effect of overcoring induced heat on lvdt stress ... - … · working report 2012-75 ... lvdt...

June 2013

Working Reports contain information on work in progress

or pending completion.

Tuomas Rantanen

WSP Finland Oy

Working Report 2012-75

Effect of Overcoring Induced Heaton LVDT Stress Measurements

EFFECT OF OVERCORING INDUCED HEAT ON LVDT STRESS MEASUREMENTS

ABSTRACT

A new overcoring based rock stress measuring device called LVDT-cell has been developed for Posiva Oy. LVDT-cell measures radial deformation of a rock cylinder due to release of stresses as it is overcored.

Overcoring based rock stress measuring methods are susceptible to temperature effects. This work is focused on measuring the temperature field during and after the overcoring and evaluating the magnitude of radial displacement caused by thermal expansion of the overcored rock cylinder using numerical simulation.

Total of seven overcoring tests were done in veined gneiss (VGN) and pegmatitic granite (PGR) at variable depths from tunnel wall. Measured maximum temperature increase on the pilot hole wall varied from 3.7 °C (VGN) to 14.5 °C (PGR). Overcored rock cylinder cooled to ambient rock mass temperature in six hours.

Thermal simulation was used to develop a transient heat model of the overcoring process. This model was used in coupled simulation to evaluate the magnitude of radial displacement of the rock cylinder during overcoring.

Results of the numerical simulations indicate maximum radial displacement of 3.57 µm for overcoring performed in VGN and 9.84 µm for overcoring performed in PGR. Difference is explained by larger heat production in PGR. According to simulations, it is suggested to not take final readings from LVDT-cell until after four hours of cooling time if temperature compensation is not used.

It is also concluded that temperature and displacement data from the LVDT cell sensors could be used to compensate for the temperature effects.

Keywords: LVDT, overcoring, temperature effect, rock stress.

IRTIKAIRAUKSEN AIHEUTTAMAN LÄMPÖTILAN VAIKUTUS LVDT JÄNNITYSTILAMITTAUKSIIN

TIIVISTELMÄ

ONKALOn jännitystilamittauksia varten on Posiva Oy:lle kehitetty uusi irtikairaukseen perustuvan kallion jännitystilamittauslaite, LVDT-kenno. LVDT–kenno mittaa kivi-sylinterin radiaalisia muodonmuutoksia kun siinä olevat jännitykset purkautuvat irti-kairauksen seurauksena.

Irtikairausmenetelmät ovat alttiita kairauksen yhteydessä syntyville lämpövaikutuksille. Tässä työssä on keskitytty mittaamaan irtikairauksen aiheuttama lämpötilakenttä kivi-sylinterissä, sekä arvioitu lämpölaajenemisen aiheuttamia radiaalisiirtymiä numeerisen mallinnuksen avulla.

Työssä tehtiin seitsemän irtikairauskoetta kahdessa eri kivilajissa vaihtelevilla syvyyk-sillä. Testatut kivilajit olivat juonigneissi (VGN) ja pegmatiittinen graniitti (PGR). Mitatut lämpötilannousut vaihtelivat maksimissaan välillä 3.7 °C (VGN) – 14.5 °C (PGR). Molemmissa kivilajeissa kivisylinterin jäähtyminen ympäristön lämpötilaan kesti noin kuusi tuntia irtikairauksen suorittamisesta.

Numeerinen mallinnus jaettiin lämpömallinnukseen ja yhdistettyyn lämpö- ja mekaa-niseen mallinnukseen. Lämpömallinnuksen avulla luotiin ylikairausta vastaava ajasta riippuva lämpötilamalli. Tätä mallia käytettiin yhdistetyssä mallinnuksessa irtikairauk-sen aiheuttamien lämpötilasta riippuvien radiaalisiirtymien suuruuden arviointiin.

Mallinnuksen tulokset osoittavat radiaalisiirtymän olevan maksimissaan 3.57 µm juoni-gneississä suoritetussa irtikairauksessa, sekä 9.84 µm graniitissa suoritetussa irtikai-rauksessa. Tulosten perusteella on suositeltavaa mitata LVDT-kennolla vähintään neljä tuntia irtikairauksen jälkeen jotta voidaan varmistua tulosten olevan riippumattomia lämpötilavaikutuksista.

Lisäksi todettiin, että LVDT-kennolla mitatun lämpötila- ja siirtymätiedon avulla voi-daan mahdollisesti kompensoida irtikairauksen aiheuttamia lämpötilavaikutuksia.

Avainsanat: LVDT, irtikairaus, lämpövaikutus, jännitystilamittaus.

PREFACE

This work is done as authors master’s thesis and it is a part of the development of the LVDT cell. The author would like to thank Topias Siren from Posiva Oy and Matti Hakala for the research subject and guidance along the way. Additional thanks go to Teemu Koskinen from STIPS Oy for designing and manufacturing temperature measuring device and Professor Mikael Rinne for comprehensive review of this work.

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TABLE OF CONTENTS

ABSTRACT

TIIVISTELMÄ

1 INTRODUCTION ..................................................................................................... 5

1.1 Scope and objectives of this study ................................................................... 5 1.2 The LVDT rock stress measurement cell ......................................................... 5 1.3 Previous studies ............................................................................................... 7

2 OVERCORING EXPERIMENT .............................................................................. 11

2.1 Motivation ....................................................................................................... 11 2.2 Arrangement ................................................................................................... 11 2.3 Measuring equipment ..................................................................................... 14

2.3.1 A device to measure hollow core surface temperature field ................... 14 2.3.2 Sensor tubes ........................................................................................... 18

2.4 Error estimation .............................................................................................. 18 2.4.1 Pilot hole temperature cell ....................................................................... 18 2.4.2 Sensor tubes ........................................................................................... 19

3 RESULTS .............................................................................................................. 21

3.1 Sensor naming conventions ........................................................................... 21 3.2 Test no. 1 – in VGN at 50 mm depth .............................................................. 21 3.3 Test no. 2 – in VGN at 50 mm depth .............................................................. 25 3.4 Test no. 3 – in VGN at 450 mm depth ............................................................ 31 3.5 Test no. 4 – in PGR at 50 mm depth .............................................................. 33 3.6 Test no. 5 – in PGR at 300 mm depth ............................................................ 38 3.7 Test no. 6 – in VGN at 50 mm depth .............................................................. 40 3.8 Test no. 7 – in VGN at 450 mm depth ............................................................ 42

4 NUMERICAL MODELLING ................................................................................... 45

4.1 Modelling paradigm ........................................................................................ 45 4.2 Geometry and material properties .................................................................. 46 4.3 Transient heat transfer model ........................................................................ 49

4.3.1 Boundary conditions ................................................................................ 49 4.3.2 Results .................................................................................................... 55 4.3.3 Sensitivity analysis .................................................................................. 59

4.4 Coupled thermo-mechanical model ................................................................ 60 4.4.1 Boundary conditions ................................................................................ 60 4.4.2 Results .................................................................................................... 61 4.4.3 Sensitivity analysis .................................................................................. 65

5 SUMMARY ............................................................................................................ 67

5.1 Summary of overcoring tests .......................................................................... 67 5.2 Summary of numerical simulations ................................................................ 68

6 CONCLUSIONS AND RECOMMENDATIONS ..................................................... 71

REFERENCES ............................................................................................................. 73

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ABBREVIATIONS

AD Analog-to-digital (conversion)

CSIRO Commonwealth Scientific and Industrial Research Organisation, Australia's national science agency.

EDZ Excavation Damaged Zone

IC Integrated circuit

LVDT Linear Variable Differential Transformer

NTC Negative Temperature Coefficient

OC Overcoring

PGR Pegmatitic granite

POSE Posiva's Olkiluoto Spalling Experiment

TBM Tunnel Boring Machine

USBM United States Bureau of Mines

VGN Veined gneiss

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1 INTRODUCTION

1.1 Scope and objectives of this study

ONKALO is an underground rock characterisation facility constructed by Posiva Oy, which is the nuclear waste management organisation in Finland. ONKALO is constructed at a depth of between 400 and 600 m in the crystalline bedrock in the Olkiluoto Island, Eurajoki. Olkiluoto is an island in the Baltic Sea coast close to mainland, covering about 10 km2 (Posiva Oy, 2009).

A key aspect in underground rock characterization is knowledge of the current in-situ stress state of the rock mass in the vicinity of the ONKALO. Excavations also change the stress field and these changes should be measured and modelled as well as possible to ensure long-term safety of the ONKALO tunnels.

Overcoring based methods to measure in-situ stress field in rock have already been available for a few decades. Common variants are the Borre Probe and the CSIRO stress cells, although many other variants are also used around the world.

With traditionally used overcoring based stress cells, such as Borre or CSIRO, it has long been know that the overcoring induced heat might cause significant inaccuracy and possibly bias the stress measurement data.

Temperature effects during an overcoring measurement can be divided roughly to be device or medium based. Device based temperature effects include all uncertainties caused by temperature variation in the measuring device, for example heating of strain gauges or lead wires. Medium based effects include thermal expansion and/or contraction of measured medium and possibly other temperature variable material properties.

The sensors in these traditionally used devices, however, differ significantly from the new LVDT cell that was developed especially to be used in ONKALO. As the LVDT cell is dimensionally larger and uses electro-mechanical sensor technique, it is by nature less prone to temperature induced inaccuracy.

This study is part of the development of the stress measurement techniques and the purpose of this study is to

a) Measure the temperature field in the rock cylinder during overcoring and b) evaluate the maximum effect of measured temperature variations to the

displacements measured by the LVDT rock stress measurement cell in ONKALO rock conditions

1.2 The LVDT rock stress measurement cell

Numerous stress measurement campaigns have been conducted in the ONKALO area using various different stress measurement methods, including hydraulic fracturing and overcoring with CSIRO cell and Borre probe (Posiva Oy, 2009). These campaigns raised a need for a new method for stress measurement due to the scale problems with foliation and heterogeneity of migmatite in Olkiluoto. Traditional overcoring methods

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also suffered from gluing problems, which encouraged the development of an electromechanical device to measure displacements. To address these issues, a new stress measurement device was developed for Posiva Oy to measure the stresses around the boundary of a tunnel surface. The new LVDT (Linear Variable Differential Transformer) cell and its components are presented in Figure 1.

LVDT stress measurement cell measures two dimensional diametric convergences induced by full or partial release of stresses. This stress release is achieved by either overcoring the LVDT cell, or by core drilling a large hole close to the LVDT cell. Latter method is also referred as sidecoring (see Figure 2). Stress state can be calculated by performing at least three, preferably four to five, measurements around tunnel perimeter. These stress measurements together with scanned tunnel geometry are used as parameters for numerical back calculation, which in turn returns full in-situ stress tensor. Elastic properties for the rock are determined from the pilot hole core samples with biaxial testing (Hakala et al. 2012).

