Mouhamad Hassoun
email: mouhamad.hassoun@ineris.fr

Pascal Villard
email: pascal.villard@3sr-grenoble.fr

Marwan Al Heib
email: marwan.alheib@ineris.fr

Fabrice Emeriault
email: fabrice.emeriault@3sr-grenoble.fr

Soil reinforcement with geosynthetic for localized subsidence problems: Experimental and analytical analysis

Keywords: Sinkhole, geosynthetic, design, trapdoor, granular and cohesive soils

This study focuses on the evaluation of the load transfer mechanism and the determination of the shape of the load distribution transmitted to the geosynthetic layer when a cavity appears under granular and cohesive backfills. Trapdoor experimental technique with high speed acquisition of digital images and continuous monitoring of load and displacements were used for this purpose.

Testing different soil types, it has been demonstrated that an approximate parabolic or inverted triangular load distribution seems acceptable for granular soil layer, whereas for cohesive backfill the load distribution could be approximately represented by two concentrated forces near the edges of the cavity. In both cases, an important overload on the soil surface could change the shape of load acting on the geosynthetic sheet. Experimental results are then approached by Terzaghi s formulation with an appropriate shape of load distribution and a convenient value of the earth pressure coefficient. Finally, recommendations are proposed to promote a better design of such 1 HASSOUN et al., March 29, 2018 reinforcement structures.

INTRODUCTION

Localized collapse (sinkhole) can occur in karstic regions, former mining areas or in zones of poor geotechnical characteristics. Ground subsidence represents a significant geo-hazard not only to infrastructures but to residential properties as well [START_REF] Heib | Large-scale soilstructure physical model (1g)-assessment of structure damages[END_REF]. The geosynthetic reinforcement (GR) is an economic and effective solution to reduce the total and differential settlement problems when limited construction time is available and small deformation is allowable [START_REF] Dggt | Recommendations for Design and Analysis of Earth Structures Using Geosynthetic Reinforcements-EBGEO[END_REF]. When a geosynthetic layer is placed on a deformable soil on which vertical loads are further applied, loads are transmitted to the geosynthetic layer relieving the subsoil unable to support them. In this case, the geosynthetic takes up the vertical loads and diffuses them in the form of tensile forces transfered by friction in the anchorage areas.

The design of such geosynthetic reinforcement layer goes through two main steps: Step (1) calculates the effective amount of the applied load on the GR layer, where a load transfer mechanism should be taken into account. Step (2) determines the maximum GR strain on the basis of the result of step (1); load distribution over GR, subsoil reaction, and GR stiffness should be known. Finally knowing the maximum GR strain and the GR stiffness gives the mobilized tensile force which should be smaller than the long-term GR tensile strength.

Several methods are available to design a GR over cavities (RAFAEL, 1997;[START_REF] Briançon | Design of geosynthetic-reinforced platforms spanning localized sinkholes[END_REF]BS8006, 2010;[START_REF] Dggt | Recommendations for Design and Analysis of Earth Structures Using Geosynthetic Reinforcements-EBGEO[END_REF], all with their own hypotheses for the design calculation.

The difficulty is that these models give results that largely differ, and may lead to more expensive design than necessary [START_REF] Huckert | Load transfer mechanisms in geotextile-reinforced embankments overlying voids: Experimental and analytical approaches[END_REF].

The objectives of the experimental study are to determine with precision the load transfer mechanism that develop within shallow granular and cohesive backfills subjected to localized subsidence and the corresponding load distribution on the GR layer and further to validate an appropriate analytical design model with these results. (1965) as the ability of a material to transfer loads from one location to another in response to a relative displacement between the locations. The pioneer work on this subject was performed by [START_REF] Marston | The theory of loads on pipes in ditches and tests of cement and clay drain tile and sewer pipe[END_REF] and [START_REF] Terzaghi | Theorical soil mechanics[END_REF] who reproduced this effect with a trapdoor test by imposing a localized displacement to a horizontal rigid support. The classical experimental setup to investigate load transfer in the trapdoor test has been reviewed by many authors, (Vardoulakis et al., 1981;[START_REF] Evans | An examination of arching in granular soils[END_REF]Stone and Muir Wood, 1992;[START_REF] Costa | Failure mechanisms in sand over a deep active trapdoor[END_REF][START_REF] Chevalier | Etudes expérimentale et numérique des transferts de charge dans les matériaux granulaires. application au renforcement de sols par inclusions rigides[END_REF][START_REF] Iglesia | Investigation of soil arching with centrifuge tests[END_REF]Cox, 2014) under normal gravity as well as under an increased gravitational field using a centrifuge facility. All researchers observe that the movement of an active trapdoor causes an instantaneous reduction of soil stresses (initially geostatic) above the trapdoor and an increase of stresses in the adjacent soil mass, but they seldom agree on the value of this reduction. [START_REF] Terzaghi | Theorical soil mechanics[END_REF] assumed that the lateral load transfer occurs through shear stresses along vertical planes going from the edges of the trapdoor to the ground surface. [START_REF] Iglesia | Determination of vertical loading on underground structures based on an arching evolution concept[END_REF] investigated the load transfer effect using a trapdoor apparatus in a geotechnical centrifuge and postulated the loading profile on the trapdoor as well as the arching evolution.

