C Didier

J P Josien

Keywords: In Frankreich stellt sich derzeit das Problem der definitiven Schliessung des meisten Bergwerke. In manchen fällen ist die Stabilität der Oberfläche nicht garantiert, speziell wenn unter Tage nicht verfüllte Hohlräume verblieben sind. Dieser Bericht zeigt die Wichtigkeit, bei der Auswahl geeigneter Langzeit-Risikomanagementmethoden, der verschiedenen Mechanismen mit denen ein eventueller Bruch an die Tagesoberfläche geraten kann

France now faces the closure of most of its mining sites. Surface instability risk may sometimes remain, especially where large underground voids exist. This paper illustrates the importance of failure mechanisms identification in order to define properly how to manage the residual risk. RESUME: La France fait aujourd'hui face à la fermeture de ses exploitations minières. Des risques d'instabilités de surface peuvent persister, notamment lorsque des vides résiduels existent au sein des travaux souterrains. L'article illustre l'importance que présente la nature des mécanismes de rupture potentiels dans le choix des mesures compensatoires à adopter pour gérer le risque résiduel à long terme.

Importance of failure mechanisms for management of surface instability risk above abandoned mines.

Introduction

After exploiting its underground mineral resources for centuries, France has witnessed the gradual closure of its mining sites. These closures have often taken place without any satisfactory analysis being performed to identify the possible consequences which could have long-term impact on persons, property and the natural environment. Of the various risks inherent in a post-mining period, the surface instability may be particularly significant where large residual voids remain underground once mining has ended. This "ground movement" risk can be evaluated by combining the predictable damage and surface occupation. It is used to judge the need to envisage the implementation of appropriate compensatory measures. Treatment such as backfilling or providing support in voids are indeed often technically and economically impossible because of the volume, depth and accessibility of the mining works concerned. Although priority is traditionally given to solutions for managing the surface occupation in sectors which are still uninhabited, when the risky zones already show surface instability, this solution is no longer adequate and moving populations would generate sociopolitical and financial problems which are difficult to solve. When there is no other suitable solution in the context, monitoring potentially unstable zones may provide an interesting alternative. This solution is intended to provide risk management by detecting the signs of unfavourable evolution of the undermined land. In this way populations can be maintained in a "doubtful" zone while making it possible to evacuate them in the event of an alert judged sufficiently serious by the competent authorities. Using a monitoring system requires the failure mechanism, particularly its dynamics, to be identified as accurately as possible, so that it can be determined whether it is possible to detect preliminary signs and whether they are adequate for operational risk management (trigger a warning). In-depth investigation must therefore be conducted on the type and dynamics of failure mechanisms liable to affect mine working and to spread to the surface (gradual or sudden, smooth or discontinuous phenomena). In this context, INERIS and GEODERIS have recently developed techniques and tools for risk analysis and management concerning mining subsidence for various mining contexts. Two examples of mine fields in the process of being closed, for which residual risks have been identified, will illustrate this article. For the Lorraine iron basin, with many examples of instability in the past, it was possible to determine methods based on back analysis to characterise failure mechanisms and classify the risky zones on the basis of a multicriteria analysis. For the Provence coalfield, on the contrary, the very small number of instabilities in the past means that a deterministic approach has to be used, based on numerical modelling of particularly complex failure mechanisms.

The Lorraine iron basin

Context and objectives:

The Lorraine iron basin extends 120 km from North to South and 20 km from East to West in North-eastern France. It covers more than 1,500 km². The iron deposit outcrops in the North and East and falls with a shallow 3% dip towards the West. The depth of mining operations varies from zero to 260 m, a range of depths particularly sensitive for surface instability effects. The age of the iron deposit is Aélenien with thickness varying from just a few to about sixty metres. It takes the form of alternating iron layers, 3 to 7 metres thick, alternating with marly or limestone beds. This very rich deposit has been exploited in several layers (up to 6), sometimes with very thin middlings (up to 1 m) which must be taken into account when determining mechanisms leading to degradation of mining underground structures. The overburden consists of several Bathonian and Bajocian marl and limestone beds, some of which show major lateral variation in facies. Depending on the thickness and the facies of the overburden, its behaviour may be relatively plastic (metric limestone bed) or brittle (decametric bed). Production, which began in the Middle Ages on outcrops, grew industrially between 1870 and 1997, when the last mine was closed. The quantity of ore extracted is estimated at 3 billion tonnes (or more than 1 billion m 3 ). The ore was extracted using room and pillar method with rooms varying in time from 4 to 7 m and pillars of varying shape in widths of 5 to 25 m. Extraction ratios are very variable from 30 to 80%. The pillars could then be reworked, stoping leading to collapse of the recovery and consequently surface subsidence. This second operation was not used in certain zones, particularly where the mining company wanted to protect vulnerable surface structures (residences, roads, railways). In this case, remains a question concerning the long-term stability of the pillars left in place underground. The Briey-Longwy basin concerns 110 villages covering an area of 110,000 ha. 50% of the area of these communities has been affected by mining operations. Variable zones, where long-term stability is not proved, cover 8,600 ha of which 1,300 ha have house buildings and 800 ha have infrastructures.

It is impossible to eliminate risk over the entire region; it would require the use of one hundred million cubic metres of backfill, which is technically and financially impossible, or the transfer of population to an extent, which is impossible at regional level. Since the number of instabilities fortunately remains small (about 1/year), risk management has been organized on the basis of the following principles: Guarantee personal safety, damage to property giving rise to compensation; Classify the risk, taking into account the importance of surface effects and the likelihood of collapse of mining structures;

Set up a monitoring system to detect preliminary signs of failure and make sure people are moved to safety;

Reserve risk management to zones where safety cannot be guaranteed. This risk management involves accurate knowledge of the nature and amplitude of possible instabilities and the mechanisms responsible for these instabilities so as to define appropriate monitoring and classification of the different zones according to their risk level. This information has been obtained by retroanalysis of old (during exploitation) and more recent events.

Surface instabilities

The shallowest zones are likely to be affected by sinkholes of several metres wide and several metres deep. The conditions of appearance of this type of instability are well known. Under natural conditions and those governing ironore mining in Lorraine, mining at a depth of more than 50 m eliminates the risk of sinkhole development.

For mine workings at greater depths, deterioration will be reflected in subsidence, the surface formation of a depression which may be several hundred metres in diameter. The maximum subsidence at the centre of the depression is proportional to the voids left in mining works (working height x extraction ratio); the coefficient of proportionality depends on the width of the zone mined. Surface subsidence is almost non-existent for a width less than 0.4 time the depth (H), remains very slight up to 0.6 H, when it begins to accelerate rapidly. This is interpreted as support given to the roof by the abutments of the mined zone, limiting subsidence. In some cases, surface phenomena no longer occur as more or less rapid formation (one day to several months) of a subsidence depression. They occur suddenly, in just a few seconds and are accompanied by a seismic shock. The rocks fall in parallel ranks above the mined zone with the formation of open fractures in a step formation on the edges. Eight cases of sudden collapses (called generalised collapses also) were recorded between 1902 and 1974. Some of these occurred during operations, others when mining had been completed for many years in the zone.

Mechanisms of instabilities

The appearance of a surface instability results from a process of degradation which has affected mining operations and the overburden. Several mechanisms have been observed. In single level mining, it is mainly the pillars which are the weak point of the structure. One zone will be characterised by the mean stress in the pillars calculated by the tributary area. Stresses in pillars are not uniform throughout the mined zone, either because of the variation in their dimensions, or because of their position (overload linked to adjacent stoping, for example). Pillar rupture can cause the hanging wall to collapse immediately, but for an instability to be reflected on the surface, the affected zone at the bottom must be sufficiently wide. The main hanging wall will collapse when the area is more than 0.6 time the depth. Pillar rupture leads to a process which, by spreading the stress, overloads the adjacent pillars. The process duration depends on the pillar stress with respect to their strength, but it is hard to believe that it could be instantaneous for a slightly fragile rock such as the Lorraine iron ore. In a multilevel operation, stability must be assessed for all mined layers and middlings. Thin middlings deteriorate in the form of roof falls or crushing failure. Rupture of middling weakens the pillars: decrease of its w/h (width/height) ratio, formation of a composite pillar formed by the various layers and middlings. Impact on stability is particularly high when the pillars in the different layers are not exactly superimposed. These phenomena are difficult to quantify during a global stability study, particularly because of the importance of detailed geometry in these workings. To be on the safe side, for every middling less than 7 m thick, the tributary area is only applied to the part of the pillars that appears after superimposing the planes of the 2 layers. These degradation mechanisms are not appropriate to explain sudden collapses. Indeed, the description of phenomena which occurred during mining operations indicated the apparently good condition of the workings during the months preceding the collapse. Precursor signs such as increased pressure, flaking pillars, were only reported during the very last days preceding the event. The presence of thick, rigid beds in the overburden, in all the cases analysed, means that a mechanism can be proposed to explain the sudden rupture of overburden and pillars within a few seconds. These beds help stability by bearing part of the overburden load, thus relieving the pillars, which explains their good condition before failure. Over a period of time, creep acting on the pillars increases the stresses inside the rigid beds which cannot follow this deformation. The rigid beds fail, leading to a simultaneous increase in load on all the pillars in the zone and causing their simultaneous failure up to the surface, sometime over several hectares. The sudden collapse mechanism involves two essential conditions [START_REF] Tincelin | Sudden collapses, rockbursts[END_REF] :

The presence of a rigid bed (thick and strong) in the overburden "Brittle" mining geometry which will deform and then suddenly break when the bed fails; it is characterised by high extraction ratios, large areas of mining operations, small, not very strong, slender pillars.

Distinction between the risk of subsidence and the risk of collapse.