In excavations made with drill and blast method, a buffer zone of 25 – 50 cm is used to avoid most of the excavation damaged zone (EDZ). In raise bored shafts and TBM tunnels the LVDT cell can be installed at the surface of the excavation. Required dimensions and installation instructions for LVDT cell are presented in Figure 3.

The LVDT cell has been tested successfully in Äspö hard rock laboratory, where in-situ stress state is relatively well-known. In-situ stress state measured by LVDT cell agrees very well with previous measurements performed with conventional methods (Hakala et al. 2012).

Figure 1. Version 2 of the LVDT stress measurement cell (Hakala et al. 2012).

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Figure 2. LVDT-cell installed in pilot hole and sidecoring hole in Äspö TBM-tunnel (Hakala et al. 2012).

Figure 3. LVDT cell installation (Hakala et al. 2012).

1.3 Previous studies

Temperature effects on overcoring stress measurement are not very widely studied. The most comprehensive study was made by Bertilsson (Temperature effects in overcoring stress measurements, 2007) as master’s thesis at Luleå University of Technology. The study is focused to the Borre Probe stress measurement cell and comprise of numerical simulations and various laboratory tests to identify the temperature effects of different components, namely the strain gages, glue and the probe itself.

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Bertilsson used numerical simulation to estimate the temperatures right below the drill bit and at the location of the strain gauges. The results show a peak temperature of about 20 °C above ambient temperature at the strain gauges, depending on the thermal conductivity of the rock mass.

Bertilsson also did laboratory tests with Borre Probe installed to a rock cylinder, which was then heated in controlled manner. In the test the thermal response and behaviour of the rock was not as expected, as the rock seemingly contracted under 30 °C temperatures and started to expand after temperature exceeded 30 °C. Bertilsson did not found clear explanation for this behaviour. However Lin et al. (2006) found out that depending on strain gauge, temperature compensation can be around -10 microstrains/°C, which can affect measurements considerably and thus must be taken into account.

Ask (Evaluation of measurement-related uncertainties in the analysis of overcoring rock stress data from Äspö HRL, Sweden: a case study, 2003) has evaluated the temperature effects in overcoring test performed in Äspö hard rock laboratory as part of his research. The temperature correction was estimated to be -8 microstrains/°C, which is about 1 MPa increase in magnitude on average. Corrections were only done for tests where temperature difference of values before and after overcoring were over 1°C. Smaller temperature differences were considered to have negligible effect to results.

Ask states that the maximum temperature difference measured was 4.3 °C, which corresponds to stress magnitude increase of 2.5 – 3.0 MPa, or nearly 10 % increase in σH magnitude, 50 % increase in σh and 24 % increase in σv magnitude. This test was done at 379 m depth with Borre Probe.

In the same paper Ask also reminds that the temperature measurement is not done at the location of strain gauges, but at the logging device, which is exposed to flushing water. It is therefore possible that the actual temperatures at the rock surface are greater than those measured. It is also unclear whether these temperature effects were related to the thermal response of the strain gauges or the thermal response of the rock annulus.

The ISRM suggested methods for overcoring stress measurements (Sjöberg et al. 2003) states that an overcoring probe should measure the temperature in borehole to assess the temperature effects, and that after overcoring the borehole should be left with no on-going activity for at least 5-10 minutes. Collecting strain data during this phase enables to assess possible temperature effects. There are no direct guidelines suggested in the paper to how to acknowledge temperature effects while interpreting the results, other than a brief note that large deviations in readings might be a result of temperature effects.

Cai et al. (Study and Test of Techniques for Increasing Overcoring Stress Measurement Accuracy, 1995) have also studied various parameters affecting overcoring measurements. In their study they claim that the strain values induced by temperature changes could not be ignored and compensation for them is necessary. The temperature problems considered in that paper are however more concentrated to strain gauges and accurate resistance measurement, not the thermal expansion of the rock cylinder. As the LVDT cell is measuring deformations electro-mechanically and calibrated after every measurement, it does not suffer from the same amount of temperature related problems

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as variable resistance strain gauges, although it is not completely tolerant to thermal variations.

The stability of the LVDT cell has been tested with various different methods, using both mechanical and thermal stress testing. Mechanical testing includes tests for dynamic stability by rolling the cell in a rock cylinder and shaking the cell. In thermal stress test the LVDT cell was installed in a rock cylinder, which was pulse heated with a heat gun. Based on these tests, the maximum effect for a 10 °C temperature increase was 0.02 mm (see Figure 4). Temperatures were measured with LVDT cells two built-in sensors, from which one measures the temperature of the electronics, and another one is a spring-loaded thermistor which measures the rock surface temperature (Hakala et al. 2012)

Overcoring temperature has been measured in a smaller scale in ONKALO by Matti Hakala (unpublished). In this test, a temperature measuring device was installed in a small diameter drill hole, which was then overcored using different overcoring speeds and different drill bit diameters. Some of the tests had problems with flushing water leaking to sensors, which was later solved by installing 10 cm thick insulating plastic foam to separate the instrument from the flushing water. Some results of these experiments can be seen in Figure 5 and Figure 6. Measured temperature difference was very small, being well under 1 °C.

Figure 4. Diametric responses of LVDT-probe when it was inside a rock cylinder which was pulse heated by heat gun (Hakala et al. 2012).

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Figure 5. Fast overcoring with 146 mm drill bit.

Figure 6. Fast overcoring with 76 mm drill bit.

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2 OVERCORING EXPERIMENT

2.1 Motivation

Calibration of numerical models and laboratory heat stress tests depend on as accurate as possible knowledge of the in-situ temperature field in the hollow core during overcoring. In theory the amount of energy consumed while drilling could be measured and by knowing the mechanical properties of the rock type could dissipated thermal energy be calculated. Accurate modelling of the overcoring induced heat is however practically impossible due to numerous variables. Thermal properties of the rock, drilling speed, type and quality of drill bit, flushing water flow and temperature, are only few variables involved in the process.

As numerical models could over or under estimate the magnitude of the temperature field by decade, the in-situ temperature measurements are only practical approach to collect reliable data. Thus one of the main emphasises in this study was to collect in-situ data from the temperature field changes during overcoring process.

2.2 Arrangement

The experiment was performed in the Posiva’s ONKALO, in the POSE investigation niche. The approximate location of the experiment holes on the east wall of the POSE drift is presented in Figure 7. The POSE drift was selected for two reasons: first, there is no on-going activity, so measurements can be done in relatively undisturbed area. Second, the thermal properties of the rock types in POSE drift are well studied.

Figure 7. Location of the POSE investigation niche (Siren, Fracture Mechanics Prediction for Posiva's Olkiluoto Spalling Experiment (POSE), 2011).

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Temperature measuring was done in total of seven separate experiments to study different conditions. Five of the tests were done in veined gneiss (VGN) and two in pegmatitic granite (PGR). The rock type selection is based on the fact the pegmatite is much harder rock. The overcoring generates more heat in pegmatite, and thus it is interesting host rock for the experiment. Downsides for overcoring in pegmatite are the longer time the overcoring process takes and wearing of the drill bit. PGR is also more susceptible to core disking, which adds uncertainties and risks to the test. For these reasons it is decided to do most of the tests in VGN.

The following five setups were chosen for testing:

Overcoring in VGN near tunnel surface Overcoring in VGN at 500 mm depth from tunnel surface Overcoring in VGN after at least 12 h cooling time after pilot hole coring near

tunnel surface Overcoring in PGR near tunnel surface Overcoring in PGR at 500 mm depth from tunnel surface

Few of the measurements failed due to flooding of flushing water and measuring device related problems. The performed tests were:

1. Overcoring in VGN after at least 12 h cooling time after pilot hole coring. (Success)

2. Overcoring in VGN near tunnel surface (Success) 3. Overcoring in VGN at 500 mm depth from tunnel surface (Failed) 4. Overcoring in PGR near tunnel surface (Success) 5. Overcoring in PGR at 500 mm depth from tunnel surface (Success) 6. Redone the overcoring in VGN near tunnel surface (Failed) 7. Redone the overcoring in VGN at 500 mm depth from tunnel surface (Partial

success)

Overcoring tests were done in late February 2012. Schedule is presented in Table 1.

The in-situ overcoring induced heat was measured during overcoring from the inner surface of the hollow core and from few control points around and under the drill bit.

Measurements were designed to be performed in as realistic overcoring conditions as possible, while keeping as much control of the process as possible. Test arrangement is illustrated in Figure 8 and Figure 9. Temperature development during overcoring is monitored with two devices. First of these devices is referred as temperature cell and it measures pilot hole wall temperature. Second device is called tube sensor and it is installed in a 12 mm diameter bore hole and used to measure temperature increase in the rock just in front of the advancing drill bit. Tube sensors are used in pairs as accurate positioning of the tube sensors is quite challenging.

Temperature field of the rock cylinder surface was monitored during overcoring process with 20 evenly distributed measuring points. Tube sensors contain four sensors each adding to total of eight sensors. Tube sensors can only be used near tunnel wall, as it is difficult to drill instrumentation holes accurately to greater depths.

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Table 1. Testing schedule.

Test Date Overcoring start 1 28.2.2012 2:04 pm 2 29.2.2012 10:47 am 3 1.3.2012 9:42 am 4 1.3.2012 3:28 pm 5 2.3.2012 10:14 am 6 8.3.2012 3:21 pm 7 9.3.2012 10:16 am

Figure 8. Test arrangement, partial cut, top view. Black device in the pilot hole is the temperature measuring device. Drawing is not in scale.

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Figure 9. Photograph of the test arrangement. A device to measure hollow core surface temperature is installed in the pilot hole. Instrumentation holes for temperature measurements of the surrounding rock are on both sides of the pilot hole (tube sensors).

2.3 Measuring equipment

2.3.1 A device to measure hollow core surface temperature field

Temperature field on the hollow core surface is monitored in 20 points with digital temperature sensors. The monitoring is performed online via data cable drawn through the axel of the drill. Online monitoring is considered necessary to detect possible problems during core drilling, such as core disking or cooling water leakage. Gathered data is also immediately recorded by computer, which decreases probability of data loss due to hardware failure.

The sensors are read with digital data logger, which basically functions as an USB interface for the sensors. The values of each individual sensor are sent to a PC at user selected intervals for display and recording purposes.

The digital sensors are Maxim DS18B20 1-Wire® protocol IC (integrated circuit) devices (Figure 10). The DS18B20 sensors produce digital absolute temperature value in degrees Celsius with 12 bit AD conversion, which gives a resolution of 0.0625 degrees. The sensors are factory calibrated to ±0.5 °C accuracy between -10 °C and +85 °C.

The 1-Wire® protocol used by Maxim DS18B20 sensors allows to connect arbitrary number of sensors to a common serial databus, from where each individual sensor is identified by its unique 64 bit serial code. This means that it is possible to monitor dozens of sensors in very simple manner.

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Figure 10. The Maxim DS18B20 digital temperature IC sensor. Pin presented as scale.

Figure 11. A device to measure hollow core surface temperature connected to associated USB interface.

