5 at depth of 500 mm.

149

Appendix 8

LVDT-cell measurements in Äspö TBM-tunnel

150

Figure A8.1. Äspö TBM-tunnel, LVDT overcoring in hole R1 at depth of 350 mm.

Figure A8.2. Äspö TBM-tunnel, LVDT sidecoring in hole R2 at depth of 240 mm.

151

Figure A8.3. Äspö TBM-tunnel, LVDT sidecoring in hole R2 at depth of 780 mm.

Figure A8.4. Äspö TBM-tunnel, LVDT sidecoring in hole R3 at depth of 350 mm.

OC_Start

10 cm

20 cm

30 cm

SC_End, 33 cm

0

4

8

12

16

20

24

28

32

36

40

44

48

52

-0.08

-0.06

-0.04

-0.02

0

0.02

0.04

0.06

0.08

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0.14

0.16

0.18

0.2

0.22

14:00 14:15 14:30 14:45 15:00 15:15 15:30 15:45 16:00 16:15

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mp

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(C)

Dia

met

ric

de

form

ati

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(mm

)

M2 - 90 (1+5)

M2 - 135 (2+6)

M2 - 00 (3+7)

M2 - 45 (4+8)

OC_Start

SC_End, 33 cm

Values for calc.

Breaks

T, Probe

T, Rock

Looking from tunnel tothe measurement hole

OC_Start 10 cm

0.19

20 cm 30 cm OC_End, 35 cm

0

4

8

12

16

20

24

28

32

36

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48

52

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0.14

0.16

0.18

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0.24

2:45 3:00 3:15 3:30 3:45 4:00 4:15 4:30

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re (

C)

Dia

me

tric

def

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atio

n (

mm

)

M3 - 90 (1+5)

M3 - 135 (2+6)

M3 - 00 (3+7)

M3 - 45 (4+8)

OC_Start

OC_End, 35 cm

Values for calc.

T, Probe

T, Rock

Looking from tunnel tothe measurement hole

152

Figure A8.5. Äspö TBM-tunnel, LVDT sidecoring in hole R3 at depth of 740 mm.

Figure A8.6. Äspö TBM-tunnel, LVDT overcoring in hole R4 at depth of 70 mm.

OC_Start

10 cm

0.19

20 cm

30 cm

OC_End, 35 cm

0

4

8

12

16

20

24

28

32

36

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19:45 20:00 20:15 20:30 20:45 21:00 21:15

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C)

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tric

def

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atio

n (

mm

)

M4 - 90 (1+5)

M4 - 135 (2+6)

M4 - 00 (3+7)

M4 - 45 (4+8)

OC_Start

OC_End, 35 cm

Values for calc.

T, Probe

T, Rock

Looking from tunnel tothe measurement hole

153

Figure A8.7. Äspö TBM-tunnel, LVDT sidecoring in hole R6 at depth of 240 mm.

Figure A8.8. Äspö TBM-tunnel, LVDT overcoring in hole R7 at depth of 60 mm.

OC_Start

10 cm

20 cm

30 cm

OC_End, 50 cm

0

4

8

12

16

20

24

28

32

36

40

44

48

52

-0.08

-0.06

-0.04

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0

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0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

0.2

0.22

18:30 18:45 19:00 19:15 19:30 19:45 20:00 20:15 20:30 20:45 21:00

Te

mp

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(C

)

Dia

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ric

def

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atio

n (

mm

)

M6 - 90 (1+5)

M6 - 135 (2+6)

M6 - 00 (3+7)

M6 - 45 (4+8)

OC_Start

OC_End, 50 cm

Values for calc.

T, Probe

T, Rock

Looking from tunnel tothe measurement hole

OC_Start 10 cm

0.19

20 cm 30 cm OC_End, 35 cm

0

4

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12

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0.26

20:45 21:00 21:15 21:30 21:45 22:00

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C)

Dia

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form

atio

n (

mm

)

M7 - 90 (1+5)

M7 - 135 (2+6)

M7 - 00 (3+7)

M7 - 45 (4+8)

OC_Start

OC_End, 35 cm

Values for calc.