While many researchers have chosen one shape to represent the zone affected by the subsoil collapse [START_REF] Guido | Plate loading tests on geogrid-reinforced earth slabs[END_REF][START_REF] Carlsson | Reinforced soil, principles for calculation[END_REF][START_REF] Hewlett | Analysis of piled embankments[END_REF][START_REF] Han | Numerical analysis of geosynthetic-reinforced and pile-supported earth platforms over soft soil[END_REF], centrifuge tests performed by [START_REF] Iglesia | Determination of vertical loading on underground structures based on an arching evolution concept[END_REF] suggested that the volume of soil supposed to act on the area of subsoil of poor mechanical characteristic goes through a series of stages between curved, triangular and prismatic shape before coming to rest or collapse. [START_REF] Stone | Effects of dilatancy and particle size observed in model tests on sand[END_REF] and [START_REF] Chevalier | Etudes expérimentale et numérique des transferts de charge dans les matériaux granulaires. application au renforcement de sols par inclusions rigides[END_REF] have reported under shallow conditions (H/B ≤ 2) a mechanism similar to that given by [START_REF] Iglesia | Determination of vertical loading on underground structures based on an arching evolution concept[END_REF]. The mechanism reported for shallow backfills observed by [START_REF] Stone | Effects of dilatancy and particle size observed in model tests on sand[END_REF] and [START_REF] Chevalier | Etudes expérimentale et numérique des transferts de charge dans les matériaux granulaires. application au renforcement de sols par inclusions rigides[END_REF] involves the development of multiple (internal and external) failure surfaces in the region above the trapdoor, with soil directly above the trapdoor remaining essentially rigid. This failure mechanism differs significantly from those observed by Costa et al.

defined internal failure surface which becomes gradually more inclined to the vertical when the downward movement of the trapdoor increases. This sequence of arching evolution explains the variation in stress acting on the trapdoor.

There are several other recent studies focused on the numerical modeling of the load transfer effect (Le [START_REF] Hello | Embankments reinforced by piles and geosyntheticsnumerical and experimental studies dealing with the transfer of load on the soil embankment[END_REF][START_REF] Chevalier | Etudes expérimentale et numérique des transferts de charge dans les matériaux granulaires. application au renforcement de sols par inclusions rigides[END_REF][START_REF] Han | Dem analysis of stresses and deformations of geogridreinforced embankments over piles[END_REF][START_REF] Pardo | Experimental and numerical study of arching soil effect in coarse sand[END_REF][START_REF] Huang | Experimental and numerical study of geosynthetic reinforced soil over a channel[END_REF]Villard et al., 2016;[START_REF] Rui | Evolution of soil arching; 2d dem simulations[END_REF]. [START_REF] Rui | Evolution of soil arching; 2d dem simulations[END_REF] observed numerically three soil evolution patterns in multiple trapdoor systems. The focus of previous studies on active trapdoor has been mainly limited to the failure mechanism, arching evolution within soil mass and the interpretation of the loading curve on the trapdoor, where no geosynthetic reinforcement has been used or fully studied. Based on true scale experiments [START_REF] Briançon | Design of geosynthetic-reinforced platforms spanning localized sinkholes[END_REF], numerical studies of the behavior of a reinforced thin granular layer (ratio between the thickness of the embankment and the width of the cavity H/B = 0.25) were performed [START_REF] Villard | Coupling between finite and discrete element methods for the modelling of earth structures reinforced by geosynthetic[END_REF][START_REF] Yu | Influence of selection of soil and interface properties on numerical results of two soil-geosynthetic interaction problems[END_REF]. Comparisons were mainly based on the load transfer mechanisms, the surface settlements and the tensile strains of the geosynthetic. The DEM-FEM model (Villard et al., 2009) provides additional results that cannot be obtained during the experiment, as for example the strain pattern within the granular layer under large deformations. The existing conclusions concern the granular material case (cohesionless frictional backfills), and cannot be easily used for a cohesive soil layer. [START_REF] Huckert | Load transfer mechanisms in geotextile-reinforced embankments overlying voids: Experimental and analytical approaches[END_REF] of the proposed tests at 1g proved difficult to ascertain [START_REF] Zhu | Model tests on interaction between soil and geosynthetics subjected to localized subsidence in landfills[END_REF]. Consequently, it will be difficult in this context to make quantitative interpretation of the experimental results. Scaling factors of the main parameters of the model test are resumed in Table 1. The use of a flexible rubber membrane for the reinforcement at the base of the soil layer does not allow measuring the total load acting on the trapdoor. Therefore, an original specific process using photogrammetry technique has been used to determine the reinforcement strain and deduce the vertical load distribution acting on the membrane all over the void.

Soil and reinforcement properties

The considered soil layers are either made of a granular (cohesionless) material or a cohesive sandy-clay. The shear strength parameters of these geomaterials have been obtained from conventional triaxial compression tests ( rubber membrane has been used to represent the geosynthetic reinforcement in order to respect the condition of similarity and to allow measuring the tensile strain using a photogrammetry technique.

The tensile behaviour of the membrane has been characterized in tensile tests on 20 ± 1mm wide specimens. Load -extension curves show a non linear elastic behavior of the tested membrane, the variation of the axial stiffness J defined per unit length of the geosynthetic layer in the range of strain measured in trapdoor tests is given in Fig. 2. Considering the scaling factor of n = 10 a value of the tensile stiffness equal to 20 kN/m in the small model corresponds to a value of 2000 kN/m in the full-scale case which is a classical value used in reinforcement applications.