The factors defined above within the context of a sudden collapse mechanism must be quantified in order to define the criteria used to classify a risk zone. A back-analysis has been done of 16 events, 8 sudden collapses and 8 cases of gradual subsidence (Table 1). The presence of thick, strong beds in all these cases did not enable this criterion to be quantified so the analysis was based only on the mining works geometry. Sudden collapses are characterised by a higher level of mean stress on pillars (14 MPa vs 10 MPa), a lower w/h ratio taking into account all the layers worked (0.9 vs 2.0) and higher openings (7 m vs 4.5 m) but only the extraction ratio really discriminates between the 2 phenomena. Sudden collapses all have an extraction ratio greater than 60%, gradual subsidence all have a level below this limit. Using statistical techniques, a linear function of the operating parameter used to discriminate between the 2 types of instability can be defined. Taking into account existing links in our sample between different variables (for example, the w/h ratio of pillars is strongly correlated to the total opening), some variables only have been selected. Of the 11 parameters known for each case, 7 have been selected as giving the best discrimination: extraction ratio, stress on pillars, depth, hydraulic diameter (ratio between the area and the perimeter of the pillars) width of rooms, presence of adjacent stopes [START_REF] El | Towards the determination of surface collapse type over abandoned mines in the Lorraine iron basin[END_REF].

Applying these criteria to risk zones beneath buildings (1,024 ha) has enabled a level of risk of gradual subsidence to be allocated to 1,021 ha. There are 103 ha remaining for which the risk of sudden collapse is not excluded by this statistical analysis. For these zones a geological study of the recovery is performed to detect whether or not there is a rigid and brittle layer.

Methodology of risk management

For zones likely to be affected by gradual land subsidence, the risk level is classified on the basis of the surface factors vulnerability, the importance of effects likely to be felt and the susceptibility of bottom workings to collapse [START_REF] Piguet | Contribution of Rock Mechanics for Risk Assessment in abandoned mines[END_REF]. The classification method is based on multicriteria analysis techniques, particularly the Electre method. This method, having been adapted to a mining context, has been systematically applied to the whole Lorraine iron basin. Risk management has thus been implemented, based on the following principles.

For zones without buildings or infrastructures, the zone is classified as unsuitable for construction while waiting for the " Mining risk prevention mapping " to be drawn. Those reglementary tools will define the terms for land usage according to risk types and levels as well as social factors.

For zones with risk of gradual subsidence, monitoring is implemented for the highest levels. This monitoring is mainly based on microseismic measurement of ruptures in rocks, made by sensors installed below the surface. This method has been qualified and calibrated by a site test when the last mine to be worked in the field was still in operation.

Seven networks are currently in operational service. Three levels of alarm have been defined according to the process of degradation of mine workings: local degradation of pillars or intervening layers, pillar rupture or roof cave-in, evolution of damage in the overburden. The adequacy of alarm criteria was verified at a site which was still accessible and where the evolution of degradation was noted at the bottom [START_REF] Bennani | Subsidence monitoring in post-mining: needs and methods[END_REF]). An alert procedure was introduced which, depending on its level, mobilises the DRIRE (regional technical management body), the Prefecture, State departments, Town halls and populations concerned depending on the specific emergency plans.

For zones in which the risk of sudden collapse is not excluded, monitoring has also been set up. However, in the context of the Lorraine iron basin, the operational efficacy of this monitoring system has not been proved. Indeed, the particular part played by the rigid bed in the shear mechanism is likely to change the alarm criteria. Considering the short period during which preliminary events precede the collapse, a rapid alert is necessary. The probability of a false alarm is difficult to manage in a longterm monitoring system. Specific experiments concerning the phase of overburden rupture would be necessary to calibrate the monitoring system in each particular case.

The Provence coalfield

Geological and mining context

The Provence coalfield is located in South-Eastern France, close to Marseille. With a mainly East-West orientation, it extends over a distance of 70 km long and 12 km wide. The coal deposit is found entirely in a massive Upper Cretaceous limestone formation, close to 200 m thick and consisting of continuous beds, often very hard. It consists of seven layers of varying thicknesses, of which only the 4 deepest have been largely mined, mainly the thickest, known as the "Grande Mine", the thickness of which varies between 2 and 5 m according to the sector of the field.

Mining began during the 16th century, and the last workings were closed at the end of January 2003, taking the total area of undermined land to slightly less than 200 km². For reasons of simplification, it will be considered that until 1930 the Grande Mine vein was worked using the room and abandoned pillar method (area of more than 3,000 ha). The extraction ratio, , close to 75%, has varied very little with time and depth (approx. 0 to 500 m). The most recent and deepest (500 m to 1,300 m) workings were mined using the longwall caving method and do not pose any major problems in terms of surface stability. Several sources of information (oral surveys, aerial photographs, archives) have revealed the low level of the Provence field's susceptibility to major instabilities on the surface. Over the past two centuries, only one major event has been discovered throughout the field: a massive collapse of several hectares in 1879. Several exploratory visits have been made to previous sections which are still accessible and were mined using room and pillar methods. Observations revealed a gradual yielding of pillars leading to significant deflection of the limestone roof with convergence values approaching one metre (Figure 1). Although these observations help eliminate the risk of dynamic pillar failure (there is no elastic energy available in the supports), on the other hand they raise questions on the ability of the rigid overburden to absorb such large deflections without rupture.

To obtain more calibration data, a borehole and logging campaign was undertaken in various sectors of the field now inaccessible, with the intention of characterizing the height of residual voids in the centre of panels judged to be representative and located at various depths.

Choice and development of the model

Simple analytical modelling of pillar stability (tributary area) indicates that pillars located deeper than 100 m have a safety coefficient of less than 1. Although this diagnosis is validated by observations (collapsed pillars, high convergence) made in relatively deep sections (approx. 350 m), it is refuted by inspections made on workings at medium depth (150 m to 250 m), which reveal an absence of convergence and hence of collapsed pillars. Such a different behaviour cannot be explained without taking into account the interaction between deformable pillars and a rigid hanging wall, which plays an essential part in the behaviour of the Provence coalfield rock mass. Indeed, the rigid overburden acts like a slab supported on the unmined edges and, thereby, relieving the subjacent pillars, which, on the other hand, make the hanging wall susceptible to the risk of failure due to deflection. Only numerical modelling of the rock mass on a global scale of mining panels can take into account " overburdenpillar" interaction. The deformation and failure mechanisms thought to occur in the Provence field are strongly influenced by the presence of heterogeneities, particularly stratigraphic. The calculation code chosen is then UDEC, as this tool allows one to take into account the existence and behaviour of the numerous discontinuities in rock masses.

Faced with the complexity of the environment and mechanisms to be modelled, a "pragmatic" approach aiming to give priority to calibration using the available operational data has been chosen. We then built a robust model able to reconstruct accurately the mechanisms observed or suspected, while retaining enough flexibility to be adapted to the various configurations encountered.

For each representative sector, a parametric analysis (mainly based on the parameters of the law governing the behaviour of coal) was used to calibrate the model on the basis of information collected during the exploratory phase (residual opening of workings). The current state can then be extrapolated to evaluate the future behaviour of undermined land.

We selected an elastoplastic softening behavioural model to model the limestone and coal. This model is used to locate meshing elements for which the plastic criterion has been reached (plastic points). In fact, the criterion for major collapse of a panel is considered to be due to a combined high density of plasticity points affecting most of the roof beds, and significant lowering of the overburden.

Main results of simulations

Sectors at medium and extreme depth (beyond 250m)

The results of many calculation simulations reveal that at medium or extreme depth, the coal pillars were, from the time of extraction (mechanical characteristics not reduced by the effect of ageing), subject to stresses too severe to be withstood without being crushed. The crushed pillars caused deflection then gradual rupture of the overburden (Figure 2). This rupture is reflected by a smooth profile on the surface (subsidence with no effect on structures).

The results of the calculation match the testimony of exminers who describe multidecimetric subsidence of the workings from the first days following excavation. The workings therefore behave more or less identically to a classical longwall exploitation with a "overburden degradation face", with the difference that the overburden did not collapse but was gradually deflected leading to the creation of fractures in the hanging wall beds. Calculations show that only major plastification of the overburden can produce a level of convergence as high as that observed in the underground workings (up to one metre).

In this configuration, once the panel has been worked and subsidence has occurred, there is no more elastic energy stored in the pillars (crushed and in a state of post-rupture), nor in the hanging wall (fractured and resting all its weight on the subjacent pillars). Any risk of sudden rupture of the overburden in the long term can therefore be eliminated. On the other hand, continuing (or further) smooth subsidence is still possible in the event of residual crushing of the collapsed pillars.

Shallow sectors (up to 250 m)

Figure 3 illustrates the evolution of convergence in shallow workings in accordance with the gradual reduction of characteristics attributed to coal pillars. In spite of the pertinence of this hypothesis, using a safetyoriented approach imposed by the presence of surface factors, it is difficult to categorically exclude any evolution in the sectors concerned toward general "hanging wallpillar" failure. The main problem is, as in Lorraine, to identify the dynamic or progressive nature of instabilities.

The experience feedback indicated that almost all surface movements affecting the Provence field have been smooth and gradual. Only one example of massive collapse has been noted for the whole field. This collapse, characterized by a fairly specific context (overburden affected by several faults), developed in the configuration identified as being most susceptible to the risk of major movement (room and pillar mining at low or medium depth).

In addition, although it is not enough on its own to determine the dynamics of the phenomenon, modelling provides some important information on rupture development. Although, in deep sections, pillar crushing and hanging wall degradation developed during mining operations and followed the extraction face, the same does not apply to shallow sections which have not suffered any major convergence. In this case, it is the eventual subsidence of the rock mass due to the ageing effect which could be the source of instability. This type of degradation would not be propagated gradually in space like a longwall, but would simultaneously affect all the workings, causing general deflection of the rigid roof with a risk of rupture along the abutments, which would transfer the weight of the rock suddenly onto pillars not dimensioned to support such stress. Therefore, in the interests of safety with respect to the surface factors, it seems to us to be reasonable not to categorically exclude the possibility of sudden movement, even if it is unlikely and even highly unlikely that it would actually occur. Appropriate measures in the context of risk management will therefore be defined and implemented.