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The fact that the sensors are stand-alone devices that do not require any external electronic components to function makes the use of digital sensors very flexible. All the sensors can be connected parallel to same wires, which enable communication between PC and sensors using only three wires: data, ground and power.

For mechanical protection sensors are encapsulated in small, thin-wall brass caps. Good contact between sensor caps and rock is ensured by pressing the sensor caps against the core surface with springs. Sensor caps are attached to the instrument body with threaded plastic inserts. Drawings as well as photographs of both the parts and the assembled sensor cap are presented in Figure 12 and Figure 13.

Maximum sampling rate of the system depends on the used sensors. Due to slow AD conversion of the sensors, theoretical maximum sampling rate is one sample per second.

Due to inevitable small deviation between sensors, the sensors are used in differential manner, i.e. the difference to a baseline measurement is used to even out small inconsistencies.

Figure 12. Exploded and assembled views of the sensor cap.

Figure 13. Photographs of the sensor cap and associated parts (without the sensor IC).

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The instrument body is a hollow tube made of polyoxymethylene plastic (POM, or polyacetal). The body is sealed from the both ends with plastic end plates. The bottom plate is a simple plastic plate with diameter of 126 mm and it is attached to instrument body with three Allen screws. The upper plate is more complicated, as it needs to provide necessary water proof seal to prevent cooling water from entering the core and effectively ruining the measurement. The body has 20 machined threaded sensor installation holes in rows of five on four opposite sides. A sketch of device body and associated parts is presented in Figure 14. Main dimensions of the temperature measuring device are presented in Figure 15.

Figure 14. Partial radial cut of the temperature measurement instrument body and associated parts. 1) Instrument body 2) Sensor cap (see Figure 12) 3) Bottom plate 4) Front plates 5) 4xM5 Allen screws for tensioning 6) 8xM4 Allen screws to connect front and bottom plates 7) Water tight cable seal 8) Sealing O-rings.

Figure 15. Assembled measuring device body dimensions.

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2.3.2 Sensor tubes

Measuring temperature change in front of the advancing drill bit is more time critical than the hollow core surface temperature field measurement. Rock as material is quite poor heat conductor and as preliminary test results confirm, the temperature changes are measurable only a few seconds before the drill bit reaches the sensor. Temperatures might also get very high, depending on rock type and quality and the speed of drilling. For these reasons it is necessary to use a sensor that has significantly lower thermal time constant than digital DS18B20.

Tube sensor consists of four epoxy-moulded NTC-thermistors. The thermistors do not have any additional mechanical protection to keep the thermal mass as low as possible. The measuring system has theoretical thermal time constant lower than 1 second, enabling very fast response to temperature changes.

The sensor tube uses four separate thermistors in row to improve chances to hit straight to at least one sensor with the drill bit. It is also desirable to continue to monitor temperatures with remaining thermistors as drill bit passes the monitoring point. Sensor tube consists of 50 cm long acryl tube measuring 10/8 mm, four NTC sensors and a data cable. A photograph of the tube sensor measuring head is presented in Figure 16.

Good contact between rock and thermistor is essential. Poor contact would slow down the thermal response time. Tube sensors are installed to 12 mm boreholes with epoxy resin, filling possible air gaps between sensors and rock. Epoxy resin also secures sensors to place.

Temperatures are sampled by external datalogger and send to a PC for recording and further analysis.

Figure 16. Measuring head of the tube sensor. The tube consists of 4 thermistors clued on plastic tube.

2.4 Error estimation

2.4.1 Pilot hole temperature cell

Pilot hole wall temperature field measurement involves static and time dependent error sources. Temperature is measured with digital sensors, so a static measurement is only affected by internal accuracy of the sensor and self-heating effects.

Digital sensors are factory calibrated to ±0.5 °C absolute accuracy at 25 °C reference point. According to manufacturer documentation, readings should not deviate more than 0.5 °C in measured absolute temperature range (from 10 °C to 25 °C). Thus for purpose of this study, the internal accuracy of the digital sensors can be determined to be at least ±0.5 °C in both absolute and relative static measurements.

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The major area of interest, however, is the time dependent temperature field during the overcoring. Temperature sensing can never be instant, as the thermal mass of the sensor itself slows down the measurement. Sensors are mounted in brass casings supported by steel springs (Figure 13), which both add more thermal mass to the overall system. The contact between rock face and brass casing also affects to the thermal response time. The thermal time constant of the overall configuration is hard to estimate.

2.4.2 Sensor tubes

Sensor tubes have four 10 kΩ NTC thermistors in half-bridge configuration with 10 kΩ ±0.1 % reference resistors. Further error is introduced by resistance of the cables, which was measured with 5 % accuracy, leading to maximum of ±1 Ω difference in final measurement. This error can be thought as negligible, as it gives only 0.01 % error to a measurement at 25 °C.

Thermistors are accurate to ±0.5 °C of absolute temperature within 25 – 85 °C. By utilizing 3rd order polynomial fitting, linearity of the sensors is better than 0.1 % at any given temperature. Differential measurements have thus excellent accuracy, but due to uncertainties in A/D conversion, accuracy for differential measurements is considered to be under ±0.5 °C.

Thermal time constant for the sensors is <3 s in air, and <0.7 s in stirred oil. Thermal time constant depends on measured medium and quality of contact with the medium. In solid material with good contact the time constant is around 1 s.

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3 RESULTS

3.1 Sensor naming conventions

It is important to take note how the sensors in temperature cell and in tube sensors are named. The naming convention presented here is used through this work.

Temperature cell has 20 sensors divided evenly in both radial and axial directions. Adjacent sensors located on the same side of the cell are referred as sensors columns. The sensors belonging in a sensor column are marked within a green coloured rectangle in Figure 17. There are four sensor columns around the cell and they are referred with letters from A to D.Radially distributed sensors at the same depth are referred as sensor row. This is illustrated in Figure 17, where sensors belonging in a sensor row are within a red coloured rectangle. There are five sensor rows and they are referred with numbers from 0 to 4. Two sensor tubes are installed in each test. These are referred as tubes A and B. Sensors are named so that they are prefixed with letter A or letter B. Number one sensor is located furthest in the hole and number four closest to tunnel surface. Tube sensor installation and numbering is presented on right hand side in Figure 17.

Figure 17. Naming convention. Temperature cell on left and tube sensor on right. Drawing is not in scale.

3.2 Test no. 1 – in VGN at 50 mm depth

First overcoring test was performed in veined gneiss in hole ONK-SH125. A 500 mm long pilot hole was drilled on previous day, to ensure at least 12 h cooling time for surrounding rock. Pilot hole temperature cell was installed in the pilot hole 50 mm from tunnel surface.

Two sensor tubes were installed around the test hole. First tube was installed horizontally to the right side of the hole. Second sensor tube was installed vertically below the pilot hole. Test arrangement and location is presented in Figure 18.

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Overcoring was performed after 15 minutes settling time with 200 mm drill bit. Overcoring to depth of 45 cm took 27 minutes and 19 seconds. Elapsed time versus overcoring advance was recorded every 5 cm. Nothing abnormal was observed during overcoring and the overcoring advance was almost linear, as can be seen in Figure 20 and Figure 21.

The Figure 19 shows uncorrected raw data measured from pilot hole wall during first overcoring test. Due to the number of temperature sensors, thermal noise and quite possibly variable quality of contact between sensors and rock, it is necessary to use averaging. Sensor readings are averaged in radial and axial directions.

Radial averaging is used to track changes in temperature radially around rock cylinder. This is done by calculating average temperature of each sensor column, from A to D, for every time step. For axial direction, similar averaging process is used. Average temperature of each sensor row, from 0 to 4, is plotted against time.

Figure 18. Test arrangement, front view. Paths of tube sensors are marked with dashed red lines.

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Figure 19. Uncorrected raw data measured from pilot hole.

Averaging can be justified because the measured data indicates that the temperature field is axially symmetrical. In case of asymmetric heating the averaging in axial direction should not be used, as this might bias the results.

Averaged temperature plots of pilot hole sensor readings as well as overcoring advance and location of the sensor rows for test 1 are presented in Figure 20 and Figure 21.

No significant temperature difference between sensor columns can be observed during overcoring. However, in the cooling period, a clear vertical temperature gradient can be seen in Figure 20. Maximum temperature difference (to baseline measurement) is reached at about 48 minutes after the start of the overcoring, being around 3.0 to 3.25 °C.

Axial temperature reveals clear depth dependence to maximum temperature reached. Maximum temperature difference between sensors rows is about 0.2 °C, from 2.60 °C (row no. 0) to 3.65 °C (row no. 4). Heat pulse is detected by first sensor row at 00:05:45 (coring advance 10 cm). Time difference between sensor row raise times is around 00:02:30 – 00:02:40, which is in line with overcoring speed.

First sensor row shows sudden temperature increase near the end of the overcoring. This might be explained by changes in flushing water circulation, as well as another notable phenomenon, which is seen in Figure 21. The maximum temperatures for each sensor row are reached at slightly different times and in reverse order. Row 4 is the first to achieve maximum temperature and row 0 the last. This gives indication that the overcoring temperature rises considerably as the drill bit penetrates further to the rock.

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Figure 20. Radial temperature variation, test no. 1.

Possible reason for this behaviour is change in flushing water circulation, as the 500 mm long drill bit is drilled almost full length to the rock. It is suspected that as the overcoring advances the flushing water can circulate longer in the hole, thus getting warmer and reducing cooling capacity.

Figure 21. Axial temperature variation, test no. 1.

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Temperature log from the tube sensors is presented in Figure 22. After installation, sensors B1 and B2 were not giving consistent readings, and it is suspected that the sensors were accidentally destroyed by excess force used in installation process, as the fit between sensor tube and installation hole was very tight.

Goal was to hit only the sensors 1 and 2 with the drill bit and leave sensors 3 and 4 in both tubes intact. Unfortunately accurate placing of the sensor tubes is quite challenging task and so all sensors except sensor A4 were destroyed during overcoring.

The maximum temperature increase was recorded by sensor B3, +22 °C before it was destroyed. The A3 sensor gave maximum reading of +19.3 °C.

Figure 22. Tube sensor readings.


Second test was performed in veined gneiss in hole ONK-SH124. Overcoring was performed right after pilot hole coring, simulating LVDT measurement. Idea was to see if additional heat induced by pilot hole coring could affect the results. As in test no.1, the temperature cell was installed in the pilot hole 50 mm from tunnel surface.

Two sensor tubes were installed around the test hole. Unlike in first test, this time both sensor tubes were installed horizontally to left and right sides of the pilot hole, as it was suspected that the temperature of the flushing water falling from the hole during overcoring could be conducted to the sensors via data cable. Test arrangement is presented in Figure 23 and Figure 24.

Overcoring was performed after 15 minutes settling time with 200 mm drill bit. Overcoring to depth of 45 cm took 23 minutes and 23 seconds. Elapsed time versus overcoring advance was recorded every 5 cm. Overcoring advance was again almost linear, as can be seen in Figure 26. No problems or anything unusual was observed during overcoring.