T, Probe

T, Rock

Looking from tunnel tothe measurement hole

154

Figure A8.9. Äspö TBM-tunnel, LVDT sidecoring in hole R9 at depth of 120 mm, note that LVDT number one is aligned with tunnel axis but upwards in all other measurements.

Figure A8.10. Äspö TBM-tunnel, LVDT sidecoring in hole R10 at depth of 120 mm.

155

Figure A8.11. Äspö TBM-tunnel, LVDT sidecoring in hole R10 at depth of 485 mm.

OC_Start

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SC_End, 35 cm

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16:30 16:45 17:00 17:15 17:30 17:45 18:00 18:15

Te

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(C)

Dia

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form

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m)

M10 - 90 (1+5)

M10 - 135 (2+6)

M10 - 00 (3+7)

M10 - 45 (4+8)

OC_Start

SC_End, 35 cm

Values for calc.

Breaks

T, Probe

T, Rock

Looking from tunnel tothe measurement hole

156

157

Appendix 9

Biaxial test results in Äspö TASS-tunnel

158

159

160

161

162

163

164

165

Appendix 10

Biaxial test results in Äspö TBM-tunnel

166

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179

180

LIST OF REPORTS

POSIVA-REPORTS 2012

_______________________________________________________________________________________

POSIVA 2012-01 Monitoring at Olkiluoto – a Programme for the Period Before Repository Operation Posiva Oy ISBN 978-951-652-182-7 POSIVA 2012-02 Microstructure, Porosity and Mineralogy Around Fractures in Olkiluoto

Bedrock Jukka Kuva (ed.), Markko Myllys, Jussi Timonen, University of Jyväskylä Maarit Kelokaski, Marja Siitari-Kauppi, Jussi Ikonen, University of Helsinki Antero Lindberg, Geological Survey of Finland Ismo Aaltonen, Posiva Oy ISBN 978-951-652-183-4

POSIVA 2012-03  Safety Case for the Disposal of Spent Nuclear Fuel at Olkiluoto - Design Basis 2012 Posiva Oy  ISBN 978-951-652-184-1 POSIVA 2012-04 Safety Case for the Disposal of Spent Nuclear Fuel at Olkiluoto - Performance Assessment 2012 Posiva Oy ISBN 978-951-652-185-8 POSIVA 2012-05 Safety Case for the Disposal of Spent Nuclear Fuel at Olkiluoto - Description of the Disposal System 2012 Posiva Oy ISBN 978-951-652-186-5 POSIVA 2012-06 Olkiluoto Biosphere Description 2012 Posiva Oy ISBN 978-951-652-187-2 POSIVA 2012-07 Safety Case for the Disposal of Spent Nuclear Fuel at Olkiluoto - Features, Events and Processes 2012 Posiva Oy   ISBN 978-951-652-188-9  POSIVA 2012-08 Safety Case for the Disposal of Spent Nuclear Fuel at Olkiluoto - Formulation of Radionuclide Release Scenarios 2012 Posiva Oy ISBN 978-951-652-189-6