Conclusions -prospects

Within the context of studies on a regional scale, as it is the case of the mining fields used to illustrate this article, detailed analysis of each panel is inconceivable and global methods of diagnosis must be defined. The two examples have clearly revealed the importance of accurate identification of failure mechanisms and the dynamics of any instabilities. In this context, the search and implementation of available experience feedback must be systematized. Sorting and interpreting the instabilities developed in the past, when there are enough of them and accurate information is available, do indeed provide valuable information on the type and occurrence of instabilities. They can also help, as was possible in the case of the Lorraine iron basin, to define rupture criteria, usually based on a combination of operating parameters (recovery factors, pillar slenderness ratios, superposition of workings, panel dimensions…) and geological data (sedimentary deposit or vein, type and thickness of the overburden).

When there is not enough well-documented experience feedback or if the mechanisms involved are particularly complex to understand and define in terms of criteria, "large scale" numerical modelling methods applied to mining operations probably provide one of the most promising prospects as long as the time and financing needed for the model calibration phase (exploratory work, material characterisation…) are available.

The calculation codes available, in constant development, and based on the capacities of increasingly powerful computers have made considerable progress over the past twenty years. Rock mechanics engineers are now excellent at defining criteria so that a working configuration offers guarantees of long-term stability.

However, there are still some major gaps in interpreting the dynamic nature of failure mechanisms which may occur and experts, required to make long-term forecasts on possible instabilities (date and occurrence of damage), soon reach the limits of the current state of knowledge and available calculation tools. Dynamical analysis of failure (essential in terms of identifying the potential danger for people exposed to it) is difficult to envisage as long as the mechanism of initiation and propagation of rupture is not better understood and reflected in the available constitutive models. This type of progress would raise the hope of using high performance dynamic modelling methods instead of the "almost statistical" methods used today.

Common research (rupture theoreticians, rock mechanics engineers, numerical analysts…) based notably on ambitious in situ experimentation and operational tools development will lead to significant progress in the field of understanding rock instability phenomena. These tools would help to overcome the still large uncertainty and to limit recourse to the principle of precaution, still practically systematic today and sometimes excessive, with its eventual consequences in terms of the development of undermined land.

Figure 1 :

 1 Figure 1: Photograph of a collapsed coal pillar.

Figure 2 :

 2 Figure 2: Deep panel with ruined pillars and shear of the overburden.

Figure 3 :

 3 Figure 3: Evolution of pillar convergence in a panel as a function of coal cohesion values.Analysis of the result of these simulations (in the knowledge that at present the sites concerned have not experienced any definite convergence), combined with the analysis of the result of an identical procedure carried out in deeper sections (where major convergence has already occurred) is used to identify the equivalent characteristics of materials most appropriate for the model developed.It has therefore been deduced that, in the medium term, to reconstruct site degradation mechanisms accurately, the most satisfactory equivalent value of coal cohesion is close to 1 MPa. We have decided to vary only coal cohesion, the internal angle of friction, close to 30° being judged less susceptible to degradation within time. It will be noted that this couple of values (c=1MPa, =30°), obtained by simple retroanalysis within the context of the model, corresponds to a range of in situ resistance which is absolutely classic for altered French coal (R C around several MPa).It is then possible to extrapolate the very long-term behaviour of the sites by reducing pillar characteristics, distinctly but credibly (possible effects of rock ageing phenomena). Simulation of the degradation of coal characteristics between c = 1 MPa and c = 0.5 MPa reveals a very distinct difference in behaviour (Figure4).

Figure 4 :

 4 Figure 4: Simulation of the behaviour of shallow workings for c = 1 MPa and c = 0.5 MPa While c = 1 MPa confirms the current level of stability of the workings (stability provided by the "slab effect" of the rigid hanging wall and not the pillar dimensions), using c = 0.5 MPa reveals the appearance of rupture points on the edge of the panel. Several scenarios can then be envisaged concerning the long-term evolution of these configurations.The first considers that the workings are stabilized because there is a very little reduction in the resistance of the rock mass, up to now. Its main argument results from the fact that the sectors concerned are more than a hundred years old, some having been successively flooded and drained.

Table 1 :

 1 Main instability cases in Lorraine iron basin.

		Audun	1902 Drifting in progress	122	70	13,5	0,3	10
		Escherange	1919 Stoping adjacent	170	65	6,0	1,0	18
	BRUTAL	Sainte Marie 1932 Stoping adjacent Moutiers 1940 4 years after drifting Roncourt 1954 Stoping adjacent Roncourt 1959 Post-mining	153 121 147 140	65 70 70 75	5,0 11,0 7,5 5,0	2,0 0,9 0,9 2,4	16 10 15 17
		Rochonvillers 1973 Post-mining	190	62	4,5	1,3	15
		Rochonvillers 1974 Post-mining	190	61	4,5	1,7	15
		Jarny	1949 Drifting in progress	200	56	5,0	1,7	11
	PROGRESSIF	Auboué Crusnes Ville Montois 1982 Post-mining 1972 Extension in progress 150 1977 20 years after mining 180 166 Auboué 1996 20 years after mining 170 Auboué 1996 20 years after mining 150	45 50 55 53 45	6,0 3,8 4,5 5,0 6,0	1,7 2,9 2,0 1,3 2,0	12 11 9 11 11
		Moutiers	1997 20 ans après traçage	120	55	3,0	4,0	8
		Roncourt	1999 50 ans après traçage	140	53	2,5	2,4	9

Acknowledgements

The authors wish to thank the Ministry of the Economy, Finance and Industry and the Lorraine DRIRE (Lorraine iron basin) as well as CdF and HBCM (Provence coalfield) for their help in providing data and financing these studies.


Lionel Perrette
email: lionel.perrette@ineris.fr

Samira Chelhaoui
email: samira.chelhaoui@ineris.fr

David Corgier
email: ddcorgier@renaultvi.com

Safety evaluation of a PEMFC bus

In 2005, a fuel cell bus manufactured by the French company IRISBUS will operate in the streets of Paris. Under the French automotive regulation, IRISBUS is allowed to run its fuel cell prototype, provided that it guarantees its safe operation. This task requires a risk analysis to be carried out. INERIS has been chosen to drive this study. This paper reviews the risk assessment methodology and results. It highlights that maintenance and use are the most critical stages because of potential aggressions on the hydrogen system. It also indicates that the greater hazard potential lies in high pressure hydrogen storage and distribution. As such, it appears that more information is required on high pressure tank behaviour faced with different thermal and mechanical aggressions. Beyond, thermal fuse reliability has to be known. However, this bus features design principles and safety barriers that bring the risk down to an acceptable level.

' Régie Autonome des Transports Parisiens

Introduction

IRISBUS is born from the joint venture between RENAULT and IVECO BUS. It launched a project named CITYCELL which consists in running different fuel cell buses in 3 European cities: Paris, Turin and Madrid. The present paper concerns the French demonstration bus that will be operated by the Parisian transport company, RATP . So as to run the bus in the French capital and carry public passengers, IRISBUS shall get due authorisation from the French Ministry of Transport.

Since there is no specific regulation for fuel cell vehicles approval neither under the French automotive regulation nor at a European level, IRISBUS was asked by decision makers to demonstrate the safe operation of the bus for prototype approval.

INERIS has been tasked by IRISBUS to perform the bus preliminary risk analysis. This analysis targets the safe use of hydrogen. This paper will review methodology and the main conclusions in terms of risks and mitigation techniques for the different running stages.

l s Technical specifications

The proposed vehicle is a twelve meter articulated bus (see figure 1 below). It results from the combination of a trolley bus (external electrical power supply) and a fuel cell system. The power system is an hybridation between a 75 kW PEMFC and batteries. Batteries are a mean to recover braking and deceleration energy. They also supply additive power for peak demand.

Mf 4; I Ï Figure 1: Trolley bus -The Civis Cristalis

The bus is expected to have an autonomy of 200 km. Seven 150 litres pressurised composite tanks are used to store hydrogen. Tanks are made of an aluminium liner wired with carbon fibre impregnated by an epoxy resin. They have a service pressure of 350 bar. The system allows to store about 20 kg of hydrogen. All together, it weights about 580 kg. Each tank is equipped with :

-a manual valve for hydrogen purging (maintenance), -a thermal fuse to discharge hydrogen in case of fire, -a safety valve.

The stack will be located at the rear of the bus. Tanks as well as high pressure lines will be mounted on the roof. Along with the bus, a urban hydrogen filling station will be experimented.

Regulatory aspects

Currently, vehicle approval refers to European directive (Council Directive 70/156/EEC on the approximation of the laws of the Member States relating to the type-approval of motor vehicles and their trailers). This European directive does not cover vehicles running on hydrogen.

Meanwhile, under the same European directive, there are two ways for a manufacturer to run its prototype on the street:

-The vehicle manufacturer can refer to its own national authorities to gain a prototype approval. This prototype approval is valid as long as the number of vehicles is kept below 500 units per year. This national approval does not entitle to exportation unless the European exporting country agrees with this approval. In this case, the vehicle runs with a "W" type registration number. As long as the bus does not carry passengers, responsibility falls on the manufacturer. In other cases, public authorities are responsible.

-For larger production and for innovative technologies, any national authority belonging to the European Economic Community can submit a technical file to the European Commission. If the evaluation output is positive, the approved vehicle will be allowed to run freely within the 15 member states for a given period of time.

The first option is tailor-made for prototypes such as the present bus. Therefore, this option was chosen. On the basis of a risk analysis, Ministry of Transports is expected to deliver due authorisation.

Both options are interesting as long as prototypes or innovative vehicles are concerned. In the longer term, the current legal context will slow down hydrogen vehicles commercialisation.

The European project EIHP aims at proposing an harmonised procedure for hydrogen vehicles approval. EIHP proposals for liquid and pressurised hydrogen storage onboard vehicles have been forwarded to an ad-hoc GRPE group in Geneva. International regulations proposals are expected to evolve from this group.

Risk analysis : methodology

Proper risk analysis requires transversal skills and knowledge to be mixed. For example, this project working group gathered people from conception (stack, bus and storage), maintenance, as well as end-users along with safety experts. Unfortunately, none of the participants was representing rescue services.