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Figure 23. Test arrangement, front view, test no. 2. Instrumentation holes marked with dashed red lines.

Figure 24. Test arrangement, top view. Tunnel wall surface on left side of the figure.

Radial temperature variation during overcoring is negligible. However, in the cooling period, a clear vertical temperature gradient can be seen in Figure 26. Similar downward temperature response can be seen at the end of the overcoring as in test no. 1. Maximum temperature of 3.81 °C was recorded at 10 min 30 s after the overcoring was stopped, at 00:33:53. Radial temperature variation for full measuring period is plotted in Figure 25

The Figure 25 reveals that the final temperature is lower than initial temperature. The final average radial temperature difference at the end of the measurement is -0.66 °C, which corresponds to absolute temperature of 12.24 °C, initial absolute temperature being 12.90 °C (average from all sensors). This difference between initial and final temperatures can be due to either residual heat from pilot hole coring or thermal mass of the measuring device itself. Temperature of the incoming flushing water was 5-6 °C.

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It can also been seen from the Figure 25 that complete equilibrium is not quite reached during the measurement, and the final temperature is slightly above ambient rock mass temperature. The possible effect of residual heat can be assessed by comparing the results with test no. 1. The test 1 was performed about 12 hours after pilot hole coring, thus avoiding residual heat from pilot hole coring affecting the measurement. The initial absolute temperature in test no. 1 was 12.14 °C.

Initial temperature from the test no.1 and final temperature measured in test no.2 indicate that there is indeed possibility that residual heat from pilot hole coring might have an effect on maximum temperature reached during overcoring. The difference is however so small, that a definitive conclusions cannot be made.

As was the case with test no. 1, the axial temperatures reveal clear depth dependence to reached maximum temperature (Figure 27 and Figure 28). In test no. 2, maximum temperature difference between adjacent sensor rows varies between 0.22 °C and 0.63 °C. Heat pulse is detected by first sensor row approximately at 00:04:00 (coring advance 9.3 cm), which is slightly faster than in test no. 1. Time difference between adjacent sensor row raise times is again in line with overcoring speed.

All sensor rows show a temporary decrease in temperature near the end of the overcoring, which is due to increased flushing water pressure for technical reasons.

The maximum temperatures for each sensor row are reached at slightly different times and in reverse order, like in test no. 1. Difference in maximum temperature between adjacent sensor rows is more linear in this test than in test no. 1.


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Figure 26. Radial temperature variation, test no. 2, first two hours.


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Figure 28. Axial temperature variation, test no. 2, first two hours.

All sensors except sensors A4 and B4 were destroyed during overcoring. Maximum peak temperature raise of 15.12 °C was recorded by sensor A4, as drill bit passed it at 00:08:44. Almost the same temperature was recorded by sensor B3. Sensors located at the bottom of the hole were likely destroyed before they could provide higher temperatures. Results are plotted in Figure 29 and Figure 30.

The installation hole is in 45° angle to drill bit, which means that the wirings of the bottom sensors can actually be cut before the drill bit reaches them. This problem was not realised in the design phase, as heat pulse was expected to reach sensors much earlier than it was finally measured. This also raised a question about adequate sampling rate and sampling rate was increased to 20 Hz (one sample every 50 milliseconds) in the final test involving tube sensors (test no. 4).

From the results of tests 1 and 2 it is clear that the temperature right in front of the drill bit is not very high. Another point that should be made clear is that the temperature of the rock in the proximity of the drill bit is higher on the outside of the bit, and lower on the inside of the bit, because flushing water is flowing to the drill bit from inside the bit.

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Figure 29. Tube sensor readings, test no. 2.

Figure 30. Tube sensor readings, test no. 2, to 40 minutes.

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Test no. 3 was performed in veined gneiss in same location as test no. 2 (ONK-SH-124). The purpose was to see if increased depth had an effect to measured temperature difference. Temperature cell was installed in the pilot hole at 500 mm depth from tunnel surface.

Sensor tubes were not used in this test, as it was not possible to drill long enough installation holes accurately with available equipment. Test arrangement is presented in Figure 31.

Overcoring was started after 30 minutes settling time with 200 mm drill bit. Elapsed time versus overcoring advance was recorded every 5 cm. As can be seen from the Figure 32, there is probably some residual heat from pilot hole coring involved, as the temperature continues to decrease. Overcoring was interrupted after 8 minutes and 56 seconds at 15 cm depth, as temperature cell stopped responding. As the cell was retrieved from the hole, it was noticed that the water tight seal at the end of the cell had leaked, filling the pilot hole with water and short circuiting electronics.

Temperature response in this test seems similar to the one seen in test no. 2. Radial temperature increase is similar in all sensor columns. A very small temperature difference can be seen between pairs A-D and B-C. Difference is however so small, that radial temperature distribution can be called even. Maximum temperature difference to baseline measurement before interruption varies between 2.6 °C and 3.1 °C. Overcoring has at this point only passed second sensor row, so expected maximum temperature for this test is considerably higher than in test no. 2.

Figure 31. Test arrangement for test no. 3. Test was performed at 500 mm depth from tunnel wall surface. Tube sensors could not be used due to the depth of the test.

Axial temperature plot (Figure 33) shows similar behaviour as in test no. 2. Heat pulse is detected by first sensor row at 00:32:28, one minute and 10 seconds after overcoring start (coring advance 3.1 cm). Maximum temperature difference is caught by the first

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sensor row and it is about 3.89 °C, which is also higher than in test no. 2 at the same overcoring depth.

Results indicate that overcoring at greater depths is also increasing the temperatures, probably due to changes in flushing water circulation, as is already discussed in chapter 3.3.


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3.5 Test no. 4 – in PGR at 50 mm depth

Test no. 4 was performed in pegmatitic granite (PGR), in hole ONK-SH126. Goal for this test was to measure overcoring induced heat in harder rock type. PGR is much harder to core drill than VGN in the previous tests. Test arrangement was similar to the test no.2. Pilot hole temperature cell was installed in the pilot hole 50 mm from tunnel surface and two sensor tubes were installed horizontally around the test hole on both sides. Figure of the test arrangement is presented in Figure 34.

Figure 34. Test arrangement, front vie, test no. 4.

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Overcoring was performed after 15 minutes settling time with 200 mm drill bit. Plan was to drill to the same 45 m depth as in test no. 2, but wearing of the drill bit in these harder conditions meant that only a depth of 27 cm was reached. Test could not be interrupted for drill bit sharpening, as this would have skewed the results.

Overcoring to depth of 27 cm took 46 minutes and 29 seconds. Elapsed time versus overcoring advance was recorded every 5 cm. Overcoring was notably more difficult than in previous tests. Apart from being rather slow process, nothing abnormal was observed during overcoring.

Unlike the tests performed in veined gneiss, some temperature difference between sensor columns can be observed at the end of the overcoring, as seen in Figure 36. Sensors on the right hand side of the cylinder measure over 1 °C at maximum higher average maximum temperature than the sensors on the left side. This is most likely due to the flushing water circulation. Cool flushing water is flowing at the bottom, from where it is lifted and circulated by rotating drill bit in clockwise direction. Cooling rate seems to comply with the results from VGN tests, as can be seen from the Figure 35.

Maximum temperature difference (to baseline measurement) is reached 100 s after the end of the overcoring, varying from 8.80 to 10.35 °C between sensor columns. Comparing to earlier tests, it is probable that a higher maximum temperature would have been reached, if the overcoring could have been continued to planned depth of 45 cm.

Slower overcoring speed allows heat to conduct further ahead of the drill bit. This is visible in the axial temperature plots (Figure 37 and Figure 38), where temperature especially in sensor rows 3 and 4 show much more gentle reaction to approaching drill bit. All sensor rows reach maximum temperature simultaneously at the end of the overcoring. First 4 rows display same maximum temperature, but the last sensor row, row no. 4, has a lower maximum temperature.


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Figure 36. Radial temperature variation, test no. 4, first two hours.


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Figure 38. Axial temperature variation, test no. 4, first two hours.

Tube sensor readings were more reliable in this test than in previous tests, because slower overcoring gave sensors more time to react to temperature changes. Sampling rate was also increased from 2 Hz to 20 Hz for better time resolution. Tubes were slightly modified to enable easier fit in the drill hole, but unfortunately one sensor (A4) broke in spite of all the precautions.

All sensors except sensors A3 and B4 were destroyed in overcoring. The first sensor to be destroyed was the sensor B3, which measured maximum temperature increase of 30 °C. This can be considered as a quite reliable result, as the sensor has been directly hit by the drill bit. Other sensors measured lower temperatures before destruction, but in those cases the drill bit has probably cut the wires of the sensors before actually reaching the actual sensors.

Readings from the two sensors which were not destroyed can be used to estimate flushing water temperature after the drill bit has passed their location. Graphs of these two sensors A3 and B4 in Figure 39 and Figure 40 show how the flushing water temperature fluctuates and slowly increases as the overcoring advances. Temperature quickly drops as soon as the overcoring stops. Final temperature measured about 16 hours after the overcoring shows approximately 1 °C decrease in temperature compared to baseline measurement at the beginning of the overcoring (Figure 39). Possible reasons are residual heat from pilot hole coring, tunnel air cooling the sensors and heat generated by hardening epoxy resin as the sensors were installed.

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Figure 39. Tube sensor readings, test no. 4.

Figure 40. Tube sensor readings, test no. 4, first 2 hours.

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3.6 Test no. 5 – in PGR at 300 mm depth

Test no. 5 was performed in pegmatitic granite (PGR), in the same hole as test no. 4 (hole ONK-SH126). Goal for this test was to repeat test no. 4, but deeper in the rock. PGR is much harder to core drill than VGN in the previous tests. Test arrangement is presented in Figure 41. Tube sensors were not used in this test.

Overcoring was performed after 15 minutes settling time with 200 mm drill bit. Overcoring to depth of 27 cm took 44 minutes and 52 seconds, which was also a considerably longer than overcoring in VGN. Elapsed time versus overcoring advance was recorded every 5 cm. Due to hard rock, overcoring advance was not linear and the speed of the overcoring varied during the overcoring.

Figure 42 and Figure 43 show clearly how harder rock type affects the temperature field. Radial temperature distribution is quite similar than in tests done in VGN, but a definitely higher peak temperature is reached. In this test the columns C and D are located at the bottom of the hole. The cooling effect of flushing water is really clear after the overcoring has stopped, as the columns C and D show relatively fast decrease in temperature, columns A and B showing much more stable behaviour. The maximum temperature measure by the columns varies from 12.3 °C (column B) to 14.0 °C (column D).