POSIVA 2012-09 Safety Case for the Disposal of Spent Nuclear Fuel at Olkiluoto - Assessment of Radionuclide Release Scenarios for the Repository System 2012 Posiva Oy ISBN 978-951-652-190-2 POSIVA 2012-10 Safety case for the Spent Nuclear Fuel Disposal at Olkiluoto - Biosphere Assessment BSA-2012 Posiva Oy ISBN 978-951-652-191-9 POSIVA 2012-11 Safety Case for the Disposal of Spent Nuclear Fuel at Olkiluoto - Complementary Considerations 2012 Posiva Oy ISBN 978-951-652-192-6 POSIVA 2012-12 Safety Case for the Disposal of Spent Nuclear Fuel at Olkiluoto - Synthesis 2012 Posiva Oy ISBN 978-951-652-193-3 POSIVA 2012-13 Canister Design 2012 Heikki Raiko, VTT ISBN 978-951-652-194-0 POSIVA 2012-14 Buffer Design 2012 Markku Juvankoski, VTT ISBN 978-951-652-195-7 POSIVA 2012-15 Backfill Design 2012 Posiva Oy ISBN 978-951-652-196-4 POSIVA 2012-16 Canister Production Line 2012 – Design, Production and Initial State of the Canister Heikki Raiko (ed.), VTT Barbara Pastina, Saanio & Riekkola Oy Tiina Jalonen, Leena Nolvi, Jorma Pitkänen & Timo Salonen, Posiva Oy ISBN 978-951-652-197-1 POSIVA 2012-17 Buffer Production Line 2012 – Design, Production, and Initial State of the Buffer Markku Juvankoski, Kari Ikonen, VTT Tiina Jalonen, Posiva Oy ISBN 978-951-652-198-8

POSIVA 2012-18 Backfill Production Line 2012 - Design, Production and Initial State of the Deposition Tunnel Backfill and Plug Paula Keto (ed.), Md. Mamunul Hassan, Petriikka Karttunen, Leena Kiviranta, Sirpa Kumpulainen, B+Tech Oy Leena Korkiala-Tanttu, Aalto University Ville Koskinen, Fortum Oyj Tiina Jalonen, Petri Koho, Posiva Oy Ursula Sievänen, Saanio & Riekkola Oy ISBN 978-951-652-199-5 POSIVA 2012-19 Closure Production Line 2012 - Design, Production and Initial State of Underground Disposal Facility Closure Ursula Sievänen & Taina H. Karvonen, Saanio & Riekkola Oy David Dixon, AECL Johanna Hansen & Tiina Jalonen, Posiva Oy ISBN 978-951-652-200-8 POSIVA 2012-20 Representing Solute Transport Through the Multi-Barrier Disposal System by Simplified Concepts Antti Poteri. Henrik Nordman & Veli-Matti Pulkkanen, VTT Aimo Hautojärvi, Posiva Oy Pekka Kekäläinen, University of Jyväskylä, Deparment of Physics ISBN 978-951-652-201-5 POSIVA 2012-21 Layout Determining Features, their Influence Zones and Respect Distances at the Olkiluoto Site Tuomas Pere (ed.), Susanna Aro & Jussi Mattila, Posiva Oy Henry Ahokas & Tiina Vaittinen, Pöyry Finland Oy Liisa Wikström, Svensk Kärnbränslehantering AB ISBN 978-951-652-202-2 POSIVA 2012-22 Underground Openings Production Line 2012 – Design, Production and Initial State of the Underground Openings Posiva Oy ISBN 978-951-652-203-9 POSIVA 2012-23 Site Engineering Report ISBN 978-951-652-204-6 POSIVA 2012-24 Rock Suitability Classification, RSC-2012 Tim McEwen (ed.), McEwen Consulting Susanna Aro, Paula Kosunen, Jussi Mattila & Tuomas Pere, Posiva Oy Asko Käpyaho, Geological Survey of Finland Pirjo Hellä, Saanio & Riekkola Oy ISBN 978-951-652-205-3 POSIVA 2012-25 2D and 3D Finite Element Analysis of Buffer-Backfill Interaction Martino Leoni, Wesi Geotecnica Srl ISBN 978-951-652-206-0