The graph below (figure 2) shows the different steps involved in the risk analysis. A risk analysis requires targets to be identified. In a second time, the system has to be thoroughly described. Hazards inherent to the system as well as those coming from its environment shall be identified. The use of past experience (conventional and CNG buses) shall help in this task. Unwanted events that generate these hazards are referenced in a table similar to the one shown below. This table also mentions unwanted events likelihood and severity. Likelihood and severity criteria are set prior to carry the risk analysis (see tables 2 and 3). Their combination sets a criticity which is then compared with a risk acceptance level (see table 3).

Table 4: Criticity matrix

Safety measures drive the system back to an acceptable risk level (white area in table 4). Both number and type (procedures or technical measures) of barriers are related to the estimated criticity.

Safety measures implementation follows usual principles that consist in : 1. eliminating the hazard (conception choices), 2. reducing the hazard (conception choices and safety measures), 3. mitigating its effects (safety measures), 4. and finally removing or moving potential targets away from the hazard.

Safety barriers are expected to comply with specifications such as :

-fail safe conception, -ability to be tested and maintained, -ability to withstand expected aggressions, -ability to work in the case it suffers a default. Finally, their conception shall rely on a proven technology.

Results and discussion

All running conditions have been studied. That is to say :

-the hydrogen filling stage, -the bus maintenance and cleaning, -the use of the bus with passengers in a urban environment, -and finally the bus parking.

Tunnel crossing has not been looked at. The working group suggested that experience should be gained before considering tunnel crossing.

The filling station itself is also out of the frame of this study.

To ease the study, the bus has been split into its main functions:

-hydrogen storage and supply to the stack, -hydrogen conversion (air supply, electrochemical conversion, heat dissipation, outflow management, ...), -electrical architecture, -vehicle. According to this, noticeable unwanted events are listed in table 5 prior to be further discussed. The hydrogen filling stage Hydrogen leakage is quite a likely event as large quantities of high pressure hydrogen are transferred from the station to the bus. Besides, many situations can cause a leakage (improper plugging, open gates, tearing of dispenser lines, ...). Hazard associated to this situation is connected to the leaking flow, its duration, as well as to hydrogen potential to accumulate. Knowing the prone ability of hydrogen air mixture to be ignited, it is taken for granted that ignition will take place. Indeed, industrial experience shows that adequate hydrogen air mixture ignites most of the time.

I

Hydrogen release can be caused by a normal or abnormal opening of PRDs or safety thermal fuses. These releases are collected in order to be vented in a safe location.

Bus fire can be the consequence of ignited hydrogen leakage, electrical default, maintenance work, ... Fire is critical since it can eventually cause the burst of tanks. Burst of one tank can induce the burst of the other ones because of flying debris. Tank burst can also release sufficient energy to punch the passenger carrier compartment.

For safety reasons, filling takes place outside in a restricted access dedicated filling station. Ignition sources are controlled through the use of adequate electrical equipment, prohibition of work, switching of bus power sources, ... Hydrogen detection, pressure drop detection, accidental disconnection, ... cause filling to stop in emergency. Finally, hydrogen flow is limited and lines are fitted with check valves. Whenever, an explosion occurs the filling station is located away from other industrial equipment and dwellings.

Maintenance and cleaning

Maintenance is one of the most critical stage. Indeed, it can imply to act directly on the hydrogen distribution circuit in confined conditions. Confinement prevents hydrogen dispersion and dilution. It is therefore liable to significantly increase explosion hazards. Mechanical or natural ventilation prevents hydrogen from accumulating. However, it has no effect whatsoever on the dispersion plume, which is linked to the leaking diameter and pressure. Moreover, in a maintenance context, fire is a likely event (use of open flames to facilitate mechanical dismantling).

Leakage on the low pressure circuit is not that critical because of quantities of hydrogen and relative explosive volume that can be formed. This remark is valid, as long as the low pressure circuit is isolated from the high pressure one. Local confinement also has to be avoided.

Tanks rupture can be feared in case of inappropriate maintenance (drilling, open flames, ...). Drop of heavy loads carried above the tanks could also punch them. Finally, attention should be paid when maintenance concerns safety equipment. Their working conditions can be altered. To mitigate explosion hazards, it has been recommended whenever possible to limit tank pressure before to enter maintenance workshop. A list of critical maintenance work requiring tank to be emptied has to be drawn. A bay will be dedicated to hydrogen bus maintenance. This bay will have special features: control of ignition sources, hydrogen sensors, and ventilation eventually. High pressure lines are shielded against mechanical aggressions.

The use of the bus with passengers in a urban environment

In this situation, targets and hazards sources are numerous. Road collisions is one of them. Experience shows that the 4 000 Parisian buses suffer annually about 10 000 accidents. Most of these accidents concern the front right or left hand side of the bus. The low part of the chassis is usually concerned. In a urban situation, many other aggressions have been thought of: drop of objects from balconies, gun shot, falling trees, vandalism, ... CNG buses are to some extend comparable to the hydrogen ones. Experience with the 53 buses operated in Paris for the last 3 years indicates that natural gas sensors went of 20 times. Too low triggering level and exhausted gases were responsible for all of these detections. Once, a thermal fuse opened untimely because of friction between the fuse and the tank cover.

Whereas in the previous stages emergency means were more or less ready for action, in a urban context, it might take time for rescue services to intervene. Besides, it turns to be more difficult to set a safety perimeter in case of an accident. On the other hand, outdoor conditions enhanced hydrogen dispersion even though high buildings may refrain secure plume dilution.

Massive leakage can generate an explosive atmosphere as well as jet flames. Thermal fuses default leads to a leak that can not be stopped unless the tank is empty. Normal PEMFC purging does not cause any hazardous situation knowing the quantity of hydrogen involved.

Explosive atmosphere within the passengers carrier compartment has also been identified. This compartment is isolated from hydrogen equipment. High pressure lines are mounted on the roof to facilitate hydrogen dispersion. Finally, on a precautionary principle, a hydrogen sensor has been installed within the passenger compartment.

In case of a fire, the driver is expected to tackle the flames with a 6 kg powder extinguisher. He also invites passengers to evacuate. Evacuation takes no more than 5 minutes. Knowing the possibility for tanks to burst, a safety zone will have to be enforced. Time required for passengers and public to evacuate the zone has to be assessed. Possible burst shall not take place before this time.

Accidents can expose live parts. Circuit breaker and isolation means are used to prevent electrical shocks to passengers or rescue services. Moreover, high voltage parts will be tagged. Rescue services will be informed about specific explosive and electrical risks induced by this new technology.

Flow limiter and small diameter piping prevent large explosive atmosphere volumes. Length of high pressure lines is kept to a minimum. R110 rules will be followed. Leakage or catastrophic rupture would induce pressure drop. Drop in pressure detection causes feeding gates to shut off. Regarding the stack itself, electrical equipment are physically isolated from hydrogen equipment. All these equipment are ventilated. Finally, use of combustible material is controlled in order to prevent fire from rapid propagation.

The bus parking

On a parking stand, internally or externally triggered fire is possible. As said before, fire can ultimately be responsible for the burst of tanks. This risk has to be taken into account, while getting the fire under control. The hydrogen bus motion shall not be bounded to other buses position, in order to be able to move it freely if it is not on fire. Outdoor parking mitigates potential explosive atmosphere volumes and effects. Hydrogen gates shall be in a closed position and electrical parts shall be discharged as the bus is parked.

The table below summarises most of the selected safety barriers for all stages. Hydrogen storage and high pressure lines are the most hazardous part of the bus. Bursting and leaking potentiality has to be better understood and quantified for all types of aggressions. The following paragraphs discuss these events.

• Mechanical aggressions

The safety study has shown that tanks can be submitted to various mechanical aggressions: punch, shock, impact, crash, ... It is worth knowing tank behaviour to these aggressions. In this particular case, tanks are located on the roof, that is the less vulnerable part of the bus. Moreover, there are tightly linked to the chassis in accordance with RI 104 rule. Regarding collision head on (collision with tunnels), tanks are protected by the chassis. A top cover also protects them from sun light and minor mechanical aggressions.

Conclusion

This safety analysis browses all situations of bus operation. It highlighted that maintenance and use are the most critical stages because of potential aggressions on the hydrogen system. It also highlighted that the greater hazard potential lies in high pressure hydrogen storage and distribution.

As such, it appears that more information is required on high pressure tank behaviour faced with different thermal and mechanical aggressions. Beyond, thermal fuse reliability has to be known.

However, this bus features design principles and safety barriers. They bring the risk down to an acceptable level as long as high pressure tank safety equipment work as expected. This safety analysis is the preliminary link of a safety chain that includes calculations on potential explosive atmosphere volumes in case of line rupture or thermal fuse opening, overpressures related to tank burst and related safety distances,... It also includes more detailed AMDEC / HAZOP studies on the stack and the storage itself.

All together, it consists of the initial step of a safety program that includes:

-the check of critical equipment prior to be mounted on the bus, -the set up of a database to collect defaults and incidents in order to improve whenever required the overall safety level, the definition of specific maintenance procedures, -and finally the training of drivers, mechanics and rescue services. This demonstration project should bring sufficient experience on bus operability, safety and related public perception. Further use in tunnel for instance will then be considered.