Test program

The model tests using gravel have been carried out for three different overburden heights, i.e., H/B = 0.25, 0.5 and 0.75 respectively. Tests 1-3 have been carried out without GR, however in Tests 4-6 the GR is used to conclude on the effect of the GR on the load transfer mechanism and backfill soil behavior. Only one GR layer with a unique tensile stiffness is tested here. A soil layer with H/B = 0.25 has been tested for the sandy-clay layerin tests 7-8. This study mainly deals with the behavior of shallow backfill over cavity where (H/B ≤ 0.75). The test program is listed in Table 3.

We note that each presented test has been repeated three times, 24 tests have been carried out in total, all results are similar and will therefore be given in the format of the mean experimental value ± the maximum average spread.

6 HASSOUN et al., March 29, 2018 In Tests 4-6 and 8, the soil surface is loaded successively by two linear uniform loads (Q 1 and Q 2 ) placed above the middle of the trapdoor using steel bars of 25 ± 1mm in width, 50 ± 1mm in height and 400 mm in length. Each steel bar Q 1 and Q 2 has a total weight of 23 N. These loads have been applied to investigate the reinforcement response against such surface loading, and purposely to produce the collapse of the cohesive backfill in Tests 7-8.

Photogrammetry

A key requirement of the research is to provide accurate measurements of soil movement and GR displacements in ordre to determine accurately the vertical load distribution acting on the GR.

Digital images have been captured remotely via a PC mounted next to the model test using 15 Megapixels cameras. Images have been processed using the image analysis software VIC 2D. Full field two-dimensional displacement and strain data are thus provided for both the soil layer and the GR sheet. Determining the strain distribution at different points on the GR sheet over cavity is a big issue since it is further used to calculate the tension force induced in the GR and also the shape and intensity of the load transmitted to the GR due to the cavity appearance as explained in the next section. The soil layer is considered as having naturally a large enough variation in texture to allow an accurate tracking from one image to another. In addition, a speckle pattern has been projected on the edge of the GR using spray cans to enable the use of digital image correlation technique (DIC) on this thin structure. The GR is sufficiently separated from the box boundaries to avoid any interaction that could disturb the results.

Note that in order to effectively filter out scatter in experimental measurements, displacement curves have been smoothed. Smoothness of both vertical and horizontal measurements is essential in order to obtain clearer estimations for strain and load distribution. Smoothing was applied using an interpolation spline, which fits a smooth curve exactly through the given data points.

The used spline is defined piecewise by polynomials, using this smoothing theory avoids Runge s phenomenon and provides the best smoothing curve [START_REF] Birkhoff | Piecewise polynomial interpolation and approximation[END_REF]). The 1standard deviation confidence in the match at a correlated point changes from 0.0125 to 0.115

pixel (1pixel = 0.105948mm

GR response: Exploitation method

The GR strain is calculated successively between every two points i and i + 1 of GR layer all along the trapdoor by comparing their initial coordinates noted (x i , y i ) and (x i+1 , y i+1 ) and their corresponding final positions after the trapdoor has moved down noted (x

′ i , y ′ i ) and (x ′ i+1 , y ′ i+1 )
respectively (Fig. 3). The strain between points i and i + 1 is obtained, considering no change in curvature between the two points. The strain ε i,i+1 between points i and i + 1 is obtained by:

ε i,i+1 = ε j = ((x ′ i+1 -x ′ i ) 2 + (y ′ i+1 -y ′ i ) 2 ) ((x i+1 -x i ) 2 + (y i+1 -y i ) 2 ) -1 (1)
The induced tensile force defined per unit width of the geosynthetic is then deduced by:

T j = J (ε)ε j (2)
Knowing final positions of points on the GR, the GR inclination with respect to the horizontal direction could be determined at any point along the GR layer as following:

tan α j = (y ′ i+1 -y ′ i )/(x ′ i+1 -x ′ i ) (3)
Using the induced tensile force and the GR sheet inclination in Eq. 2 and Eq. 3, the vertical component of the tensile force can be calculated with:

T v, j = T j sin α j (4)
Finally, the vertical load acting at any point of the GR layer over the cavity is obtained considering the vertical equilibrium of a part of the GR of length △x (Fig. 3):

q j, j+1 = q k = △T v △x = T v, j+1 -T v, j x j+1 -x j (5) 8 HASSOUN et al., March 29, 2018
Where q k expressed in k N/m 2 is the vertical surface load between two points j and j+1 on the GR over the cavity, whose abscissa difference is noted △x, and △T v is the associated variation in vertical component of the tensile force defined by meter width. It should be noted that shear forces along GR over the cavity originating from soil / geotextile interaction are not taken into account in the derivation of Eq.5.

Experimental validation of the procedure for the load distribution determination

In this section, we report on experiments designed to test the effectiveness of the exploitation procedure presented above to deduce the distribution of load transmitted on the GR over cavity.

