Background

The need to diversify the energy mix has, among other things, led to attempts to generalise the use of the Earth’s heat for the production of heat and/or electricity. Moving beyond the “classic” geothermal energy stage associated with active volcanic areas and specific aquifers whose hydraulic and thermal characteristics are directly and economically exploitable is a new field of geothermal energy known as the Engineered/Enhanced Geothermal Systems (EGS). These systems use deep underground (high-temperature) rock formations as heat exchangers through promoting the circulation of a natural fluid (MIT 2006). In the absence of a typical aquifer (permeable porous medium), natural fluids circulate in complex hydraulic systems made up of fracture and fault networks that are more or less well connected, depending on the considered scale. The aim of the new exploitation techniques is to provoke fluid circulation between an injection well and one or more production wells, thus modifying the local circulation dynamics in order to obtain an economically viable flow rate and temperature. Given their weak injectivity and initial productivity, the wells in these environments commonly require a development phase based on hydraulic and/or chemical stimulation through overpressurised injection of a cold fluid into the hot fractured medium. During this development phase, predominant physical processes in the fracture network depend on the stimulation scenario: hydromechanical processes are of first order during hydraulic stimulation, whereas hydrochemical processes drive the behaviour during chemical stimulation. Following the development phase, the cost effectiveness of these systems lies in their sustainability over time, i.e. at least 20 years of operation without any substantial reduction in well injectivity and/or productivity or any thermal short circuit due to a localised increase in the permeability of the deep fractured rock mass. During this exploitation phase of the EGS, the problem of permeability evolution in a natural fracture (basic element of the hydraulic system in question) due to fluid–rock interactions, within a varying thermal and mechanical context depending on the distance from the well, is a crucial issue.

For over 20 years, the Soultz-sous-Forêts experimental site in Alsace (France) has been dedicated to the scientific study of these new geothermal systems (Genter et al. 2010). The various research projects carried out at this site have shown the importance of understanding both the natural and the induced circulation of fluids in the fractured and/or faulted granitic basement and its evolution in situations of specific thermal and mechanical stress. This problem, which is relatively new in the field of geothermal energy, is similar to that posed over the past 30 years in connection with the underground storage of radioactive waste (Rutqvist and Stephansson 2003) and more recently in the oil industry for the exploitation of fractured reservoirs.

The evolution of a fracture’s hydraulic behaviour under both normal stress and shear has already been widely studied in relation to the storage of nuclear waste in the 1980s and 1990s. The experimental studies resulted in the relationship between hydromechanical behaviour and fracture morphology being taken into account more or less explicitly and in more or less detail. The morphology of the walls and their degree of match, quantified through various approaches, has proved to be an important parameter as regards both normal stress (evolution of the contact surfaces with increasing stress) and shear (evolution of the friction surfaces according to their “angularity”). Over and above understanding the mechanical behaviour, the evolution of the fracture morphology determines the evolution of its permeability in expressing the deformation of the hydraulically effective volumes. Conversely, the thermomechanical behaviour of fractures has been little studied and remains a completely open and critical issue for understanding phenomena in the EGS context, independently of the thermal stimuli themselves.