Figure 2 :

 2 Figure 2: Logical steps in risk analysis

Function : hydrogen storage Date: Product or equipment : tanks N°Unwanted event Cause Consequences G Existing safety measures F Added safety measures G FTable 1 :

 1 Risk analysis table

Table 5 :

 5 List of unwanted events

Table 6 :

 6 List of proposed safety barriers

				:	"	,	i	l se ot lliu bus with
		,, ,	..,,.	,	liuS IllilllltlMlUIICC .111(1		.	,	'	"	, .
		llvcln>»en tilling stage	i	1 •• ' .			passengers in :i urban	:	Bus p:irkui[>
		-Outdoor filling station		-When possible limitation		environment -Implementation of Rl 10	-Gates ensure isolation of
		-Dedicated filling		of tank pressure			rules		tanks from high pressure
		dispensers			-Mechanical protection of		-Length limitation of high	circuit
		-Check valve on tanks		high pressure lines			pressure lines		-Outdoor parking
		-Limitation of equipment on		-Training			-High pressure lines are
		H 2 lines			-Tests of high pressure			located on the roof; they
		-Welding check		circuit after maintenance		are also protected
		-Safety filling pipe		-Dedicated bay for			-Leaking flow does not
		-Training			hydrogen bus maintenance		exceed 4 times stack H 2
		-Prohibition of any		-Control of ignition sources		consumption
	llyriruren storage il ml supply to lilt stuck	maintenance work during filling filling station -Restricted access to the		-Emptying of tanks for heavy or timely -Hydrogen venting is maintenance			-Pipe diameter is minimum -Pressure drop detection -Possibility to isolate tanks
		-Control of ignition sources		collected and discharged in
		-Filling is not possible if the		a safe location
		bus is running
		-H 2 detection, pressure
		drop, disconnection,
		induced emergency stop of
		filling
		-Pressure control
		-Flow limiter
		-Safety distances with other
		equipment
	Hydrogen				-Forced ventilation			-H 2 detection		-Ventilation of stack
	I'oir-vrsidii				-Detection					compartment before start-up
		-Internal bus power sources		-Tag and lock procedure			-Electrical isolation		-Discharge of electric
		are switched off (H 2 bus		-Discharge of electric			between passenger		capacities
	Klci'lrical	sensors are kept in		capacities			compartment and		-Internal bus power sources
	Hirliilui'tiire	function)						electrical system		are switched off
								-Floating mass
								-Circuit breaker
		-Fire fighting measures and		-Fume detection			-Tag of high voltage		-Freedom of movement
		equipment			-Fire fighting measures and		equipment		-Fire fighting measures and
		-Evacuation procedures		equipment			-H 2 detection in passenger	equipment
					-Critical safety equipment		compartment		-Evacuation procedures
					have been identified for a		-Isolation between H 2
	\ «hide				specific maintenance			circuit and passenger
					procedure			compartment
								-Tanks and high pressure
								lines on the roof
								-Use of non combustible
								materials

R 110 : Agreement concerning the adoption of uniform technical prescriptions for wheeled vehicles, equipment and parts which can be fitted and/or be used on wheeled vehicles and the conditions for reciprocal recognition of approvals granted on the basis of these prescriptions

High pressure hydrogen storage : how safe is it safe ?

Jet flames, explosive atmospheres or burst of tank can be induced namely by: -local (jet flames) or global thermal aggressions (bus on fire); -mechanical aggressions (road accidents, falling objects, flying debris, gun shot, .. .).

The graph below illustrates some of the causes leading to unwanted effects.

Accident on a pressurized H2 tank

Thermal aggressions

Thermal aggressions can ever be local (jet flame from an opened thermal fuse) or global (bus on fire). Whatever extensive the aggression is, safety measures shall be taken to prevent tanks from bursting. For global aggressions, the following questions have to be answered:

-What is the likelihood for a thermal fuse to remain closed ? -What is the influence of the fuse location in its ability to operate properly ? -Is there any critical thermal flux for which the tank can burst before it is depressurised ? -In the event of thermal fuse failure, how long will it take for the tank to burst?

Untimely fuse opening likelihood is also to be quantified. Combination of fuse reliability and sensibility shall indicate how many fuses are to be fitted to a pressurised tank.

Some tests 3 have shown that bursting can be induced by local aggressions. These tests indicate once more that number and location of fuses should be adequately set. It also underlines that tank structure must withstand the local aggression during the time required to get its pressure down. Tank bursting is not an acceptable event, unless either its likelihood is speculative or people have sufficient delay to get themselves in a safe location.

J . CHAINEAUX, « Security of highly pressurised tanks equipping GH 2 fuel road vehicles », Contract n°1 3461-97-11 F1ED ISP F, november 2000


Zbigniew Pokryszka
email: zbigniew.pokryszka@ineris.fr

David Grabowski
email: david.grabowski@ineris.fr

ON THE SURFACE NOXIOUS GAS EMISSIONS FROM OLD IRON MINES

Keywords: gas, emission, iron mine, pyrite. 1

A study carried out in the iron basin of Lorraine (North-East of France) highlighted the existence of under-oxygenated and noxious gas emissions (carbon dioxide, radon…) in vast built-up areas related to former mine workings.

Site investigations showed that the gas flow was mainly due to natural thermal air circulation between old mines and external atmosphere. The main hazards associated with this flow of gas and its potential accumulation in confined volumes in buildings are asphyxiation and intoxication. A further study is in progress in order to better understand the origin of the gas generation within the old iron mine workings. Following the first hypothesis, the carbon dioxide emission is mainly due to the specific hydro-geo-chemical process associating the pyrite oxidation with the acid dissolution of carbonates.

INTRODUCTION

Some residential areas of the iron basin of Lorraine (in the North-East of France) are affected by noxious gas emissions (under-oxygenated and noxious gas loaded mixtures). The most spectacular phenomenon occurred in the built-up area of Moyeuvre-Grande in Moselle (figure 1), especially in the district of the town located very close to the former underground mine workings in which some inhabitants observed faulty working gas cookers and boilers. In 1999, following the Lorraine local office of the French Ministry of Industry, Research and Environment (DRIRE) request, INERIS (the French National Institute for Safety and Risks) started some investigations to assess the geographical extent of the problem. First of all, the research was focused on the underground old mine workings area of Moyeuvre-Grande and also some neighbouring closed mines (Rosselange, Franchepré and Orne) whose main features are given in table 1. Volumes (Mm 3 ) 1,3 0,7 >30

Operating periods 1873 -1993 1890 -1935 1867 -1932 1871 -1901 18 ? ? -1978 Number In some cases, these emissions affected directly inhabited buildings in connection with the mine entrances or with the old superficial mine workings.

Moreover, a more exhaustive series of measures has been taken in other non-flooded mining reservoirs of the North of the ferriferous basin: Mexy, Lexy and Godgrange (figure 1 and table 1). It showed that similar emissions were occurring in these areas. However the intensity of the phenomenon was far less spectacular than those observed in Moyeuvre-Grande area (Pokryszka and Grabowski, 2002).

INERIS is now carrying out research to better understand the phenomenon of gaseous emissions from iron mines. These works are performed in the framework of a program funded by the French Ministry in charge of Industry in collaboration with the Lorraine local administration (DRIRE).

THE RISKS CREATED BY THE GASEOUS EMISSIONS

In an average atmosphere, the volumetric oxygen content is 20.9 % and the carbon dioxide one is 0.03 %. The consequences of a CO 2 enrichment and an O 2 impoverishment on health are numerous. The significance depends on the contents of these two gases: it ranges from headaches for CO 2 contents between 3 and 4 % and O 2 content between 14 and 16 %, to death for O 2 content lower than 6 % and CO 2 higher than 10 % [START_REF] Monomakhoff | Dangers résultant de la présence du grisou dans l'atmosphère de la mine. Rapport CERCHAR "Grisou -Grisoumétrie -Anémométrie[END_REF].

Fortunately, the composition of the air in inhabited buildings of the investigated area did not reach a level that could affect health. However in some cases, the measured contents were very close to the critical level for which harmful effects can appear by O 2 deficit and CO 2 excess.

Concerning the radon, it acts essentially on lung epithelium to further the development of cancer in case of longlasting exposure with content higher than 500 -1000 Bq/m 3 (Barrier and Lorentz, 2001).

Apart from these clinical effects, the observed dysfunctioning of combustion appliances (boiler, gas cooker) due to CO 2 polluted and under -oxygenated air are likely to increase the explosion and CO emission risks in concerned houses.

These observations induced INERIS to carry on its investigation, in order to better understand the mechanism of the gaseous emanations and to find reliable solutions to this problem.

GAS FLOW MECHANISM

In order to find the origin of the flow driving the gaseous exchanges between the old mine workings and the atmosphere, two measuring stations were installed at the entrance of two former galleries connected to the closed mine in Moyeuvre-Grande area.

These stations continuously monitored the intensity and the direction of the air flow, the temperature and the content of some components of the emitted gas (O 2 , CO 2 and radon) as well as the outer temperature and the atmospheric pressure. These results, in agreement with the numerous measurements done during the preliminary reconnaissance phase between 1999 and 2000 (Pokryszka and Grabowski, 2002), clearly show a link between the direction and the air flow rate, and the outer temperature. No other parameter having a significant effect on the flow (like the atmospheric pressure) has been identified. This fact led us to consider the natural thermal convection as the phenomenon creating the gaseous flow between the mining reservoir and the surface. This phenomenon is due to the thermal differential existing between the outer atmosphere and the old workings.

The temperature in the latter is almost constant (in that case, it was between 12 and 14 °C), unlike the atmospheric temperature whose level varies with the seasons.

In the same way, the surface topography above the mining reservoir of Moyeuvre-Grande and other concerned reservoirs seems to favour the natural thermal convection.

Indeed, the investigated areas are globally characterised by a very contrasted topography with rather large plateau distinctly intersected by an irregular network of valleys. The ferriferous deposit is located at the base of the plateau. Outcrops can be frequently seen on the hillsides.

During summer, a part of the atmospheric air in contact with the upper old workings, by unsealed openings or cracks, is subject to a progressive cooling. An increase in the bulk density due to the cooling causes the gas migration to the lower parts of the old emerged mine workings. Then the gas can leave the reservoir by the various openings and apertures mainly situated along the deposit outcrop on the hillsides. An air circulation becomes established in this way the principle of which is presented on figure 4. The flow direction is shown for the summer period.

Te > Ti

According to this model, the reservoir working can be divided in two distinct stages: the summer stage during which the flow is globally coming out of the massif. The outer temperature is significantly higher than the inside one (between 12 °C and 14 °C); the winter stage during which the flow is globally coming into the old mine workings. The flow rate is clearly setting when the outer temperature is lower than the old mine workings temperature.

These two stages are separated by transitional periods during which the flow is fluctuating.

Figure 5 represents the O 2 and CO 2 rates measured in an opening of the Moyeuvre-Grande reservoir as a function of temperature. It clearly shows that the O 2 impoverished and CO 2 enriched atmosphere appears for temperatures higher than 12 °C to 14 °C, that corresponds to the coming out flow stage. Furthermore, a simulation based on the simplified model of Budryk (Budryk, 1932) showed that the aeraulic charge due to this phenomenon was estimated in the region of 20 to 30 Pa for a thermal differential of 10 °C. Such a charge seems to be large enough to create the gas circulation in this context.