The presented experimental results aim to secure that the exploitation method is satisfying with respect to representativeness, reproducibility and repeatability. Validation experiments consist in the loading of the GR layer with a cylindrical metal tube with a weight of 22N and a diameter equal to 50mm spreading over the model box width (400mm) with a gap of 2mm. Three tests (Tests (a), (b) and (c)) have been carried out in the same conditions and compared in terms of strain distributions. No soil mass is used in these tests, the GR layer is stabilized with a 300N metallic plate on each anchorage side next to the trapdoor edges. The plates have the same dimensions as the anchorage zones and their weight is supposed uniformly distributed over the whole anchorage area (400mm x 400mm) (Fig. 4).

Strain distributions of GR for three identical tests are presented in Fig. 5a. Experimental values show a good convergence with almost a linear strain distribution over the trapdoor. These results are close to the theoretical estimation (ε = 1.25%), obtained from the analytical resolution of the GR sheet equilibrium under cylindrical tube loading assuming no friction between the flexible rubber membrane and the tube, a constant value of the GR tensile strain over the void and an horizontal displacement of the GR in the anchorage area (0.28 mm as experimentally observed).

Strain distribution is slightly antisymmetric over the cavity, this could be explained by the fact that the anchorage conditions were not perfectily identical on both sides of the anchorage zones, and that the GR layer was not fully mobilized. Fig. 5b shows the vertical tension curve, obtained from the average smoothed strain distribution in Fig. 5a applying Eq.4. An approximately constant value HASSOUN et al., March 29, 2018 of T v, j is observed between the two edges of the cavity.

If we note A and B the points of the GR at the edge of the cavity, we obtain T v,A = 30.3N/m and T v,B = 24.7N/m ( Fig. 5b). The weight of the cylindrical tube could be estimated using:

Q 2T v = 0.4(T v,A + T v,B ) = 22N
which is equal to the real weight of the cylinder.

Using Eq.5, the shape of load distribution transmitted to the GR can be determined as shown in Fig. 5c. Transmitted load reaches a maximum at the tube/GR interface, then drops suddenly to nearly zero along the GR over the rest of the trapdoor. The weight of the cylinder could be also calculated by the integration of the load distribution curve over cavity saying :

Q Int = x=+100 x=-100 z=400 z=0
q k dx dz = 22.56N. The integration of the load distribution curve over cavity seems also satisfying to estimate the real applied load on the GR. Q 2T v and Q Int are relatively equivalent.

The exploitation procedure is therefore judged reliable and fulfilling the intended purposes. All the results presented thereafter are based on the average strain between repeated tests.

UNREINFORCED GRANULAR SOIL: TESTS 1-3

Loading response -Arching theories vs. experimental results

Fig. 6 shows the trapdoor average vertical stress p normalized by the geostatic stress p 0 = γH, against the trapdoor displacement δ normalized by the trapdoor width B.

Similar trends are obtained in Tests 1-3 with a good general agreement with [START_REF] Chevalier | Etudes expérimentale et numérique des transferts de charge dans les matériaux granulaires. application au renforcement de sols par inclusions rigides[END_REF] and [START_REF] Iglesia | Investigation of soil arching with centrifuge tests[END_REF] findings. As the trapdoor is lowered, the stress on the trapdoor drops rather abruptly to reach a minimum value generating maximum arching and then gradually increases (approximately linearly) before being roughly constant until the trapdoor cannot move down any further. The most significant difference between Tests 1-3 is the stress reduction amount and the ultimate recovered stress. The displacement corresponding to the minimum load on the trapdoor tends to fall within a wider range (δ/B ≈ 1.35%) for all tests. This range is in full agreement with results reported by [START_REF] Chevalier | Etudes expérimentale et numérique des transferts de charge dans les matériaux granulaires. application au renforcement de sols par inclusions rigides[END_REF]. Initial stress is reduced to 68 %, 55% and 35% of p 0 for respectively H/B = 0. with soil depth. This could be explained by the fact that the load transfer action is not identical and prevented from effectively forming in Tests 1-3, i.e., arching cannot be fully sustained in regular manner when the overburden is relatively shallow and the ratio of soil height to trapdoor width (H/B ≤ 0.75) is lower than the lower limit (1.5 ≤ H/B ≤ 2) estimated by [START_REF] Terzaghi | Theorical soil mechanics[END_REF] to have a stable arching action.

In addition, for the ultimate state reached for large trapdoor displacement δ (i.e. δ/B = 100%),

Test-1 with H/B = 0.25 shows that the initial geostatic pressure is approximately fully recovered, whereas in Tests 2 and 3 with H/B = 0.5 and 0.75 a reduction of 13% and 20% respectively is still observed.

In this section, more attention is payed to the estimation of the maximum stress reduction generated by load transfer mechanisms. A summary of the equations proposed in previously published arching theories to estimate minimum and ultimate loading caused by active trapdoor movements is provided in Table 4.

Among the minimum load expressions proposed in the literature, Engesser (1882), [START_REF] Iglesia | Investigation of soil arching with centrifuge tests[END_REF], [START_REF] Bierbaumer | Die Dimensionierung des Tunnelmauerwerkes: Studien[END_REF], [START_REF] Evans | An examination of arching in granular soils[END_REF] and [START_REF] Guido | Plate loading tests on geogrid-reinforced earth slabs[END_REF] immediately into contact with each other and allow the formation of a force (arch) path oriented toward stable areas. The moving area is located above the cavity in the shape of an inverted parabola (Fig. 7a). At this point, the maximum load transfer mechanism is obtained. [START_REF] Chevalier | Etudes expérimentale et numérique des transferts de charge dans les matériaux granulaires. application au renforcement de sols par inclusions rigides[END_REF] reported that this stage corresponds to soil decompaction.