POSIVA 2012-26 Climate and Sea Level Scenarios for Olkiluoto for the Next 10,000 Years Natalia Pimenoff, Ari Venäläinen & Heikki Järvinen, Ilmatieteen laitos ISBN 978-951-652-207-7 POSIVA 2012-27 Geological Discrete Fracture Network Model for the Olkiluoto Site, Eurajoki, Finland: version 2.0 Aaron Fox, Kim Forchhammer & Anders Pettersson, Golder Associates AB Paul La Pointe & Doo-Hyun Lim, Golder Associates Inc. ISBN 978-951-652-208-4 POSIVA 2012-28 Safety Case for the Disposal of Spent Nuclear Fuel at Olkiluoto - Data Basis for the Biosphere Assessment BSA-2012 Posiva Oy      ISBN 978-951-652-209-1 POSIVA 2012-29 Safety Case For The Disposal of Spent Nuclear Fuel at Olkiluoto - Terrain and Ecosystems Development Modelling in the Biosphere Assessment BSA-2012 Posiva Oy ISBN 978-951-652-210-7 POSIVA 2012-30 Safety Case for the Disposal of Spent Nuclear Fuel at Olkiluoto - Surface and Near-surface Hydrological Modelling in the Biosphere Assessment BSA-2012 Posiva Oy ISBN 978-951-652-211-4 POSIVA 2012-31 Safety Case for the Disposal of Spent Nuclear Fuel at Olkiluoto - Radionuclide Transport and Dose Assessment for Humans in the Biosphere Assessment BSA-2012 Posiva Oy ISBN 978-951-652-212-1 POSIVA 2012-32 Safety Case for the Disposal of Spent Nuclear Fuel at Olkiluoto - Dose Assessment for the Plants and Animals in the Biosphere Assessment BSA-2012 Posiva Oy ISBN 978-951-652-213-8 POSIVA 2012-33 Underground Openings Line Demonstrations Stage 1, 2012 ISBN 978-951-652-214-5 POSIVA 2012-34 Seismic Activity Parameters of the Olkiluoto Site Jouni Saari, ÅF-Consult Oy ISBN 978-951-652-215-2

POSIVA 2012-35 Inspection of Disposal Canisters Components Jorma Pitkänen, Posiva Oy ISBN 978-951-652-216-9 POSIVA 2012-36 Analyses of Disposal Canister Falling Accidents Juha Kuutti, Ilkka Hakola & Stephania Fortino, VTT ISBN 978-951-652-217-6 POSIVA 2012-37 Long-Term Safety of the Maintenance and Decommissioning Waste of the Encapsulation Plant Olli Nummi, Jarkko Kyllönen & Tapani Eurajoki, Fortum Power and Heat Oy ISBN 978-951-652-224-4 POSIVA 2012-38 Human Factors in NDT of the EB-Weld ISBN 978-951-652-225-1 POSIVA 2012-39 Safety Case for the Disposal of Spent Nuclear Fuel at Olkiluoto: Radionuclide Solubility Limits and Migration Parameters for the Canister and the Buffer. Wersin, P., Kiczka, M. & Rosch, D.. ISBN 978-951-652-219-0 POSIVA 2012-40 Safety Case for the Disposal of Spent Nuclear Fuel at Olkiluoto: Radionuclide Solubility Limits and Migration Parameters for the Backfill. Wersin, P., Kiczka, M., Rosch, D., Ochs, M., Trudel, D., ISBN 978-951-652-220-6 POSIVA 2012-41 Safety Case for the Disposal of Spent Nuclear Fuel at Olkiluoto: Radionuclide Migration Parameters for the Geosphere. Martti Hakanen, Heini Ervanne & Esa Puukko ISBN 978-951-652-221-3 POSIVA 2012-42 Summary Report. Microbiology of Olkiluoto and ONKALO Groundwater Karsten Pedersen, Microbial Analytics Sweden Ab Malin Bomberg & Merja Itävaara, VTT ISBN 978-951-652-222-0 POSIVA 2012-43 In Situ Stress Measurement with LVDT-cell – Method Description and Verification. Matti Hakala, KMS Hakala Oy Topias Siren & Kimmo Kemppainen, Posiva Oy Rolf Christiansson, SKB Derek Martin, University Of Alberta ISBN 978-951-652-223-7