It is also important to underline that other natural events (such as the atmospheric pressure variations, the wind effect, the aquifer fluttering …) add to the natural thermal draught phenomenon and can disturb temporarily the two main stages characterised previously, without being enough to reverse them on a long term basis.

ATMOSPHERE MODIFICATION MECHANISM IN THE MINING RESERVOIRS

4.1.

Atmosphere modification in the studied reservoirs

The old mine workings gas composition, monitored continuously in the Moyeuvre-Grande reservoir and punctually in other reservoirs of the North of the ferriferous basin, have shown that the atmosphere changes within the old mine workings were consisting in a O 2 consumption accompanied by a less than proportional production of CO 2 .

This result, in accordance with a literature review on geochemical reactions observed in similar mining development cases [START_REF] Feuga | Contribution à l'estimation des termes entrée sortie des gaz transitant dans d'anciens travaux miniers[END_REF], led us to consider the pyrite oxidation (found in the ferriferous deposit) coupled with the dissolving of calcite by the produced sulphuric acid as the most convincing hypothesis to explain the CO 2 emission and the O 2 consumption.

The reaction balance can be written:

FeS 2 + 3,75 O 2 + 2 CaCO 3 + 1,5 H 2 O  Fe(OH) 3 + 2 CO 2 + 2 SO 4 2-+ 2 Ca 2+
Moreover, an other study revealed the presence of reactive pyrite in the rich in calcite marly interstratifications of the Lorraine ferriferous field [START_REF] Collon | Chemical changes in groundwater due to flooding of an iron mine in a non-acid-producing environment[END_REF].

Furthermore, another way to check the assumption about the pointed out geochemical mechanism occurred with the opportunity of the Tressange pilot site experimentation, which consists in a controlled submergence of a small part of the old mine workings.

Atmosphere modification measured in the massif during the submergence

In order to estimate the flooding effect on the abandoned mining structures and on the water and atmosphere quality, a specific experiment has been carried out within the framework of the GISOS research program (GISOS, 2003). It is realised in real and in situ conditions, in the old mine of Tressange.

For the needs of that experimentation, a part of a former mine stripped by the room and pillar method used in two geological levels has been isolated, fitted out and prepared to be flooded in a controlled way.

Only the deeper layer and the base of the marly interstratifications have been flooded. A simplified diagram of the experimental site is presented on figure 6.

INERIS took advantage of this opportunity to check the effect of the ferriferous massif flooding on the possible gas production and on the atmosphere composition within the rock.

An experiment has been set to follow the atmosphere evolution (O 2 and CO 2 contents) in the isolated enclosure, during the flooding. The atmosphere has been monitored thanks to two boreholes drilled through the marly interstratifications and emerging into the mine workings of the brown layer to be flooded. Therefore, three main stages can be distinguished:

Before water reaches the marly interstratifications level (35 first days), the atmosphere composition measured at the base of the boreholes does not evolve; As soon as the water level reaches the marly interstratifications base and progresses inside this layer (end of April 2002), the air contained in the borehole is becoming poorer in O 2 and richer in CO 2 ;

Once the water level is almost stabilised in the marly interstratifications, the atmosphere composition does not evolve any more.

Such results clearly show an obvious link between the atmosphere composition modification (O 2 consumption and CO 2 production) and the marly interstratifications flooding on the experimental site.

The atmosphere modification would be resulting from a chemical reaction between the pyritic marl and the flooding water, which tends to confirm our hypothesis according to which the CO 2 emission would be due to a geochemical reaction combining pyrite from marl with calcite dissolved by produced acids.

Hydrogeochemical analysis being carried out in laboratory are in progress in the framework of the GISOS scientific program to provide additional information about the mineralogical massif composition and so to confirm this in-situ obtained result.

CONCLUSIONS

The non or partly flooded mine workings constitute a CO 2 laden polluted air source. The flows are generated by natural thermal draught and the initial atmosphere composition modification is very likely due to a chemical reactions set involving marly interstratifications, pyrite and calcite, in contact with water and air.

This hypothesis seems to be confirmed by the investigations carried out in real mine conditions (Tressange experimental site) where the evolution of the atmosphere in contact with the marly interstratifications has been observed.

Figure 1 .

 1 Figure 1. Localisation of the studied mining reservoirs

Figures 2

 2 Figures 2 and 3 shows some interesting results obtained from the measures made from April 2000 to December 2002.

Figure 2 .

 2 Figure 2. Outside temperature, air flow rate and radon content evolutions observed in a gallery connected to the Moyeuvre-Grande reservoir.

Figure 4 .

 4 Figure 4. Schematic diagram of the natural thermal draught during the summer period (Moyeuvre-Grande reservoir case)

Figure 5 .

 5 Figure 5. Correlation between the outside temperature and the O 2 and CO 2 contents observed in an opening of the Moyeuvre-Grande reservoir

Figure 6 .

 6 Figure 6. Schematic section of the atmosphere watch over system during the flooding of the Tressange experimental site

Figure 7 .

 7 Figure 7. Evolution of the massif gaseous composition during the flooding of the Tressange pilot site.

Table 1 .

 1 Features of the studied reservoirs (old mine workings)

		Moyeuvre
	Name of the reservoir	Orne	Rosselange	Lexy	Mexy	Godbrange
		Franchepré
	Area (ha)	355,7	252


O Bour

E Taffoureau

R Therrien

LATERAL LANDFILL GAS MIGRATION: CHARACTERIZATION AND PRELIMINARY MODELING RESULTS

Lateral landfill gas migration occurs in the surroundings of a MSW landfill with capped old waste cells. Three areas were identified by poor vegetation growth. Methane flux chamber measures confirm the extent of these areas, as observed in the cultivated land. These measures and some complementary physical measures were used to build a conceptual model of lateral landfill gas migration in the geological layer. Prior to drilling new boreholes on the site, a preliminary simplified model will be built with the numerical code TOUGH2-LGM. A description of the geological units, methane flux and the results of the preliminary modeling are presented. The aim of this work is to identify the key parameters used to predict the travel distance of landfill gas.

INTRODUCTION

The estimation of lateral migration of landfill gas (LFG) is necessary to predict fires and explosion hazards and damages for the environment. Lateral LFG migration occurs preferentially in the case of old cells without sealing. In France, capping of old cells is common because of current French landfill regulations. Capping stops the upward migration of LFG and requires a LFG recovery network. However, drainage of LFG can be more difficult to use in the case of old cells, where the degree of saturation of the waste and mixing of the waste with fine grained soils could reduce the permeability and the capacity of a collection network. In this case, lateral LFG migration could occur, especially in the case of semi-permeable to permeable soil with buried waste.

Investigations of methane flux and composition of the gas in borehole are usually conducted to estimate LFG travel distance. Modeling of methane production and migration can help predict the methane plume migration in the soil [START_REF] Poulsen | Modeling lateral gas transport in soil adjacent to old landfill[END_REF]. That type of modeling requires detailed characterization of soil properties for the entire vadose zone concerned with the plume migration. For this project, it was chosen to perform preliminary modeling of the LFG migration and focus field investigation on the upper geological layer to reduce both costs and time required.

CHARACTERIZATION OF THE MIGRATION

Site description

The landfill site was located in Malle ville-sur-le Bec (Eure region, France). The annual capacity of the site is approximately 70 000 tons. Exploitation began in 1971 in the south area, where waste was buried without building a compacted clay liner at the bottom of the cells. Exploitation has been ongoing since 1994 within cell 1 (but with a bottom seal) as well as within the north area of the site. After capping the old cells with a geosynthetic liner, areas of poor crop production (zone Zl, Figure 1) have appeared in the vicinity of the site. The landfill was then equipped with a LFG collection network but because drainage in the old cells was poor, they were not connected to the collection network in the first years after the deposit of the waste. The study of a landfill site began in 2001 with the investigations of two different diminishing crop production areas, zone Zl and zone Z2 (Figure 1). The site was initially exploited for clay and later filled with domestic waste. The geology at the site consists of, downward from the land surface :

• A superficial cover of silty clay loam ;

• A residual flint clay layer ;

• A chalk aquifer.

The thickness and the permeability of the first two layers vary within the site. The log of previous boreholes reveals that the thickness of the silty clay loam generally decreases from north to south. The thickness of the residual flint clay is approximately 12-14 meters, but the variation in thickness for the site is less known than that of the upper silty clay loam. Large variations of the thickness and permeability of the two top geological layers are likely to occur in the site. Several decimetric layers of flint blocs almost without clay could also be observed in the residual flint clay layer. The permeability of the silty clay loam and the residual flint clay range respectively from 1. 10" 6 m/s to 3. 10" 8 m/s and from 3. 10" 7 m/s to 1. 10" 9 m/s in the north part of the site. A rising trend for the silty fraction from east to west was also observed. Additional investigation of the top layer with manual auger in the center of the zone 1 confirms these trends.

The rain water reaches the lower chalk layer. The water table is located 75 meters below the ground level in the chalk aquifer. The contrast of permeability between the chalk aquifer and the more clayey level of the residual flint clay could probably permit a temporary thin accumulation of groundwater. This hypothesis will be used later when we focus our attention on the first two geological layers in the vadose zone (Figure 2).

Flux measures

The flux chamber method [START_REF] Savanne | Comparison of different methods for measuring landfill methane emissions[END_REF] was used in October 2001 and August 2002 to investigate methane flux coming from the surface in the area of the crop damage. At each measure location an area of 0.25 m 2 was caped by a chamber with recirculating gas. A flame ionisation detector enables to observe the enrichment of methane in the chamber. In the north area (zone Zl), measures were located on a 5 meters spacing grid which covers the area of the crop damage and the surface between the major outlines. The methane flux measures range from 0 to 82 ml/min/m 2 . In zone Z2, Z3 transects with a 10 meters spacing between measures were t* investigated. The methane flux measures range from 0 to 90 ml/min/m but there were only 6 measure points with positive values and two hot spots of methane flux. The maximum distance between a hot spot of methane flux and the landfill limit reaches 80 meters. These methods demonstrate that, during the first campaign in October, there was good agreement between the area of the crop damage and the occurrence of the methane flux, especially in the zone 1 (wheat culture).