Loading recovery: this stage is a transition from the maximum load transfer (minimum loading)

to the ultimate state. The moving zone of soil is larger and reaches the soil surface, the initial parabolic displacement pattern cannot be sustained and evolves towards an approximately trapezoidal one (Fig. 7b). Thus, Fig. 6 shows that stress variation in this stage is almost linear, the slope represents the way the load transfer decreases progressively as trapdoor move down. Fig. 6 shows that load recovery index decreases with the ratio of the height of the soil layer to the trapdoor width.

This can be explained by the fact that arching is more efficient with thicker backfills.

Ultimate state: as the trapdoor continues to move. the whole height of the backfill above the trapdoor is affected by the displacement of the latter, the arch pattern doesn't exist anymore and two vertical shear bands develop at the edge of the trapdoor instead (Fig. 7d).

The soil immediately above the trapdoor is moving at the same rate as the trapdoor displacement δ. On the opposite, the ratio of the settlement at the soil surface to the trapdoor displacement δ is significantly lower than 1, indicating the zone of soil above the trapdoor is dilating.

REINFORCED GRANULAR SOIL: TESTS 4-6

For these tests, due to the presence of the GR, neither subsidence displacement δ nor the applied stresses on trapdoor have been measured.

Vertical displacement analysis

Vertical settlements are significantly smaller than those obtained in the unreinforced soil case not only at soil surface but also at different depths in the backfill. For Test 6, the maximum vertical displacement at soil surface appear in Fig. 8 as reduced 35% compared to the unreinforced backfill in Test 3 for the same final trapdoor displacement (9.5% of B) at which the GR takes off from the trapdoor. This observation is explained by the mobilization of the GR layer that prevents the soil above the edges of the cavity to move downwards and thus reduces its deformation and expansion.

GR behavior: Strain and load distribution

Tensile strain and load distributions along the GR layer determined by photogrammetry analysis are presented in this section. The results of Tests 5 and 6 showed similar trends, therefore tensile strain and load distributions are given for only Tests 4 and 6. For Tests 5 and 6 (under backfill weight and overloads), the tensile strain of the GR at the edges of the trapdoor are generally larger than those obtained at midspan, while the tensile strain variation of the GR in the central part of the trapdoor is quite small even after surface loading (Fig. 9). The reason of the higher strains in GR at the dges of the trapdoor than in the adjacent zone is due to the vertical stress concentration at the edge of the trapdoor shown hereafter,which is similar to the abrupt-increase strain in geosynthetic at the edges of piles firstly discovered by [START_REF] Han | Numerical analysis of geosynthetic-reinforced and pile-supported earth platforms over soft soil[END_REF]. Obtained strain patterns, above the cavity and in the anchorage zones, are similar to the analytical, numerical and full-scale experiment results founded by [START_REF] Briançon | Design of geosynthetic-reinforced platforms spanning localized sinkholes[END_REF], [START_REF] Villard | Coupling between finite and discrete element methods for the modelling of earth structures reinforced by geosynthetic[END_REF]. On the opposite, the GR tensile strain in Test 4 is approximately constant over the length of the trapdoor for unloaded case (Fig. 10). This phenomenon is explained by the large amount of sliding of the GR layer in the anchorage areas where the confinement stress is very low and unable to effectively counterbalance the load transmitted to the GR on top of the trapdoor. The tensile strain in the GR reaches a maximum in the center of the trapdoor and decreases gradually towards the trapdoor edges, the load transfer mechanisms are thus strongly affected. 2016) and [START_REF] Han | Numerical analysis of geosynthetic-reinforced and pile-supported earth platforms over soft soil[END_REF].

Further loading of the soil surface with Q 1 and Q 2 , the shape of load distribution in Test 5 does not significantly change, a small increase in the stress in the center and at the trapdoor edges is observed. The backfill in Test 6 appears to be less influenced by surface loading. On the other hand, for unreinforced Test 4 the load distribution does not reflect an effective load transfer mechanism.

Loading the soil surface drastically changes the shape of the load distribution acting on the GR, with a concentration of stress near the center of the trapdoor (Fig. 12).

The stress reduction ratio (SRR) is defined as the ratio of total load Q transmitted to the GR above the trapdoor and calculated experimentaly with Q = 0.4(T v,A + T v,B ), to the initial average geostatic stress applied by the backfill weight (E) plus eventual surface loads Q 1 and Q 2 . The lower the value of SRR, the greater the arching effect, a SRR of 1.0 implies no arching.

SRR = Q/(E + Q 1 + Q 2 ) (6)
As expected, Fig. 13 shows that SRR decreases with an increasing backfill height, a larger soil arching effect being underlined for thicker soil layers. Stress reduction ratio is approximately 70%, 45% and 35% in unloaded Tests 4, 5 and 6 respectively. Arching in Test 5 is shown to be maintained after Q 1 was applied. Arching effect is even more relevant in Test 6, where a constant SRR is calculated for all loading cases.