The evolution of a fracture’s permeability as a result of fluid–rock interaction is an even more recent problem and is still little studied from either the experimental or the modelling standpoints. Although coupled models are beginning to be developed from a theoretical or empirical point of view, their validation is far from being realized, notably due to the lack of a sufficient number of adequately instrumented laboratory experiments to take into account the different interfering hydraulic, mechanical and thermal aspects. This validation requires, among other things, an understanding of the chemical phenomena occurring within the fracture and the changes in the fracture’s morphology. A number of recent studies (Polak et al. 2003; Yasuhara et al. 2006; McGuire et al. 2013; Zhao et al. 2014) are devoted to chemical interactions in fractures and, in particular, to the different types of mineral dissolution patterns. Two types of dissolution phenomena are identified: those of the free face dissolution type and those of the pressure solution type. The latter correspond to chemical corrosion of the fracture asperities in contact resulting from the localised concentration of pressure at these points leading to greater mineral solubility (Zhao et al. 2014). The dominance of this pressure solution effect is notably due to mechanical loading and the effective pressure applied to the fracture (McGuire et al. 2013). The two types of dissolution phenomena may impact differently on the hydraulic behaviour and permeability of the fracture: the channelling effects induced by free face dissolution and giving rise to increase in the fracture’s permeability can be offset by the chemical attack on the fracture asperities in contact which, conversely, brings about a decrease in the hydraulic opening and thus the permeability (Polak et al. 2003). When the fracture is at the same time submitted to a normal stress, a mechanical closure is superimposed to these closures/apertures due to chemical phenomena. Several relevant processes can explain the mechanical closure of a fracture under normal stress. Indeed, in addition to its elastic part, an irreversible closure of the fracture can occur due to mechanisms such as damage by brittle fractures or plastic flow of contacts (Brown and Scholz 1986), viscous creep of the contacts inducing time-dependent closure (Matsuki et al. 2001) or stress corrosion process (Yasuhara and Elsworth 2008). This latter process induces a compaction of the fracture by combining mechanical and chemical phenomena: the tensile stresses resulting from the compressive loading of contacting asperities induce “subcritical” or “quasistatic” cracking in the matrix around them and the presence of fluid can lead to a growth of these fractures due to chemical reactions. In addition to these processes, when the fracture is submitted to temperature, a thermal over-closure can occur (Barton and Makurat 2006). All these processes can superimpose and the predominance of one or another will depend on the loading conditions (level of normal stress, temperature), the type of rock, the fluid composition and the morphology of the fracture. As far as the latter is concerned, numerous experimental works are carried out with “artificial” in lab-made fractures. Depending on the way they are created, such “fresh” fractures will be free of coating and mineral deposits on their wall and their asperities will have a high angularity. When the fracture is submitted to a normal stress, this point can increase the stress concentration and then intensify the irreversible mechanisms such as damage, stress corrosion and pressure solution. In contrast, natural fractures can exhibit coating due to their history—fluid circulation or chemical phenomena—and their asperities can be rounded, which can limit stress concentrations.

Using experience gained from laboratory studies of a fracture’s hydromechanical behaviour under normal stress and preliminary developments for studying the effect of acidification under controlled temperature, we built and gradually improved an experimental apparatus so as to provide a percolation cell in a fracture under imposed stress and temperature conditions (Gentier et al. 1998). This apparatus, after various stages of adaptation and experimental validation, enables us to study the complex physical phenomena that can play a fundamental role in the success of EGS.

In addition to describing the experimental apparatus and the first results obtained on the evolution of fracture permeability, we aim to highlight the need for a strict upstream experimental protocol for the reactive percolation so as to enable meaningful interpretation of the results. The work presented here was carried out on a natural fracture in a core obtained from a depth of 1890 m in drill-hole EPS1 at Soultz-sous-Forêts.

Methods

The aim of the in-fracture reactive percolation tests under an imposed normal stress and temperature is to characterise the evolution of the fracture’s hydraulic and hydromechanical behaviour with, in the first instance, the evolution of its permeability induced by the fluid/rock interactions resulting from the injection of an imposed chemical fluid. To achieve this aim, the work carried out over the last 10 years has been focused on develo** the experimental apparatus and on establishing a methodology for obtaining the necessary data and information through a series of tests and characterisations carried out before and after the reactive percolation test itself.

Principle of the tests and experimental apparatus

The tests are performed on cylindrical samples of fractured rock cored such that the mean plane of the fracture is perpendicular to the cylinder’s axis of symmetry. The principle of the tests is to percolate a fluid through a natural fracture contained in a rock sample, under imposed and/or controlled THM conditions. The fluid, of known and constant chemical composition (the percolated fluid not being recycled), is injected into the centre of the fracture by boring into the lower wall, resulting in a divergent radial flow within the fracture. The evolution of the fluid’s chemical composition is then characterised after passage through the fracture. The tests were performed within a containment cell at an imposed temperature and with a normal stress loading on the sample perpendicular to the fracture plane.

The experimental apparatus is shown schematically in Fig. 1:

Fig. 1
figure 1

Diagram of the experimental apparatus

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  • The hydraulic part of the test consists in injecting the fluid at a prescribed flow rate with a chromatography pump, which also measures the injection pressure. Once the fluid has percolated through the fracture, it is recovered in an annular reservoir surrounding the rock sample. With the containment cell being pressurised (0.1 MPa), each time the solenoid valves located outside the chamber (at atmospheric pressure) are opened, the pressure difference flushes the fluid from the annular reservoir via capillaries (with a diameter of 1/16th of an inch) to the systems for measuring the physicochemical changes in the fluid, for collecting samples and for monitoring the outflow. The last, which is done by weighing the fluid, enables one to verify, for an imposed injection rate, that there is no fluid loss in the circuit whether through leakage or excessive evaporation. To limit evaporation, the nitrogen used for pressurising the cell is water saturated through bubbling.