The first crop damages are located in the zone Zl and are contemporaneous with the capping of the old cells with liner. The source of LFG is also probably coming from one of these old cells. Under this assumption, the LFG should migrate perpendicularly and under new cell n°l. LFG reaches the distance of 130 meter for the farthest hot spot of methane. Travel distance of several hundred meters have been reported in the literature (Nastev & al., 1998). For the other areas (zones Z2 and Z3), sources are more difficult to locate. Nevertheless it is more than likely that preferential migration is associated with the flint blocks layers located in the upper part of the residual flint clay and the more silty fraction of the silty clay in the west area. A documented report of an explosion that occurred when drilling piezometer Pz3 in this material confirms this assumption.

Properties of the vadose zone material

The first meter of the silty clay loam was investigated in order to obtain physical parameters and a secondary parameter correlated with the methane flux. Water content of undisturbed soil samples was investigated in three profiles. Capillary pressure was also measured by tensiometers. These parameters permit to adjust Van Genuchten parameters for the silty clay loam. A geophysical campaign was also conducted in zones Z2 and Z3 to detect the influence of more resistive material in the superficial layer. A geophysical survey done with a EM 38 device was not conclusive in zone Z2, most probably because of recent ploughing of the land. In zone Z3, the contrast was better and some correlation could be done between the geophysical and flux measures. The only hot spot of methane flux was located in the transitional stratum between the silty clay loam and a more silty layer with some flints (Figure 3).

PRELIMINARY MODELING OF LFG LATERAL MIGRATION

Model and boundary conditions

The aim of this preliminary modeling was to simulate LFG production and migration in the soil under the new cell and in the area of the field (zone Zl). Limited data is available to characterize the buried refuses in the old cells and the flint strata. Due to the restricted data set used, the major aim was to better understand the migration along the transverse section. The simplified assumptions used do not permit to compare the distribution of the methane flux along the transverse section and the flux methane measures. Nevertheless, knowledge of the date of the deposit and the description of the waste permits the use of a first order kinetic model for the rate of methane production, with a half life equal to 7 years. The lack of data concerning the location of the flint strata permits only to use a simplified homogeneous description of the geological material.

For this study, steady state water saturation is assumed. Water infiltration in the field was modeled with a medium infiltration rate of 100 mm/year, which is approximately half of the real annual recharge. Actually, recharge at the site occurs only during half of the year and the net infiltration rate is null during the month of the flux measures. Also the choice of a medium rate permits to maximize the gas saturation of the soil void and then the relative gas permeability.

The numerical model TOUGH2-LGM used was developed by Nastev & al. (1998) to simulate LFG production and migration. The model contains a new equation of state to account for 4 fluid components (air, water, methane et carbon dioxide) and heat. An average temperature of 10 °C was chosen for the entire profile. The mesh of the transverse section used 18 grid blocks in the vertical direction and 28 grid blocks in the horizontal direction. A saturated bottom boundary with no charge of water was used in order to limit the migration in the first 12 m modeled (Figure 3). The waste located above the ground level were supposed not to contribute to the migration. A zero flux limit was also chosen for the left lateral boundary concerning the waste and silty clay loam left column.

CONCLUSIONS

The flux characterization and preliminary modeling presented in this study allow a better understanding of lateral landfill gas migration. In the field, methane flux concentrated on hot spots. We need additional evidence to demonstrate the leading part of the transitional strata between the silty clay loam and the residual flint clay layer. Further studies including calibrations with the methane measures in boreholes and better description of the geological strata are necessary to improve the efficiency of the modeling predictions of the travel distance of methane.

Figure 1 .

 1 Figure 1. Scheme of the location of the crop damage and transversal section studied.

Figure 2 .

 2 Figure 2. Geological and conceptual scheme of the lateral LFG migration in the transversal section A-B.

Figure 3 .

 3 Figure 3. Scheme of the location of the hot spot of methane flux in the Z3 zone.

Table 1 -Figure 3 .Figure 4 .

 134 Figure 3. Scheme of the boundary conditions used for the modeling


Thoraval Alain

Vincent Renaud

ENSMN Ineris

HYDRO-MECHANICAL UPSCALING OF A FRACTURED ROCKMASS USING A 3D NUMERICAL APPROACH

A new upscaling method has been developed using 3D numerical tools (RESOBLOK & 3DEC). This method has been successfully compared with standard analytical approaches in the case of a simple fracture network. This method has been applied to determine the equivalent permeability, stiffness and Biot tensor of a real fracture rock-mass at different scales. The effects of the fracture network properties and of the state of stress on the result have been investigated.

CONTEXT OF THE WORK AND OBJECTIVES

This work is part of the INERIS contribution to work package 3 (WP3) of the European project BENCHPAR. The purpose of WP3 is the better understanding the impact of upscaling on the THM processes on Performance Assessment of nuclear waste disposal problems.

The INERIS contribution to WP3 consist in:

The development and evaluation of a new upscaling method, and its application to WP3 fracture data;

Comparison between H, HM, THM simulations considering an equivalent homogeneous medium with FLAC3D (whose properties will have been determined previously by the upscaling process). This paper presents only the upscaling work.

DESCRIPTION OF THE UPSCALING APPROACH & VERIFICATIONS

A new upscaling method has been proposed by INERIS to determine the equivalent hydro mechanical properties of a fractured rock-mass. This method is based on the 3D numerical simulations of the behaviour of a "sample" of fractured rock-mass submitted to different hydromechanical boundary conditions. The simulations are defined in order to determine the equivalent permeability, stiffness and Biot tensor of a fractured rock-mass.

Two numerical tools are used: RESOBLOK [START_REF] Heliot | Generating a Blocky Rock Mass[END_REF] to generate the 3D fracture network; 3 DEC [Itasca Consulting Group, 1994;[START_REF] References Damjanac | A three-dimensional numerical model of water /low in a fractured rock mass[END_REF] to make the hydro-mechanical computations.

Generation of the fracture network

The code used to generate the fracture network (RESOBLOK) is based on the assumption that the fracture can be considered as polygon. This code is able to make determinist or stochastic simulations. To make the stochastic simulations done for the WP3 exercise (figure 1), the following assumptions are done:

Fracture density (number of fracture per unit volume of rock -P 3 i) is assumed to follow a Poisson law;

Fracture orientation is assumed to follow a Fisher law. Average dip & dip direction as well as the Fisher coefficient (k f ) has to be given for the different fracture sets;

Fracture length is assumed to follow a power law distribution of the form N= CL~D, where N is the number of fractures which are equal to or greater than a given trace length L, D is the fractal dimension and C is a constant.

Determination of the equivalent permeability tensor Kij

The hydraulic gradient J and the fluid flow Q are linked by the relationship:

Qi --Kij Jj (1)
To determine all the tensor terms with 3DEC, 3 sets of (hydraulic) boundary conditions have to be defined (figure 2). After each computation, the "equivalent" vectorial flowrate is averaged from the flowrate value at different locations of the fracture network. The different terms of the equivalent permeability tensor are then computed from equation (1).

The results of this new upscaling approach have been compared with an analytical solution given for the simple geometry (figure 3).

L Figure 3. Fracture network used for comparison

with analytical upscaling approach.

The analytical equivalent hydraulic conductivity tensor can be expressed in this case as following: The equivalent permeability given by 3DEC is:

K,j3DEC = (6pga q 3 DEc/AP) 5% (3) = 2.35 1Q-7 m/s
where q 3DEC = 4 10" 2 m/s is the discharge (m/s) computed with 3DEC for AP = 10 6 Pa.

We can see that the 3DEC result is very close to the analytical solution.

Determination of the equivalent Stiffness tensor Tijkl

The stress tensor CTJJ, the elastic strain tensor e^, the Biot's tensor By and the Biot coefficient G are linked by the Biot's equation:

P = -G(B M eu -$
where P is the fluid pressure and production.

(4)

(5) the fluid To determine all the tensor terms with 3DEC, 6 sets of (mechanical) boundary conditions have to be defined (Figure 4 and5). During each 3DEC runs, the average strain tensor is computed from displacement differences between opposite face of the model (100 pairs of point is used for each strain term).

A verification has been done again for the same simple geometry (figure 3). From Oda theory [1986], it can be shown that, in this case, (T^)" 1 is equal to: The determination of this matrix with the numerical approach gives the following result: Tijkl 3DEC -3.205 1.000 0.220 0.220-0.014-0.014-0.005] 0.220 1.000 0.220 -0.014 -0.005 -0.014; 0.220 0.220 1.000 -0.005 -0.014 -0.014 -0.001 -0.002 0.000 0.780 0.000 0.000 0.001 0.000 0.000 0.000 0.780 0.000 0.001 0.003 0.003 0.000 0.000 0.780 It can be noticed that, in all cases, the differences between the analytical and the 3DEC solution are very small (less than 1.5%).

V "Ë 0 0 0 v ~~~Ê v E _L " E + LK n 0 0 0 0 0 0 n F" + 7JT o o 0 0 0 0 0 0 0 + v n 1 -- E LK, 0 0 l + v n

Determination of the equivalent Blot tensor Bij

The numerical procedure is to impose a fluid pressure built up from P to P + AP within the model, assuming no mechanical deformation at the boundaries. In this case the Biot's equations (4 & 5) can be rewritten as follows:

= -BiAP & AP-Q [START_REF] References Damjanac | A three-dimensional numerical model of water /low in a fractured rock mass[END_REF] with AC = VpVo/Vt, where VQ the initial fluid volume (under P), V] is the final fluid volume (under P+AP) and V t the total model volume.

For the geometry of figure 3, it can be shown:

ô and G=-LK»(l-2v)+nE

With 3DEC, we determine the average stress variation Aay and volumetric water content variation AC induced by the hydraulic pressure AP.

We have found again in this case that the difference between analytical and numerical solution are below 2%.