For Test 4, the large displacement of the granular layer due to the sliding of the GR in the anchorage areas, combined with the frictional mechanism at the interface between GR and the granular particles, lead to complex load transfer mechanisms that increase with additional overloads. A remarkable point, that needs to be pointed out, is that SRR values for Tests 4-6 (with reinforcement)

14 HASSOUN et al., March 29, 2018 are approximately equal to the maximum stress reduction ratio in unreinforced Tests 1-3 (Fig. 6).

This observation leads to conclude that the reinforcement layer has maintained the "maximum load transfer" stage mentioned earlier even for larger subsidence.

Comparison with analytical approaches of the GR deflection and strain

Several analytical models for the design of GR layer are available in literature. An analytical model requires two elements: an arching model that provides the stress applied on the GR layer and the shape of the load distribution on the GR above void. From these elements, the equilibrium equations can be established and the corresponding GR tensile strain and induced tension are calculated. These models generally differ in arching values and load distribution shapes. Considering the equilibrium of a differential rectangular element on trapdoor, [START_REF] Terzaghi | Theorical soil mechanics[END_REF] estimates the ultimate vertical applied stress σ v for unloaded cohesionless backfill as follow:

σ v = q 0 = γB 2K tan φ (1 -exp 2K tan φ H B ) (7)
Where K is the ratio between the horizontal and vertical stresses (K = σ h σ v ).

Terzaghi formula provides a stress estimation that depends to a large extent on the value of K.

Various authors have used this theory but with different assigned values for K, which is difficult to be determined experimentally with an accurate value. The different proposed values of lateral earth pressure coefficient K are shown in Table 5.

Fig. 14 shows a comparison of average soil pressures acting on the GR calculated by Terzaghi formula for different K expressions given in Table 5 with Tests 4-6 results. The sensitivity of Terzaghi equation (Eq.7) on different formulations of K is clearly illustrated, high disparity between different estimations of vertical stress is observed. The calculated soil pressures using K proposed by [START_REF] Aubertin | Interaction between backfill and rock mass in narrow stopes[END_REF][START_REF] Chen | Modified terzaghi loozening earth pressure based on theory of main stress axes rotation[END_REF] underestimate tests results. On the opposite, [START_REF] Handy | The arch in soil arching[END_REF] and [START_REF] Marston | The theory of loads on pipes in ditches and tests of cement and clay drain tile and sewer pipe[END_REF] highly over-estimate experimental measurements. [START_REF] Terzaghi | Theorical soil mechanics[END_REF] proposition appears as a secure choice. [START_REF] Huckert | Experimental and numerical approaches of the design of geosynthetic reinforcements overlying voids[END_REF] The shape of the load distribution on GR layer is also of high importance. Here various geometries for the distribution of vertical stress have been tested in terms of maximum GR deflection and tensile strain. All the considered distributions have the same average stress intensity, as defined by

Terzaghi in (Eq.7), with K = 1.3 as proposed by [START_REF] Huckert | Experimental and numerical approaches of the design of geosynthetic reinforcements overlying voids[END_REF]. φ is assumed to be the peak friction angle of the granular material. The use of φ peak is based on the results obtained in the previous section where a maximum arching effect was proved at the end of Tests 4-6. Only the cases without loading at the soil surface are considered in this section. Analytical developments proposed by [START_REF] Huckert | Load transfer mechanisms in geotextile-reinforced embankments overlying voids: Experimental and analytical approaches[END_REF] and Villard et al. (2016) have been used for comparison with the experimental results. The main assumptions of the proposed analytical model are a non uniform load on the GR over the cavity, the sliding and the friction of the GR in the anchorage areas on each side of the cavity.

Parabolic and inverted triangular load distributions are proposed to match the experimental results. Equilibrium equations for the corresponding load distribution are given in Table 6. The effect of both null and low (denoted NZ for Non-Zero) values of the stress at the center of the trapdoor has also been studied. For the NZ cases, the vertical stress at the trapdoor center is assumed to be equal to 0.22kPa as shown in Fig. 12 and Fig. 11.

Importance of GR displacement at the edge of the cavity (U A ) on its maximum deflection and tensile strain is also highlighted by using experimental values of U A (U A = 1.34, 0.82, 0.77mm) for Tests 4-6 respectively. Comparison between analytical and experimental results for Tests 4-6 in terms of maximum GR deflection and strain is presented in Table 7.

The comparison shows that NZ parabolic or inverted triangular load distributions are required to estimate accurately the maximum deflection and tensile strain in the GR. This conclusion is valid for granular cohesionless backfills and considering an adequate value of K = 1.3 as proposed by [START_REF] Huckert | Experimental and numerical approaches of the design of geosynthetic reinforcements overlying voids[END_REF] and the displacements of the GR layer at the edges of the cavity as recommended by [START_REF] Briançon | Design of geosynthetic-reinforced platforms spanning localized sinkholes[END_REF] in such case. Few experimental tests have been carried out so far to simulate the sinkhole opening under a cohesive soil layer. The mobilized failure mechanisms in such cohesive soil under localized sinkhole being still not clearly identified, the main objectives of this section are a better understanding and description of the total transferred load, of the shape of the load distribution on the GR layer and of the GR behavior in anchorage zones.