The axial temperature plot shows very clear how the difficult conditions result in rather unpredictable thermal profile. The overall behaviour appears to be the same and it is expected that the sensor rows at the bottom should have reached higher temperatures than the rows close to the tunnel surface, if the overcoring could have been continued to planned 45 cm depth. Maximum temperature varies between 10.44 °C and 14.47 °C (row no. 0 and 3, respectively). Heat pulse is detected by first sensor row approximately at 00:10:00, one minute and 37 seconds after overcoring start (coring advance 2.3 cm).

Figure 41. Test arrangement, front view, test no. 5. Temperature cell is installed at 300 mm depth in tunnel wall. Tube sensors were not used due to the depth of the test.

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The most interesting phenomenon in test no. 5 is again the maximum temperature, which is nearly 4 °C higher than in test no. 4. This is again expected to be due to interference in flushing water circulation, as test no. 5 is performed deeper in the rock then test no. 4.

Test was interrupted at 03:32:48 (02:39:33 after the overcoring stop) due to schedule reasons, but as can be seen from Figure 42, the cooling period behaves as expected, and can be analytically estimated further in time using logarithmic fitting.



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The goal for tests number six and seven was to repeat tests number two and three, by performing two measurements in the same hole at different depths. Sixth overcoring test was performed in veined gneiss in the hole ONK-SH-138. Temperature cell was installed in the pilot hole at 50 mm depth from tunnel surface. Test arrangement is presented in Figure 44.

Due to reliability problems with temperature cell, detected just before testing, all the sensors in column A were disabled as a precaution.

Overcoring was performed after 10 minutes settling time with 200 mm drill bit. As can be seen from the beginning of the measurement in Figure 45 and Figure 46, the settling time was probably not long enough compared to the initial temperature of the cell, causing a notable gradual decrease in measured temperature. Elapsed time versus overcoring advance was recorded every 5 cm. The overcoring advance was almost linear.

Temperature cell stopped functioning at 00:23:56, 19 minutes and 52 seconds after overcoring start. At this point coring had advanced to 34 cm depth. Overcoring was however continued in case the cell recovered, but at 44 cm depth at 00:31:36 after overcoring start overcoring was finally stopped as it became obvious that the temperature cell was broken.

As the cell was retrieved from the pilot hole, the reason for malfunction became obvious, as the hole was half full of flushing water. Closer examination revealed two possible reasons for water leakage. At first it was suspected that the seal of the temperature cell was failed, for instance, due to inadequate or asymmetric tensioning. Another possibility was discovered when the hollow core was removed from the hole. An axially orientated crack was found penetrating through the core, following rock type contact. The crack was not visible prior to overcoring and it is uncertain whether it could have opened enough during overcoring to leak such amount of flushing water as was discovered from the pilot hole.

Figure 44. Test arrangement, front view, test no. 6. Test was performed at 50 mm depth from tunnel wall surface.

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The radial data however does reveal a sudden temperature increase in column C, which was located right at the suspected crack, thus giving evidence that the crack might have been involved in the flooding.

Radial temperature profile shows a small time difference between columns, but as the slower reacting columns are also located at the bottom of the hole, it suggests that the leaking flushing water is first cooler and then rapidly heats up as the overcoring advances. It should also be noted that the temperature response of all sensor data is slower than in previous tests, also possibly indicating water leakage from right at the beginning of the overcoring.

Estimated maximum temperature increase could have been around 3 to 4 °C. Measured maximum axial temperature increase was 2.73 °C and maximum radial increase was 3.47 °C. Despite the problems with flushing water leakage, both the overall heat generation signature and maximum temperature differences conforms with previous measurements done in veined gneiss (tests 1 to 3).


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The last overcoring test, test no. 7, was performed in veined gneiss in the same hole as test number six, ONK-SH-138. Pilot hole temperature cell was installed in the pilot hole 500 mm from tunnel surface. Test arrangement is presented in Figure 47.

The temperature cell was completely disassembled, dried and cleaned after being flooded during the test no. 6. Previously problematic sensor column A was also restored and the temperature cell was working flawlessly during extended periods of testing before overcoring test number seven.

Figure 47. Test arrangement, front view, test no. 7. The test was performed at 450 mm depth from the tunnel wall surface.

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Overcoring was performed after 10 minutes settling time with 200 mm drill bit. Unfortunately the temperature cell again stopped working half way through the test at 00:36:00, 23 minutes and 19 seconds after overcoring start. As in test no. 6, overcoring was proceeded until it was clear that the cell was not going function anymore. Overcoring was stopped to 40 cm depth at 00:40:31 (00:27:50 after overcoring start). Elapsed time versus overcoring advance was recorded every 5 cm. There were no problems associated with overcoring and the advance was almost linear.

The likely cause for the unstable behaviour of the temperature cell was later traced to be slightly oxidized screw connectors connecting individual temperature sensors to the main data bus line. This could have been avoided by using soldered connections at each data bus junction. The fact that the stability problems begun after the first flooding of the cell during the test number three and gradually became worse and worse, also suggests that the water contact and moist tunnel air were oxidizing the contacts and thus creating stability issues.

The radial and axial temperature plots (Figure 48 and Figure 49) show similar behaviour with tests done with same lithology. This time it is obvious that the temperature cell was not settled to the tunnel wall temperature before beginning of the test. Part of the temperature decrease can be explained by residual heat from pilot hole coring, but the total drop of approximately 0.5 °C from start of the measurement to detection of the heat pulse is far too much to be completely explained by that phenomenon.

Radial temperature plot shows similar behaviour as in previous successful tests. Columns C and D, which are located at the bottom, are responding a bit slower but still seem to achieve same maximum temperature.

Axial temperature plot also reveals similar behaviour as measured in earlier tests in veined gneiss. Sensor rows at the bottom of the hole achieve greater maximum temperature and also achieve that temperature faster.


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4 NUMERICAL MODELLING

4.1 Modelling paradigm

The performed seven overcoring heat tests gave good indication of the temperature field of the rock cylinder during and after the overcoring. The measuring method however did not allow performing actual LVDT measurements at the same time with temperature measurements. A numerical analysis was thus used to further study the magnitude of displacements caused by thermal expansion of the rock during the overcoring.

Physically the overcoring process is quite complex to model. The physical phenomenon during the overcoring includes, for example, relatively fast advancing and rotating drill bit, complicated flushing water flow, possibly varying hardness of the rock, varying force applied to the drill bit and wearing of the drill bit. The amount of heat generated at the drilling face is unknown and hard to estimate. Heat also transfers along the drill bit body by conduction. The drill bit interacts with the flushing water by transferring heat to and from the drill bit in very complex and unknown manner.

Two of the measurements were chosen to be the reference data for the numerical simulations. Test no. 2 overcored in veined gneiss was chosen to be the primary calibration data for the models, as the overcoring advance in that test was almost linear and because the cooling period was fully recorded. Test no. 2 is considered to be a good approximation for an overcoring in relatively good conditions. Second reference test was the test no. 4, which was overcored in pegmatitic granite. This test was chosen to provide estimation for the worst case scenario, where overcoring is slow and heat production is high.

The data from these two tests were further analysed to find a possible analytical function, which could explain reasonably well measured data. As the radial data suggests that the overcoring process is axi-symmetrical, it was noted that the actual problem reduces from four dimensions (three physical and the time) to only three (two physical dimensions and the time). This three dimensional data was curve-fitted with different analytical formulas using a dedicated software, TablePlot 3D. The best fit was achieved using combinations of nonlinear peak functions, namely lognormal distribution and Gaussian distribution.

None of the user generated or functions suggested by the software was able to model the complete thermal behaviour accurately. Satisfactory analytical function was not found, but the curve-fitting did provide some information to be used in numerical simulations.

Two and three dimensional modelling techniques were evaluated to be used in numerical simulation. As previously discussed, the modelling problem can be reduced into two dimensions in physical domain. It was thus considered to be justified to perform FEM-modelling of the problem with 2D axi-symmetrical model. True three dimensional modelling was not considered to provide any critical value the simulation.

After test simulations the FEM model was further simplified to include only the overcored rock cylinder and rock mass right behind it (Figure 51). Full model containing also the surrounding rock was found to complicate the modelling without providing enough value to the simulation. This approach has a few obvious limitations.

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First, there is no mechanical support on the outer edge in front of the advancing overcoring, as in the real life situation, where rock cylinder is gradually separated from the rock mass by core drilling. This could potentially affect to the modelled displacements. Another concern is similar, but related to the heat transfer. Heat transfer to the surrounding rock mass during overcoring cannot be modelled and this might yield to incorrect thermal model.

First of these two problems is handled by introducing mechanical support mathematically by using a stiff boundary condition on the boundary ahead of the overcoring. Second problem is a more complicated one. It was noticed in the test simulations and in the field tests that the speed of the drill bit is relatively fast compared to the thermal conductivity of the rock. This means that the heating of the surrounding rock mass during overcoring does not play an important role in the heating of the rock cylinder.

Heat transferred to the rock cylinder can be modelled similarly as a time-varying boundary condition. It should also be noted that the heat transfer of the flushing water circulation is unknown and contributes significantly to the total heat transfer. It is very complicated to take all of these into account and the heat transfer is simplified by utilising all the phenomena into a single thermal boundary condition.

Surrounding rock mass can have some influence to the cooling of the rock cylinder via radiation or convection through air. This effect is considered to be small and cooling is thought to happen mostly by conduction from the base of the cylinder and by convection to the tunnel air. It is important to note, that at the time of the field measurements (late February), the rock mass was warmer than the tunnel air. A small temperature gradient was measured in a fully cooled rock cylinder in the field tests. This effect was also considered to have negligible effect and could be thus excluded from the simulation. Thermal mass of the temperature measuring system is biggest concern, as it can bias measured values considerably.

Numerical simulations were divided to two phases. Thermal simulation was first used to understand and model the heat distribution during overcoring. Thermal boundary conditions were then introduced to coupled thermo-mechanical model, which was in turn used to estimate maximum displacements induced by heat generated during overcoring.

4.2 Geometry and material properties

Numerical modelling was performed with Comsol Multiphysics 4.3 finite element code. Separate models were used for thermal and thermo-mechanical modelling. The basic model consists of two domains, a hollow rock cylinder and supporting rock mass. This model is used in the thermo-mechanical modelling. In transient heat transfer model the temperature measuring device is also modelled to introduce thermal mass, which can potentially affect the measured temperature readings. Due to 2D axi-symmetry of the model, brass cups surrounding the sensors are modelled as rings. This is demonstrated in Figure 50, where 2D cross section and 3D revolution of that cross section are presented.

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Length of the rock cylinder is 450 mm. Inner diameter is 127 mm and outer diameter is 190 mm. Supporting rock mass length is 600 mm and diameter is 190 mm. The length of the measuring device is 230 mm and wall thickness of the device is 19.5 mm. Cross-section of the brass rings between the rock cylinder and the measuring device body is a 5 mm by 5 mm square.

Measuring device was removed from the coupled model to prevent it affecting displacements. Dimensions for the model are the same as in the heat transfer model. Length of the cylinder is 450 mm with inner diameter of 127 mm and outer diameter of 190 mm. Rock mass domain length is 600 mm and diameter is 190 mm. The model is presented in Figure 51.