APPLICATION OF THE UPSCALING TECHNIQUE TO WP3

The objective of this section is to compute, with the numerical approach already presented and validated, the equivalent permeability tensor of the fracture network of the WP3 case. The set of computation presented here concerned the fracture network of formation 1 (computation has also been done for formation 2 and the fault zone).

The size of the zone of interest where the fracture network is simulated has been progressively increased from 2 m to the greater possible size according to the numerical problem encountered. Five simulations have been done for each model size to get an idea of the standard deviation.

3.11nput of the fracture network geometry into 3DEC

RESOBLOK simulations

The fracture network is generated by RESOBLOK from the data gathered in Table 1. The simulations have been done considering initial value of P 3 i equal to the value of PU (for a Fracture Length Threshold of 0.5 m) given in Table 1. Average values of ? 2 2, computed from several RESOBLOK simulations, were compared to ?22 measurements (5m/m 2 for a fracture length threshold of 0.5 m) to evaluate the realism of the simulations. The results are gathered in Table 2.

We can notice that the computed P 2 2 value at 3 m or 4 m are close to the measured value (5m/m2).

Interfacing RESOBLOK to 3DEC

A specific command of RESOBLOK is used to generate a 3DEC file including all information regarding fracture network geometry.

3DEC does not allow generating fractures that stop inside a block. During the 3DEC cutting process, the plane where the RESOBLOK polygonal fracture is lying is continued up to the next fracture (or block face). However, 3DEC automatically distinguish 2 zones (from information output from RESOBLOK): an "active" zone inside the polygon area (real joint), an "inactive" zone outside the polygon (fictitious joint). So 3DEC is able to "give" different properties to "real" joint and fictitious joint.

Checking the effect of fictitious joints on the equivalent permeabity tensor

We have shown, from various 3DEC runs, that the prolongation of joint up to the next block face (that artificially increased the fracture network connectivity) could lead to overestimate the permeability by a factor of 2 or 3.

The hydraulic aperture used for the real joint is supposed to be 6.5 10° m. We have considered a hydraulic aperture of 10" 6 m for the fictitious joint. We have checked that the equivalent hydraulic permeability tensor computed considering a fictitious joint network is 5 orders of magnitude smaller than if we consider a real joint network.

Checking the effect of fictitious joints on the equivalent stiffness tensor

The joint stiffness used for the real joint is supposed to be 4.34 10" 11 Pa/m. We have considered a joint stiffness of 10~1 2 Pa/m for the fictitious joint. We have checked that, for this value, the equivalent stiffness tensor is very close to the intact rock stiffness tensor.

We have evaluated the effect of considering the fictitious joints as real joints on the equivalent stiffness tensor Tjjki. We can notice that the terms of the equivalent stiffness tensor terms are smaller if we consider 3DEC fictitious joints are real. For the WP3 case (where high joint stiffness values are considered) the differences remain small. It increased up to 30 % for Kn = 4.34 10~1 0 Pa/m.

Limitation of 3DEC code

We have noticed that some problems occur during the cutting process into 3DEC (version 3.00.073), when the fracture number becomes important (and maybe when the angle between fracture become small). I happens that in some cases, the cutting is rejected or an error appears later in the process of block geometry recognition. Some efforts have been done (in collaboration with ITASCA) to overcome those limitations but, up to now, no satisfying answer has been found

The consequence of this limitation is that the determination of Ky and TJJM was not possible (for WP3 data) for a model size higher than 2x2x2 m 3 if we want to avoid artificial prolongation of joint into 3DEC. The limit is push back to 5x5x5 m 3 for Tjjki (and unchanged for Ky) a artificial joint prolongation is considered.

Determination of the WP3 "equivalent" permeability tensors (Kij)

Assuming constant hydraulic aperture

We have assumed that the hydraulic apertures were constant and equal to 6,5 10" 5 m. The equivalent permeability tensor is given below at 2 m scale for the fracture network of formation 1 and considering a fracture length threshold of 0.5 m: Because of 3DEC limitations, the upcaling was done only at 2 m scale; so the existence of a REV (and its value) could not be established at the present time.

Fracture network simulations were done considering the relation (established from measurements by Nirex [START_REF] Andersson | Understanding the impact of upscaling THM processes on performance assessment[END_REF] between the number of fractures per meter and the fracture length threshold. With a view to the fracture network simplification, we have tried to determine the effect of the fracture length threshold (used to make the RESOBLOK simulations) on the upscaling process. Unfortunately, we have shown that, at 2 m scale, it was not possible to conclude about the effect of the fracture length threshold on KJJ. The reason is that the connection of the fracture network to the model boundaries, where the pressure gradient is applied, is artificially increased when the model size is small in comparison with the fracture length threshold.

Assuming stress dependent hydraulic aperture

For HM computation, 3DEC consider a relation between the hydraulic aperture "a" and the mechanical aperture "u" that can be written: a = ao + Au, where ao is the zero stress aperture. A maximum and residual aperture is considered for numerical stability reason. We have assumed that a.nax = a 0 = 6.5 10" 5 m and that a res = 1.8 10" 5 m. Figure 6 represent the evolution of the diagonal terms of the equivalent permeability tensor with stress applied on the model boundaries (at 2 m scale). We can see a decreasing of the permeability when the applied stress increases. This decreasing is much faster if the normal joint stiffness is assumed to be smaller (4.43 10 10 Pa/m instead of 4.43 10" Pa/m). It can be notice that there is a residual equivalent permeability (around 10" 9 m/s) that has to be related to the residual hydraulic aperture. That residual permeability is reached for a > 20 MPa (» 800 m) if the joint normal stiffness In this work we have assumed no relation between the joint stiffness and the stress or with the fracture length. This assumption could be revisited later on. We have also considered an isotropic state of stress. Others stress conditions could be applied and would probably involve other kinds of evolution of the equivalent permeability tensor.

Determination of the WP3 "equivalent" stiffness tensor (T ijk i)

Tjjki has been first determined considering a prolongation of the joint up to the next joint (all the cutting plane generated by 3DEC are assumed to be active). The results for formation 1 are gathered in Table 3 (we have distinguish the average values from the standard deviation of the different terms). From this result, we can see that the REV is higher than 5 m. Model sizes become too important (~ 250 Mo) to allow us to check if stabilization is about to be reached.

Tjj k i has also been determined at 2 m scale considering the real length of the joint. Differences with the previous simulation remain below 10 %. This could lead us to think that this factor is not very important for the equivalent mechanical behaviour. This result is consistent with the fact that no significant fracture length threshold effect has been shown on Tyki.

The results is highly related to the choice of the fracture stiffness value. This invites us to investigate later on the effect of stress on the result considering stress dependent stiffness.

CONCLUSIONS

An upscaling approach based on the 3D numerical simulation has been set up. Very good correlations with analytical solutions have been found for a network made of 3 sets of parallel fractures.

This approach has then been applied to fracture data given in the WP3 specifications in order to get the equivalent hydromechanical properties of the fractured rockmass.

The equivalent permeability tensor KJJ has been computed at 2 m scale considering the real length of the joint (we have shown the importance to avoid artificial joint prolongations). The REV could not be estimated due to code limitation reasons. The relation between permeability and stress (considering an isotropic state of stress) has been determined and is highly dependent on the joint stiffness value.

The equivalent stiffness tensor Tjj k i has been computed up to 5 m scale. No clear conclusion can be given for the mechanical REV. We have notice that the fracture network connectivity does not seem to have a strong influence on T uk i.

The new method presented in this paper is challenging because it is based on 3D explicit simulations of the hydromechanical behaviour of a fractured rock-mass. Some improvements are still to be done to make this method really operational. However those 3D results could contribute (through comparisons) to the evaluation of the more operational 2D methods used in BENCHPAR WP3. Accordingly this contribution could be profitable to the upscaling world.

ACKNOWLEDGEMENT

We wish to thank Ki-Bok Min from Royal Institute of Technology (KTH), Sweden for a helpful collaboration about mechanical upscaling using UDEC and 3DEC, and Johan Andersson from JA Streamflow AB, Sweden for many fruitful advices.

We acknowledge with gratitude the financial support of the European Community and of the MEDD and MEFI french ministries.

Figure 1 .

 1 Figure 1. Stochastic fracture network generated by RESOBLOK.

Figure 2 .

 2 Figure 2. Boundary conditions used to determine the permeability tensor with 3DEC.

  Kij analytical = (p 9 ^ / 2 // L) S i} (2) = 2.45 m 7 m/s where: a (= 10~4 m) is the hydraulic aperture of the fracture; u, (= 10~3 Pa s) is the dynamic viscosity of the fluid; (L = 20 m) is the edge length of the model; p = 1000 kg/m 3 ; g =9.81 m/s 2 .

  of stress (a = Ie6 Pa) is applied into the model; a constant stress increment (Acr = Ie5 Pa) is apply on the grey faces

Figure 4 .

 4 Figure 4. First set of boundary conditions used to determine Tijkl with 3DEC.

  Assuming E (Young's modulus)= 5.294 GPa; v (Poisson ratio ) =0.324; Kn & Ks (normal &shear joint stiffness) =1 GPa/m; n (number of fracture in each fracture set) = 3; L = 20 m, we get (in GPa)

Figure 6

 6 Figure 6 Equivalent permeability tensor variation with stress.

  is Kn = 4.43 10 11 Pa/m and for CT > 2 MPa (« 80 m) Kn = 4.43 10 10 Pa/m.

Figure 7 Figure 7 .Figure 8 .

 778 Figure 7 and figure 8 gives the evolution of the average value and standard deviation of the diagonal terms of Tjj k] terms from 2 m to 5 m scale.

Table 1 :

 1 Joint set properties for formation 1

Table 2 :

 2 Determination of fracture density from RESOBLOK fracture network simulations.

	«Sample » size	Set	Computed P22 density (m/m 2 )
		1	0.11
	3 X 3><3 m3	2 3	1.37 1.87
		4	0.97
		Total	4.39
		1	0.26
	4x4><4 m 3	2 3	1.27 2.57
		4	0.97
		Total	5.01

Table 3 :

 3 Equivalent stiffness tensor for formation 1 at 2 m scale

	a -average values