In Test 7, an unreinforced layer of manually compacted cohesive soil with H/B = 0.25 has been observed bridging the trapdoor cavity with very small maximum deflection (0.35%B) after the trapdoor was moved down. The cohesive backfill behaves as a slab with small deflection at its center. No cracks appeared neither at the edges nor at the center of the cavity. In this case, the soil layer ensures its own stability thanks only to its cohesion (Fig. 15a). Similarly to reinforced granular Tests 4-6, when the trapdoor reaches its lower position, loads Q 1 and Q 2 have been applied

on the ground surface to cause the backfill failure. Q 1 appears as the minimum load required for the collapse of the unreinforced soil layer associated with the development of several relatively big blocks on the trapdoor as observed in Fig. 15b.

The previous collapse mechanism has not been fully observed when the cohesive layer is reinforced (Test 8). Firstly, the cohesive soil layer lost instantaneously the contact with the trapdoor and the GR layer deflects due to its own specific weight, the slab behavior is the same as in the unreinforced case. Loading the soil layer, a large volume of cohesive soil falls principally in one block on the GR. Approximately the same overload (Q 1 ) has been required to cause the failure as for the unreinforced backfill, no influence of the GR is shown at this stage. The main fallen block of cohesive soil has a characteristic geometry (trapezoidal one), limited by major cracks propagating from the edges of the trapdoor up to the area close to the loaded surface (Fig. 16a). Increasing the surface load to Q 2 , cracks within the soil layer continue to develop with greater thickness and the GR layer is more curved at the trapdoor centre under the block of collapsed soil (Fig. 16b).

The speckle pattern projected on the GR layer allows to determine the profiles of horizontal and vertical displacement after the rupture of the cohesive soil layer and to deduce therefrom the tensile strain distribution in the GR layer using (Eq.1). 17 HASSOUN et al., March 29, 2018 It can be seen that tensile strain distribution in GR shows a constant value above the trapdoor under Q 1 and a slight variation under Q 2 (Fig. 17).

Applying the same procedure as in the granular case, after calculation of different tension components with the corresponding GR inclination, the vertical stress q k acting at all points of the GR layer is obtained using the relation with vertical component of the tension T v in (Eq.5). The stress distributions obtained after the collapse of the cohesive soil layer due to loading Q 1 and Q 2 are presented in (Fig. 18).

For the two loading phases, a high concentration of stress is observed near the edges of the cavity due to the block collapse mechanism. On the other hand, the stress is very low in the central part for loading with Q 1 but more important under Q 2 . The small vertical stress value on the central part of the GR layer especially under Q 1 could explain the flat shape of the GR layer in this part (Fig. 16a).

Integrating the vertical stress distributions in Fig. 18, the total load transmitted to the GR for Q 1 and Q 2 loading cases is respectively 34 N and 60 N, i.e. 54% and 70% of the total collapsed weight above the trapdoor (weight of the collapsed soil block + overload). These values clearly lower than 100% could be explained by the collapse mechanism, where the cohesif soil layer breaks in blocks and that the forces are transmitted from one block to another and to the GR layer only at the contact points. The soil layer does not collapse completely on the GR and the soil blocks touch the edges of the trapdoor.

The obtained collapse mechanism is very close to that proposed by [START_REF] Huckert | Experimental and numerical approaches of the design of geosynthetic reinforcements overlying voids[END_REF]. [START_REF] Huckert | Experimental and numerical approaches of the design of geosynthetic reinforcements overlying voids[END_REF] proposed to represent its effects by a localized action (stress) applied by the collapsed block on the GR layer. The distance between the two localized force is equal to the length of the base of the collapsed soil volume.

The analytical approach proposed by [START_REF] Huckert | Approches expérimentale et numérique du dimensionnement de renforcements géosynthétiques sur cavités et inclusions rigides[END_REF] (Fig. 19). The collapsed block is further considered rigid, undeformable and dimensionally stable.

The proposed analytical model consists in solving Eq.8 for the horizontal tension T h .

U A + [ 1 + ( P T h ) 2 -1](l -B/2) = T h J (ε) [(1 + ( P T h ) 2 )(l -B/2) + l] (8)
Huckert assumed that the length 2l between the forces representing the effect of the collapsed block could be estimated with: 2l = (l1 + 2H tan φ) where l1 is the upper base of the collapsed block that could be approximated by the overload application length (equal to 50 mm for Q 2 loading case). The horizontal displacement U A is here replaced by its experimental value (0.51 and 1.55mm

for Q 1 and Q 2 respectively).

The mechanical behavior of the GR layer, at the right of the cavity, is then given by the expressions of maximum deflection, tension and GR deformation as follow:

f = P T h (B/2 -l) (9)
T max = T A = T B = T h 1 + ( P T h ) 2 (10) ε A = T A /J (ε) ε B = T B /J (ε) (11)
The comparison of the analytical results obtained with Huckert's model and the experimental results for the two loading phases Q 1 and Q 2 in Test-8 is given in Table 8, the geometrical properties of the collapsed soil block being deduced from experimental observations. Despite its simplicity, the analytical approach proposed by [START_REF] Huckert | Approches expérimentale et numérique du dimensionnement de renforcements géosynthétiques sur cavités et inclusions rigides[END_REF] In the case of a granular soil, the main finding is that:

• Among the different propositions of load distribution, the inverted triangular or parabolic shapes appear to be the most adapted to describe accurately the experimental results.