Figure 50. Thermal model geometry including simulated temperature cell as thermal mass. Overcored rock cylinder is located on top of rock mass. Black cylinder is simulated measuring device, which is connected to the overcored rock cylinder by thin brass rings presented in orange.

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Figure 51. Thermo-mechanical model geometry. Temperature measuring device is not simulated in this model, as it was only used to calibrate thermal boundary conditions.

Thermal and elastic deformation properties of the Olkiluoto rock types are extensively studied by various authors. Comprehensive study of the thermal properties of the Olkiluoto rock types is published in the Posiva work report WR 2011-17, Thermal Properties of Rocks in Olkiluoto: Results of Laboratory Measurements 1994-2010 (Kukkonen et al. 2011). Thermal properties for the two rock types used in this study are from the previously mentioned publication and are presented in Table 2.

Elastic deformation properties of the rock types in Olkiluoto are published in (Andersson et al. 2007). These properties for veined gneiss (VGN) and pegmatitic granite (PGR) are also presented in Table 2.

Table 2. Thermal and elastic deformation properties of the two tested rock types.

Parameter Veined gneiss (VGN)

value ±stdev Pegmatitic granite (PGR)

value ±stdev Young’s modulus [GPa] (1 62 ±9 65 ±9 Poisson’s ratio [-] (1 0.25 ±0.04 0.29 ±0.05 Density [kg m-3] (1 2734 ±37 2611 ±25 Thermal conductivity [W m-1 K-1] (2 2.83 ±0.53 3.20 ±0.41 Specific heat [J kg-1 K-1] (2 725 ±33 689 ±17 Diffusivity [10-6 m2 s-1] (2 1.37 ±0.25 1.75 ±0.18 Coefficient of thermal expansion [10-6 K-1] (3 7.1 – 12.3 3.2 – 10.8 1) Andersson et al. 2007 2) Kukkonen et al. 2011 3) Åkesson 2012

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4.3 Transient heat transfer model

4.3.1 Boundary conditions

First part of the numerical modelling was to understand and model different phenomena introduced in overcoring heat tests. The purpose of thermal modelling was to find suitable boundary conditions to be used in thermo-mechanical modelling. Thermal model was also used to locate and analyse possible problems related to measurements in overcoring tests.

Different combinations of boundary conditions were tested to find most suitable set up for the thermal model. The rock material in the model is initially set to a constant temperature. Residual heat from the pilot hole coring is also neglected. Inner surface of the rock cylinder is set as insulated, which is justified because the temperature cell effectively blocks air circulation inside the pilot hole thus enabling only a very small convective heat transfer to air inside the pilot hole. Open end of the rock cylinder has a convective cooling applied to it to simulate flushing water and convective cooling to tunnel air temperature. Outer surface of the rock cylinder is also cooled by convection. Other boundaries are set to infinite outflow to simulate continuous rock mass.

Boundaries of the temperature measuring device are set insulated. Contact between device body and brass rings is perfectly conductive. Contact between brass rings and rock cylinder is insulated by thin, thermally resistive layer. Thickness of the layer is set to 1 mm and thermal conductivity to 0.2 W m-1 K-1, which resemble plastic material. Thermal boundary conditions are presented in Figure 52.

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Figure 52. Thermal boundary conditions. Rock mass is surrounded by infinite outflow, inside of the overcored rock cylinder and the measuring device are insulated. Rock cylinder is heated from outer boundary. End of the cylinder is cooling by convection. There is a thin, thermally resistive layer at the contact between brass rings and rock cylinder.

Fundamental assumption in thermal modelling was that the overcoring induced heat can be modelled as a heat pulse. Maximum temperature is reached at the contact of the drill bit and rock, where the heat is generated due to friction. Behind the drilling face the flushing water quickly decreases the induced heating power. This quickly decreasing heat source is modelled as a heat pulse. The heat pulse is mathematically presented as a half of Gaussian pulse advancing on the outer boundary of the rock cylinder at the speed of the overcoring. An example of such normal distributed heat pulse advancing at 0.35 m depth is presented in Figure 53. Standard deviation of pulse used in modelling is 0.05 and mean is 1. The pulse is used to distribute given power along the overcored boundary.

The actual heating power generated in overcoring is hard to estimate. Test data from the overcoring heat tests is thus used to calibrate the thermal model to produce similar results compared to results measured in the field tests.

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Figure 53. Heat pulse on the outer boundary at 0.35 m depth.

Data from the test no. 2 were used as reference data for numerical modelling. In this test the overcoring advance was almost linear and data exists for the whole cooling period. Axial data from the test no. 2 were analysed to distinguish different physical phenomena involved both in the overcoring and temperature measuring. The axial data from test no. 2 is presented in Figure 54.

Closer inspection to this data reveals a few important details about the overcoring process and in particular the temperature measurements. First thing to notice is the time when each adjacent sensor row detects heat pulse. Numerical model, which is adjusted to produce approximately the same maximum temperature, shows about 120 s faster reactions to heat pulse. This effectively indicates about 2 minute delays in measurements, probably caused by the thermal mass of the measuring system. Poor contact between sensor and rock face might also introduce even more delays.

Due to the delay, the numerical model does not return results comparable with reference measurements. The problem is handled in the model by adding a thin thermally resistive layer between the inner surface of the cylinder and data probe.

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Figure 54. Axial data from the test no. 2.

Thermal mass of the simulated measuring device and brass rings has most influence to the thermal model. The mass act as a heat/cold reserve and both cools the sensors in the heating period and slows down the cooling in cooling period. This thermal mass also affects maximum temperatures and is thus important.

Another important feature is that the maximum temperature recorded for a sensor row increases with depth. This is important as it suggests that the total heating power increases with depth. Various different functions were used to evaluate most suitable approach in modelling this behaviour. Final choice was to use hyperbolic tangent function, as it could be easily modified to produce different gradually increasing power profiles with smooth transitions. Following function (1) can be used to provide different transitions between 0 and 1 along by changing parameters a, b and c.

, , , ,∙

(1)

Here parameters a, b and c are positive constants and length is the length of the boundary. Parameter x denotes location at the boundary. This function was selected merely because it is easily parametrically adjusted to represent a variety of different curvatures. Some examples of different configurations are presented in Figure 55.

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Figure 55. Different power profiles with pulse location on x-axel and percentage of full power on y-axel.

Power profile presented in lower-right corner in the Figure 55 is combined with constant term and constant base power to provide final power profile, which is gradually increasing as function of depth. The constant term is necessary, as the relative maximum temperature difference between adjacent sensors would otherwise be far too large compared to reference data. This temperature difference can be adjusted by changing the ratio between constant term and hyperbolic power function. Tests showed that a small residual power is also needed to simulate measured temperatures.

The resulting total power function is thus

∙ , , , , (2)

Where PTOT is the total power applied to boundary by Gaussian function PR is the residual power PB is the base power CV is variable portion of power, related to CC (constant part) by equation 1 = CV+ CC

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Table 3. Modelling parameters for the two modelled tests.

Case a b c Length PR PB CV CC Test 2 5 2 2 0.45 75 430 0.2 0.8 Test 4 N/A N/A N/A N/A 120 120 0 1

Resulting power distribution is demonstrated in Figure 56, where total heating pulse power is plotted at 100 s intervals. Values for all the parameters used to simulate tests 2 and 4 are presented in Table 3. Note that in the simulation of the test no. 4, the variable power part CV was set to zero and so variable power function parameters a, b, c and Length were not used.

Cooling back to the ambient rock temperature is modelled by convective cooling on the outside surface of the rock cylinder, as previously discussed. The heat flux magnitude of this convective cooling is adjusted to match measured cooling from the reference data (the test no. 2).

Modelling the delay of the measuring system is of great importance, as it potentially also effects to maximum temperatures. The measured temperature profiles reach maximum temperature from 10 to 20 minutes after the overcoring has stopped. Thermal mass of the system slows measurements and that the rock cylinder has time to cool down before the sensors reach maximum. As previously discussed, this phenomenon is modelled by insulating the measuring points in the numerical model with a thin, thermally resistive layer.

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Figure 56. Simulated total heating power at 100 s intervals on the outer boundary of overcored rock cylinder. Total heating power is used to combine overcoring induced heat and cooling caused by flushing water to a single transient thermal boundary condition.

4.3.2 Results

The method described in previous chapter seems to approximate measured temperature profiles quite well. Results are presented in Figure 57, which can be compared to the measured temperature profiles in Figure 54. The model was calibrated to give closely the same maximum temperature and cooling rate as measured.

The most obvious and concerning shortcoming in the model is the inability to model accurately the time each sensor row reaches the maximum temperature. Similar effect as what can be seen in the data of test no. 2 can be achieved by introducing additional thermal mass to the model. However, according to test simulations, this also means that the actual temperatures at the rock cylinder surface will rise much higher than measured. This would mean that the temperature measurements are greatly affected by the thermal mass of the measuring device and actual maximum temperatures could be at least twice as high as measured. There might also be some uncaptured phenomena involved, which can explain observed behaviour.

The body of measuring device and thermally resistive layer both add some delay to the temperature measures in the model, but not as much as is measured in the test no. 2.

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This is another indication that either the thermal mass (or initial temperature) of the measuring device is not estimated correctly, or the insulating layer is inadequate.

It should be made clear that used approach is only a one possibility to model thermal behaviour of the overcoring process. Similar results could potentially be obtained by using different boundary conditions.

Figure 57. Modelled axial temperature plot with thermally resisted sensors and side-by-side comparison with measured data.

0

0.5

1

1.5

2

2.5

3

3.5

4

4.5

5

0 600 1200 1800 2400 3000 3600 4200 4800 5400 6000 6600 7200

∆T

[°]

C

Time [s]

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The effect of the thin, thermally resistive layer can be examined by comparing modelling results to similar model without the layer. These results are presented in Figure 58. As can be seen, the maximum temperature is 0.5 – 1 °C higher and sensors detect rising temperature earlier than with the layer.

There is not very large difference in the temperature plots between insulated and uninsulated cases. In addition to higher maximum temperatures, uninsulated model has larger relative difference between the maximum temperatures for each sensor row. Sensors also detect heat pulse slightly earlier than in insulated model.

Tests done in pegmatitic granite were also interesting, as the measured maximum temperatures were much higher and the overcoring speed was much slower. The same model was used to model test no. 4, by changing the overcoring speed and length to correspond with the data from that test. The total heating power and the ratio between variable and constant power were also changed so that the maximum temperatures of the first sensor row matched measured data. The function presented in Equation 1 was not used, but a constant coefficient was used to model power increase as over coring advances.

Figure 58. Axial temperature plot, raw temperatures at the inner boundary of the rock cylinder.