  • For mechanically monitoring the tests, a force-controlled press is used to apply a force in the axis of the cylindrical sample and thus load the fracture under normal stress. Four displacement sensors distributed around the sample measure the relative displacements of the walls. These measurements include both the deformation of the walls and the closing/opening of the fracture; however, for rocks whose matrix can be considered as poorly deformable in terms of the applied normal stress level, these measurements provide direct access to the opening or closure of the fracture. This has to be checked considering the Young’s modulus of the rock matrix and the level of normal stress applied.

  • For thermally monitoring the tests, a temperature-controlled heating resistor installed on the wall of the containment cell enables one to regulate the cell temperature. The apparatus does not guarantee a uniform temperature within the containment cell, which is why the temperature is measured at various points using PT100 probes distributed over the height of the cell (bottom, middle and top), and using thermocouples for measurements at the rock sample and fluid contacts (i.e. at the point of injection into the fracture, at the contact with the fluid in the annular reservoir and at the level of the upper wall rock matrix).

  • For monitoring the physicochemical evolution of the fluid, the experimental apparatus enables both online monitoring of certain of the fluid’s parameters after its passage through the fracture and sampling of the fluid in order to perform targeted analyses of the chemical elements required for monitoring changes in the fluid’s chemical composition (the chemical composition of the injected fluid being known). The fluid’s pH and Eh are measured online using pH/Eh sensors at the exit of the containment cell, thus enabling the reactivity of chemical processes to be checked. The sampling is done by an automatic fraction collector that extracts the fluid volumes required for analyses undertaken in suitable packaging (open tubes, vacuum-sealed tubes) and according to a sampling frequency adapted to the test’s reactivity.

Methodology

To assess the fluid/rock interactions and their influence on the evolution of the fracture’s permeability, we developed a methodology associated with the operation of the experimental apparatus. It is based on a set of morphological, petrographic, chemical and physical characterisations and on hydromechanical behavioural tests under normal stress and temperature (summarised in Fig. 2) carried out before and after the reactive percolation test.

Fig. 2
figure 2

Methodology of the percolation tests. The red arrows indicate the chronological sequence, the green ones mean “induce” and the thin grey ones indicate which data or characteristics are compared in order to determine the evolutions referenced on the right

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To determine the fracture’s petrographic, physical, chemical and morphological characteristics, we

  • analysed and described the minerals on the fracture walls and in the rock matrix;

  • mapped the fracture’s voids (Gentier and Billaux 1989) by analysing images of the void casts;

  • mapped the topography of the fracture walls by profiling with a laser profilometer; and

  • mapped the chemical elements of the fracture walls using X-ray microfluorescence.

The compositions of the fluids used during the characterisation phases of the fracture’s hydromechanical behaviour are determined after the characterisations and before the reactive percolation test (Fig. 2). Moreover, to avoid any fluid/rock interactions during these phases, the fluids are determined so as to be chemically inert with respect to the sample’s minerals, whether in the matrix or on the fracture walls.

Once the sample is placed in the experimental apparatus, a loading protocol is applied so as to rematch the fracture walls. The protocol consists in carrying out cycles of mechanical loading/unloading followed by a hydraulic test. At the end of each loading/unloading cycle, the mean residual irreversible displacement due to the cycle is determined from the LVDT displacement sensors. The mechanical criterion of the rematch is when the mean irreversible displacement after a cycle tends to zero. For the hydraulic test, a given flow rate is imposed following a mechanical cycle and the injection pressure is measured. The hydraulic criterion of the rematch is when the pressure given by two successive hydraulic tests is constant. The fracture’s rematch is considered effective when both the mechanical and hydraulic criteria have been achieved. Once the fracture has been rematched, and to ensure this state throughout the test, a minimum normal stress, termed pre-load stress, is applied continuously to the sample. In order to avoid/limit any mechanical damage of the asperities in contact during the rematching, the maximum value of normal stress applied will be determined by considering the in situ normal stress submitted by the fracture.

The mechanical, hydraulic and hydromechanical behaviour of the fracture is characterised through injection tests. By injecting chemically inert fluids at different rates under several normal stress levels, one can characterise the fracture’s closing/opening and the evolution of injection pressure versus flow, and also estimate the flow regime within the fracture.

Acquisition of the fracture’s morphological features

The mechanical behaviour and flow properties of fractures depend largely on the surface roughness of their walls and their match (Barton and Choubey 1977; Gentier et al. 2000; Crandall et al. 2010): the walls of natural fractures are surfaces with ripples of different wavelengths that can be as much as the asperities directly associated with the minerals or component elements. In particular, Hopkins (2000) highlighted the determining role of the contact zone characteristics (shape, size, number, distribution and resistance) on the fracture’s mechanical properties and of the structure of the free space between the walls (void space) on the hydraulic properties.