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Results for thermally resisted sensors are presented in Figure 59 and for raw temperatures in Figure 60. The model can predict observed behaviour quite well. The fourth sensor row does not rise to same maximum temperature as the first three rows. The drill bit worn badly during this test and that might be one explanation. The model also predicts the more gentle temperature increase detected in last sensor row.

Few other heat source configurations, including constant power and different linear equations were tested to find a more suitable heat model, but the current model remained as the best approximation.

Figure 59. Simulation of test no. 4. Axial temperature plot with thermally resisted sensors.

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Figure 60. Simulation of test no. 4. Axial temperature plot, raw temperatures at the inner boundary of the rock cylinder.

4.3.3 Sensitivity analysis

Largest uncertainty in the thermal model is related to the boundary conditions. The model is relatively sensitive to any parameter changes concerning boundary heat source, introduced thin thermally resistive layer or convective cooling power. However, these are all more or less artificial parameters, which are selected so that they fit to the measured reference data. The model is fairly sensitive to any changes in these boundary conditions, so induced maximum temperatures can vary considerably from real world situation.

For these reasons model sensitivity analysis is therefore limited to the material parameters. Analysis is performed by investigating the model response to changes in material property values, keeping the same boundary conditions. In other words, heat production is assumed to be same in all configurations and only the material properties are changed.

The material parameters used in thermal modelling are specific heat, thermal conductivity and density. Deformation due to thermal expansion is considered in thermo-mechanical modelling. These material parameters, given in Table 2, are the mean values for the given rock types and can thus vary considerably. Minimum and maximum values for each property were calculated by using standard deviations presented in the same table. Different combinations of these extreme values were used

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to evaluate the heat transfer model. Sensitivity analysis was performed for veined gneiss (VGN).

The most interesting result in the thermal modelling is the maximum temperature reached at the location of the LVDT sensor, which corresponds to the first temperature sensor row (row 0). Another interesting result is the absolute maximum temperature. Table 2 lists these maximum values for different combinations of low and high material property values.

Table 4. Maximum temperature with different material parameter combinations.

Configuration Specific heat [J kg-1 K-1]

Thermal conductivity [W m-1 K-1]

Density [kg m-3]

Max temperature [°C]

Baseline 725 2.83 2734 3.50 1 758 3.36 2771 3.39 2 758 3.36 2697 3.46 3 758 2.30 2697 3.35 4 692 2.30 2771 3.53 5 692 3.36 2697 3.73 6 758 2.30 2771 3.28 7 692 3.36 2771 3.65 8 692 2.30 2697 3.60

Based on the sensitivity analysis, the model is not very sensitive to variation of thermal properties at the given range. The maximum deviation from the model calculated with mean values is approximately ±6.6 % (±0.23 °C).

Highest maximum temperature values were achieved with configurations 5 and 7 and lowest values with configurations 3 and 6. The model is most sensitive to variation of material density. Lower density values give higher temperatures and higher density gives lower maximum temperature values. Variation of thermal conductivity seems to have almost no effect to the maximum temperature.

Thermal properties of pegmatitic granite do not vary considerably from the veined gneiss, so it was decided to not to perform separate analysis for that rock type.

4.4 Coupled thermo-mechanical model

4.4.1 Boundary conditions

Resulting boundary heat source model from the heat transfer simulation was coupled with linearly elastic mechanical simulation to estimate displacements introduced to LVDT device during overcoring.

Model was constrained from bottom of the rock mass domain by fixed constrain. Other boundaries were free to deform. Overcoring boundary was supported in front of the overcoring by stiff spring condition, for which spring constant k equals 2E. However, testing indicated that the spring supported boundary had only very little effect on the induced displacements, so it was discarded for the sake of simplicity.

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Thermal boundary conditions were similar as in transient heat transfer model. Outer boundary of the rock mass domain was set to infinite outflow, outer boundary and the face of the rock cylinder were set to cool by convective cooling, where h = 10 W m-2. The boundary heat source is the same as in heat transfer model in both simulations.

The coupled models use the mean elastic deformation and thermal property values of the rock types, with the exception of coefficient of thermal expansion, which was selected to be the maximum value of the given range.

4.4.2 Results

Two coupled thermo-mechanical models were calculated, first one to simulate the test no. 2 and second to simulate test no. 4. Thermal boundary conditions of the models were the same as in heat transfer model.

Simulation of the test no. 2 gives maximum radial expansion of 3.57 µm. Expansion is fast at the beginning of the overcoring and slows down as the heat pulse advances, following temperature development as expected. First two hours of development of radial displacement in the simulation of the test no. 2 is presented in Figure 61. Full simulation period is presented in Figure 62.

Simulation predicts that radial displacement reduces to less than 1 µm after two hours after the overcoring has been stopped and to less than 0.2 µm after four hours. As the boundary of the rock cylinder is cooled to the temperature of the tunnel air, which is lower than rock mass temperature initially set for the rock cylinder, radial displacement descents below zero (contraction) in final state.

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Figure 61. Radial displacement at the location of the LVDT sensors, simulation of test. no2, simulation duration 2 h.

Radial displacement as a function of temperature increase is presented in Figure 63. The results indicate radial displacement, which follows the development of temperature in almost logarithmic manner. The relation between radial displacement and temperature in the cooling period is clearly linear. Linear relation enables accurate prediction of the replacements quickly after the maximum temperature has been reached and rock cylinder has started to cool down.

Numerical model however represents overcoring in ideal, static conditions. In practice, the cooling rate of the rock cylinder will vary especially right after the overcoring is stopped due to drying of the wet rock, difference in convective heat transfer along rock cylinder boundary and temperature variation and circulation of tunnel air around the rock cylinder.

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Figure 62. Radial displacement at the location of the LVDT sensor, simulation of test no. 2, simulation duration 20 h.

Figure 63. Radial displacement as function of temperature at the location of the LVDT sensor, simulation of test no. 2, simulation duration 20 h.

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Similar simulation was also performed for the model representing the conditions in test no. 4. Maximum radial expansion in this simulation was 9.84 µm. If the simulation of radial expansion presented in Figure 64 is compared to the results of the thermal simulation presented in Figure 59, it can be seen that the radial displacement follows the development of temperature very closely. This confirms the results from the thermo-mechanical simulation of the test no. 2 as the boundary heat source and overcoring advance were quite different in both cases.

Higher maximum temperature also results in longer cooling time in this simulation. The displacement for the full simulation period is plotted in Figure 65. Radial displacement decreases to 1 µm in 2 hours and 25 minutes. Four hour cooling time further reduces radial displacement to less than 0.2 µm.

Figure 64. Radial displacement at the location of the LVDT sensors, simulation of test no. 4, simulation duration 2 h.

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Figure 65. Radial displacement at the location of the LVDT sensor, simulation of test no.4, simulation duration 20 h.

4.4.3 Sensitivity analysis

Thermo-mechanical model is sensitive to both thermal and elastic deformation properties of the used material. Thermal expansion depends on induced temperature increase, so the largest uncertainty in the coupled model is also the boundary heat source. For this reason the sensitivity analysis is performed only on material property values.

The properties affecting to radial displacement are Young’s modulus, Poisson’s ratio, density and coefficient of thermal expansion. The values of these properties and their standard deviations are presented in Table 2 for both rock types.

Thermal expansion is in this case mainly dependent on temperature and coefficient of thermal expansion. For the purpose of this study, we are mainly interested in the maximum radial displacement. Thermal material property values giving the maximum temperature in thermal sensitivity analysis are therefore used in the sensitivity analysis for coupled model. Coefficient of thermal expansion was fixed to maximum value for the same reason. According to the thermal simulation sensitivity tests, density of the material affects greatly to the reached maximum temperature, so variation of density was included in the sensitivity analysis to study the relative sensitivity of thermal and elastic material parameters.

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Maximum radial displacement was calculated for different combination of extreme values of material property values for the location of LVDT sensors (in 8 cm depth). These maximum radial displacement values are presented in Table 5.

The results of sensitivity analysis show that the maximum radial displacement varies only between two values, 3.67 µm and 3.76 µm. Closer inspection reveals that density is the only parameter that contributes noticeably to the maximum value of radial displacement. This difference is due to differences in maximum temperature value, which are 3.73 °C for 2697 kg m-3 and 3.65 °C for 2771 kg m-3.

From this analysis it can be concluded that the model is not sensitive to variation of elastic deformation property values in the given range, and the largest uncertainties are related to thermal boundary conditions and partly to thermal material properties.

Table 5. Maximum displacement with different material parameter combinations, maximum temperature model.

Configuration E

[GPa] ν

[-] α

[10-6 K-1]Density [kg m-3]

Max displacement [µm]

Baseline 62 0.25 12.3 2734 3.71 1 71 0.29 12.3 2771 3.67 2 71 0.29 12.3 2697 3.76 3 71 0.21 12.3 2697 3.75 4 53 0.21 12.3 2771 3.67 5 53 0.29 12.3 2697 3.76 6 71 0.21 12.3 2771 3.67 7 53 0.29 12.3 2771 3.67 8 53 0.21 12.3 2697 3.75

Specific heat: 692 [J kg-1 K-1] Thermal conductivity: 3.36 [W m-1 K-1]

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5 SUMMARY

5.1 Summary of overcoring tests

Total of seven overcoring tests were performed in different conditions to measure overcoring induced heat. A special device was constructed to measure rock cylinder temperature from pilot hole wall from 20 evenly distributed measuring points. In three tests there were also eight additional temperature sensors (“tube sensors”) installed into 14 mm drill holes, which were used to monitor temperature in front of the advancing drill bit.

Overcoring tests and test conditions are summarised in Table 6. From the seven performed tests, four succeeded, one partially succeeded and two failed. Failure in tests 3 and 6 were due to leaking flushing water, which both affected the measurements and cut the connection to the measuring device. The partial failure of test no. 7 was due to instability of the measuring device after it had been soaked during test no. 6.

According to the tests, temperature field of the rock cylinder can be considered axially symmetrical when the overcoring advances relatively fast and induced temperatures are low. A small radial difference was noticed in the tests 4 and 5, which were performed in harder rock type. In these cases the left hand side of the rock cylinder was hotter especially at the end of the overcoring. This can be explained by flushing water circulation with clockwise rotation of the drill bit.

Tests done in veined gneiss (VGN) showed clear depth dependence in maximum temperature. This seems to suggest that the overcoring induced heat rises as the overcoring advances. Tests done in pegmatitic granite did not follow this behaviour, but in those tests the maximum temperature was achieved simultaneously at all depths. It should be noted though, that due to hard rock the overcoring length was not as long as in tests done in VGN. Maximum temperatures achieved in each overcoring test are presented in Table 7.

Table 6. Overcoring test conditions.

Test no. Status Rock type OC length Start depth OC duration 1* Success VGN 45 cm 5 cm 27 min 19 s 2* Success VGN 45 cm 5 cm 23 min 23 s 3 Failure VGN 15 cm 50 cm 8 min 56 s 4* Success PGR 27 cm 5 cm 46 min 29 s 5 Success PGR 27 cm 35 cm 44 min 52 s 6 Failure VNG 15 cm 5 cm 19 min 52 s 7 Partial Success VNG 40 cm 45 cm 23 min 19 s *) Test includes tube sensors

Table 7. Maximum temperature increase measured in each overcoring test at 8 cm depth and absolute maximum temperature increase.

ΔT Test 1 Test 2 Test 3(1 Test 4 Test 5 Test 6(1 Test 7(1

Max at 8cm [°C] 2.66 2.99 3.89 10.05 10.44 1.63 2.47 Absolute max [°C] 3.69 4.72 3.89 10.22 14.47 2.73 3.80

1) Note: test failed

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Tube sensors were installed in tests no. 1, 2 and 4. The results from the tests 1 and 2 do not show very high maximum temperatures. Maximum temperature reached in test no. 1 was 21.93 °C and in test no. 2 15.27 °C.

Overcoring advance in test no. 4 was slower and tube sensor data from this test are considered more reliable compared to tests 1 and 2. Maximum temperature measured by tube sensors in test no. 4 was 30.08 °C.

Internal accuracy of the measurements is very good, but the large body of the pilot hole temperature measuring device can have affected the measurements by adding a considerable amount of thermal mass around the sensors. Contact between the rock and sensors can also be poor. Tube sensors do not suffer from these error sources, but the results of tube sensors might under estimate overcoring temperatures if the overcoring advance is significantly faster than the thermal time constant of the sensors.

5.2 Summary of numerical simulations

Field measurements showed almost no radial temperature variation during over coring, so the temperature field in the rock cylinder can be thought to be axially symmetrical. For this reason an axi-symmetrical 2D finite element model is used to model heat transfer and displacements during overcoring.

Modelling was divided to two parts. First task was to develop a transient heat model capable of simulating measured temperature field during overcoring in two test cases. This calibrated heat transfer model was then coupled with linearly elastic mechanic model to estimate induced radial displacements at the location of LVDT sensors.

The model geometry consisted of a rock cylinder on cylindrical supporting rock mass. Thermal model also included thermal mass of the measuring equipment, which is in contact with rock cylinder via thin rings representing brass covers of the temperature sensors in 2.5 dimensions. This extra geometry was not included in the coupled model.

Overcoring induced heat was modelled as a variable power Gaussian pulse. The pulse moves on the outer boundary of the rock cylinder at the speed of the overcoring. Both generated heating power and the cooling power of flushing water were combined to this single function, as the complicated combined behaviour of heat generation and flushing water circulation is not known. Cooling of the rock cylinder was simulated by adding convective cooling to the outer boundary of the cylinder. Cylinder is also partly cooling by conduction from the supporting rock mass.

Two field tests were simulated by using the numerical model, tests no. 2 and no. 4. Test no. 2 was overcored in veined gneiss and test no. 4 was overcored in pegmatitic granite. Boundary conditions were adjusted until the developed numerical model was able to simulate measured temperature field reasonable well.

Sensitivity analysis performed on thermal material properties revealed that the model is not very sensitive to variation in material property values within one standard deviation from the mean value. Maximum difference to baseline model was about 6.6 %. Density had the largest effect to the maximum temperature and heat capacity had second largest effect. Variation in thermal conductivity value did not seem to have effect to the

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simulated maximum temperatures. The model is most sensitive to changes in in thermal boundary conditions. The model has large uncertainties regarding the boundary heat source, which includes for example changes in heat pulse width, total heating power and cooling power.

The effect of overcoring induced heat on LVDT measurements was evaluated by calculating radial displacement at the location of the LVDT sensors using coupled thermo-mechanical models. Models use the calibrated thermal boundary conditions from thermal simulations. Two models were created to simulate test no. 2 and test no. 4.

The results of the simulation of test no.2 show that the radial displacement increases fast and non-linearly at the beginning of the overcoring. During the cooling period radial displacement decreases linearly as function of temperature. Maximum displacement is achieved approximately at the maximum temperature. Transient heating pulse and the fact that the rock cylinder is cooled from the outer boundary together create a complex thermal expansion behaviour responsible of previously mentioned observations. The maximum radial displacement was 3.57 µm at the temperature of 4.4 °C.

Simulation of test no. 4 gave similar response to radial displacement. Increase of displacement was however more linear at the beginning of the overcoring, compared to simulation of test no. 2. The maximum radial displacement was 9.84 µm.

The model sensitivity was tested against elastic deformation material properties. Different combinations of property values one standard deviation above and below the mean value were used together with thermal material properties giving the maximum temperature. Results show that the model is not sensitive to variation in elastic deformation property values in given range.

Numerical simulations give an estimation of the overcoring heat induced radial displacement, but there are large uncertainties related to thermal boundary conditions. Moderate variations in material property values do not seem to have great effect on simulated displacements.

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6 CONCLUSIONS AND RECOMMENDATIONS

Performed tests give good indication about the thermal behavior of the rock cylinder during overcoring. Although some of the overcoring tests failed, the amount of data was sufficient to be used as a reference data for numerical simulations. Numerical analysis was also used to assess the delays in temperature measurements due to large thermal mass of the measuring system. It is however still possible that measurements from the pilot hole wall might be biased and can under estimate true temperature increase.

Reliability of the tube sensor data is good. The over coring advance in these tests was relatively fast compared to the thermal conductivity of the rock. For this reason the heat pulse is detected by the sensors just a few seconds before they were destroyed. This might lead to a situation where sensors are unable to react to sudden change in temperature fast enough. Test done in PGR was much slower and thus this data is considered reliable.

The most significant observations from the overcoring heat test campaign are as follows:

Temperature increase in good overcoring conditions is not very high. This motivates to perform LVDT measurements in 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 pilot hole is increasing 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 to note that a temperature plot ΔT(t) for sensor rows can be roughly modeled with log-normal distribution. This behavior is valid for fast overcoring. Slow overcoring speed levels out temperature field in cylinder.

Temperature increase in rock cylinder is radially symmetric. Thermal expansion in the cylinder is therefore axially symmetric. In mathematical sense this also reduces overcoring induced heat on the pilot hole wall to three dimensions: depth, temperature and time. A slight asymmetry was detected in PGR tests.

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

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

Rock mass outside of the drill bit reaches higher temperatures than overcored rock cylinder. As the cooling flushing water is injected from inside the drill bit, it provides considerable more cooling power to rock cylinder inside the drill bit than rock mass around the drill bit. This might yield to an interesting asymmetric temperature distribution during side coring measurements.

Numerical simulations suggest that 4.4 °C increase in temperature at the location of the LVDT sensors (as in test no.2) results to maximum of about 3.75 µm radial expansion. This result is for a fast overcoring in VGN. The maximum temperature increase was measured in tests 4 and 5, when overcoring was performed in pegmatitic granite. High amount of quartz in this rock type makes it very hard to core drill, thus creating more

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heat and slowing down the drilling. According to the numerical simulations, 10 °C temperature increase measured in tests 4 and 5 results to maximum of 9.84 µm radial expansions in rock cylinder at the location of LVDT sensors. It should be noted that these displacement values apply only to the maximum temperature.

Based to temperature measurements and numerical simulations it is suggested to leave LVDT cell to measure displacements for at least 4 hours after the end of overcoring, preferably for 6 hours. This would ensure that temperature effects do not affect to measurements. In practice it might be possible to remove LVDT cell after 2 to 3 hours as, according to numerical simulations, total displacement magnitude halves during the first 2 hours of cooling.

Numerical simulations also suggest that the relation of radial displacement and ΔT is linear during the cooling period. It should be noted this is achieved by constant cooling power in numerical model, which might not represent real world situation, especially at the beginning of the cooling period. However, as the LVDT cell is capable of measuring both radial displacement and temperature, it might be possible that the results could be temperature corrected by utilizing simple linear extrapolation for the collected data. This way the LVDT cell could be recovered from the pilot hole even after 30 minutes to one hour after the overcoring is stopped.

Performed tests and numerical simulations give some indication about the overcoring heat induced displacements. These results are obtained without the LVDT cell itself, which might introduce even more uncertainties to those discussed, such as the elongation of the LVDT sensors due to the overcoring heat and the heat generated by the LVDT electronics. It is thus advised to perform a laboratory test based on the results in this study.

Knowing the approximate temperature distribution within the rock cylinder as function of time, it is possible to study the variation in LVDT values in isolated laboratory conditions. This can be used to estimate the magnitude of error caused by heating using the LVDT cell to measure the displacements. Results of this heating experiment are also of great value when evaluating reliability of the numerical models.

The laboratory heating experiment would consist of the LVDT device, a hollow core rock cylinder and a moving heating cable. The basic principle of the experiment is to measure deformations in test cylinder while heating it from outside with moving heating cable, thus simulating the overcoring process. The heating power is calibrated to induce similar temperature distribution within the test cylinder as measured in overcoring experiments.

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Ask, D. (2003). Evaluation of measurement-related uncertainties in the analysis of overcoring rock stress data from Äspö HRL, Sweden: a case study. International Journal of Rock Mechanics & Mining Sciences , 40, 1173-1187.

Bertilsson, R. (2007). Temperature effects in overcoring stress measurements. Luleå University of Technology, Civil and Environmental Engineering / Rock Mechanics. Luleå Tekniska Universitet.

Cai, M., Qiao, L., & Yu, J. (1995). Study and Test of Techniques for Increasing Overcoring Stress Measurement Accuracy. International Journal of Rock Mechanics and Mining Sciences & Geomechanics Abstracts , 32 (4), 375-384.

Hakala, M. (2006). Quality Control for Overcoring Stress Measurement Data. Posiva Oy.

Hakala, M., Kemppainen, K., Siren, T., Heine, J., Christiansson, R., Martin, C., et al. (2012). Experience with a new LVDT-Cell to measure in-situ stress from an existing tunnel. Eurock 2012. ISRM.

Kukkonen, I., Kivekäs, L., Vuoriainen, S., & Kääriä, M. (2011). Thermal Properties of Rocks in Olkiluoto: Results of Laboratory Measurements 1994-2010. Posiva Oy.

Lin, W., Kwasniewski, M., Imamura, T., & Matsuki, K. (2006). Determination of three-dimensional in situ stresses from anelastic strain recovery measurement of cores at great depth. Tectonophysics (426), 221-238.

Posiva Oy. (2009). Olkiluoto Site Description 2008. Eurajoki: Posiva Oy.

Siren, T. (2011). Fracture Mechanics Prediction for Posiva's Olkiluoto Spalling Experiment (POSE). Posiva Oy. Posiva Oy.

Siren, T., Martinelli, D., & Uotinen, L. (2011). Assessment of the Potential for Rock Spalling in the Technical Rooms of the ONKALO. Posiva Oy. Posiva Oy.

Sjöberg, J., Bertilsson, R., & Christiansson, R. (2008). Overcoring in Deep Surface Boreholes - Recent Experiences and Lessons for the Future. The 42nd U.S. Rock Mechanics Symposium (USRMS) .

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