To quantify the potential changes in the void space and contact zones following a reactive percolation, the roughness at sample scale is depicted by topographic maps of the two fracture walls and by a thickness/height map of the voids. The combination of the three maps enables the fracture morphology to be characterised by magnitudes derived from statistical and geostatistical calculations. To enable both a spatial repositioning and a superpositioning of these maps, three cylindrical Teflon inserts, one millimetre in diameter, are embedded in each of the walls, with spatial correspondence of the points once the two walls have been rematched. The inserts are unique points, both chemically and topographically, that are easily identifiable on the different maps.

Map** the wall topography

The most classic methods for analysing fracture surface topography are based on topographic profiles (mechanical [sensor or needle] profilographs, optical profilographs [such as light section microscopy], interferometry, speckle metrology and laser profilometry; Maerz et al. 1990; Ge et al. Full size image

Map** the void thicknesses (or heights)

Estimating the thickness (or height) of the voids left after matching the walls is difficult. Recalibrating the topographic data acquired on each of the walls is problematic due to the thickness of the very small voids compared to the different undulations and irregularities of each wall. From a mechanical standpoint, some authors have limited themselves to estimating the contact zones: area and distribution. Among these methods are those based on the distribution of temperatures using thermocouples (Teufel and Logan 1978), on resistance to the passage of an electric current from one wall to the other (Power and Hencher 1996) and on the impression of the contact zones obtained with pressure-sensitive paper (Duncan and Hancock 1966) or deformable film (Iwai 1976; Bandis et al. 1983). Methods of injecting the fracture using a metal alloy with a low melting point, such as Wood’s metal (Pyrak-Nolte et al. 1987; Yasuhara et al. 2006), also restrict the information obtained to binary data: injectable or contact zone. Furthermore, it does not enable reuse of the fracture after injection.

From a hydraulic standpoint, knowledge of the zones in contact and otherwise is not really sufficient because the potential interactions that can affect the contacts can also modify the existing voids. The morphology of the void space can be obtained (1) through X-ray tomography as used by Keller (1998) and Re and Scavia (1999), although this does not allow the detection of voids less than 0.5 mm thick, (2) through injecting a coloured Epoxy-type resin (Gale 1987), which involves destruction of the sample, or (3) through injecting a soft coloured resin (Gentier et al. 1989) allowing the mould to be removed cleanly from the fracture, which is then reusable. The last method, which was used in this study, enables multiple casts on the same sample and, more specifically, both before and after the reactive percolation. Moulding the voids is done by expulsing the coloured silicon resin from the fracture previously filled with the fluid resin when adjusting the two walls. The resin’s colour is adapted to the range of the fracture’s widths; the thicker the mould, and thus the thicker the void, the darker the resin. At the same time, a calibration wedge with a bilinear thickness variation is moulded with the same resin preparation for calibrating the relationship between colour and thickness (Fig. 4). Images of the void moulds and the associated calibration wedge are obtained by light transmission using the same protocol; the resulting images are 16-bit coded RGB images (that is 65 536 or 2562 unique values for each red, green and blue component) in which each pixel corresponds to a square of 35 μm sides, providing a horizontal resolution of the same order of magnitude as the topographic maps of the walls (25 μm). The images are corrected for non-uniformity of the light source so as to eliminate any bias in capturing the transmission images. The image of the calibration wedge is then analysed in order to assign a thickness to each grey-scale value, coded on 2562 unique values, based on which the image of the void moulds is transformed into a void thickness map (Fig. 5). Studying the dispersion of the grey levels for each void thickness against the image of the corrected calibration wedge mould made it possible to estimate the thickness error as ±10 µm, which is identical to that of the topographic maps of the walls.

Fig. 4
figure 4

Calibration between colour and thickness. a View of the calibration wedge (exaggerated thickness); b plan view of the calibration wedge; c theoretical curve giving the thickness of the calibration wedge according to distance from the wedge’s first graduation

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Fig. 5
figure 5

Map** of the void thicknesses. a View of a void mould (exaggerated thickness); b curve for calibrating void thickness from the grey levels in the calibration wedge and void mould images; c map of the void thicknesses before the percolation test

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Rotating the void thickness maps made it possible to superpose them and study the differences before and after the percolation test, as well as to align them with the topographic maps of the walls based on the position of the reference pins.