• An additional surface loading leads to a change in the load distribution on the GR close to the middle of the cavity in the case of small H/B ratios (i.e. 0.25 in Test 4) whereas the shape is not modified for larger values of H/B.

• The GR displacement and deformation in the anchorage zones are well accounted for. The intensity of the load transfer mechanism can be approached by Terzaghi's formulation using appropriate value for the ratio between horizontal and vertical stresses (K = 1.3).

• Depending of the ratio H/B the load transfer coefficient SRR varies from 35% to 70%.

A preliminary analysis has also been proposed in the case of a cohesive soil showing that:

• The load distribution on the GR is only measured when a surface load is applied and large enough to induce the failure of the soil layer that bridges the cavity.

• The resulting block applies loads on the GR that are localized close the cavity edges.

• The measured strains and tension in the GR validate to a certain extend the analytical approach proposed by [START_REF] Huckert | Experimental and numerical approaches of the design of geosynthetic reinforcements overlying voids[END_REF].

) K = K a = (1 -sin φ)/(1 + sin φ) Terzaghi (1943) K = 1 Handy (1985) K = 1.06(cos 2 θ + K a cos 2 θ)
(x) = 2q 0 B x q(x) = ( 12q 0 B 2 x 2 ) Deformed shape Y (x) = 2q 0 3BT h (x 3 -B 3 8 ) Y (x) = q 0 T h ( x 4 B 2 -B 2 16 ) β = dy dx (x = B/2) β = 2q 0 B 4T h β = q 0 B 2T h Max. tension T max = 2q 0 B 4β 1 + [ 4βx 2 B 2 ] 2 T max = q 0 B 2β 1 + [-β x 3 B 3 ] 2 Max. deflection f = βB 6 f = βB

  (sinkhole) of ground support generates a local redistribution of stress and deformation within the soil mass, often known as arching. Soil arching is defined by Mc Nulty

  have experimentally simulated sinkholes under both granular and cohesive soil layers reinforced with geosynthetic. Different collapse mechanisms have been underlined for granular and cohesive backfills. A particular load transfer mechanism has been proposed as well as the corresponding deflected shape of the geosynthetic layer above void in the case of cohesive soil, while literature analytical assumptions inspired from the granular backfill case have been shown unsuitable for the design of the geosynthetic layer. Therefore, researches investigating such behavior are strongly needed.TRAPDOOR APPARATUS AND MODELTEST 4 HASSOUN et al., March 29, 2018 Trapdoor apparatus As shown in Fig.1, the test set-up consists in a rigid base, a subsidence simulator and a box. The box is 1000mm in length, 400mm in width, and 200mm in height. Only shallow embankments are considered in this study. The front wall of the box consists in a transparent Plexiglas plate used to enable visualization of the models during testing, while the other sides were made of wood. The subsidence simulator (Trapdoor) has a fixed width B = 200mm and is located in the bottom of the box in its center. The trapdoor can move downwards to 50mm with a speed of nearly 0.016 mm/s. This causes the creation of a void within the soil mass with a predetermined vertical displacement δ measured with linear variable displacement transducer (LVDT). The displacement transducer has an accuracy of 0.05 mm. The vertical effort, called (P), applied by the soil material on the trapdoor was measured using a load cell with 1N of precision. In order to limit the wall boundary effects on the measurements, (P) is measured in the central part of the trapdoor of 200 mm x 200 mm. The corresponding average vertical stress (p) is deduced from (P) and the dimensions of the central part of the trapdoor. Test model is designed to reproduce a prototype scaled by factor n = 10 on lengths, where the real state of stress is not respected. This type of scaling neglects the influence of the modulus of the backfilled soil and of gravity on the backfill behavior. Rigorous scaling factors

  Fig.7 shows the contours of vertica granular soil displacements (S v ), for trapdoor displacements equal to (a) 0.75 %, (b) 1.65 %, (c) 3.5 % and (d) 9.5 % of the trapdoor width B. The general pattern of displacements is shown to change with δ. Loading curve response and vertical displacement fields illustrate the three main stages of the behavior of a shallow granular backfill.Initial or maximum load transfer: as soon as the trapdoor is moved, the soil grains come

  The mechanism observed for the three backfill heights (H/B = 0.25, 0.5 and 0.75) used in the previous section have been reconsidered this time with a geosynthetic reinforcement layer. The same GR layer has been used in the three tests. Direct measurements of the displacement have been performed on both granular soil layer and along the GR layer using DIC technique. The influence of GR on the displacement field in the overlying soil layer is qualified, and analytical estimations of the load acting on the GR layer proposed in the literature are compared to the experimental results. Placed directly on the rigid base of the trapdoor model, the GR layer has been anchored on each 12 HASSOUN et al., March 29, 2018 side over 400mm from the trapdoor edge under only the weight of the soil backfill. Data acquisition stops when the trapdoor loses any contact with the GR layer.

  .11 concludes that the load distribution in the GR over the trapdoor determined by the established experimental analysis procedure approaches for Tests 5 and 6 an inverse triangular distribution. Arching mechanism effect is well observed. Contrary to RAFAEL (1997), BS8006 (2010) and EBGEO (2011) hypothesis, a non uniform distribution was shown. The obtained shape of the load distribution meets experimental and numerical conclusions ofVillard et al. (

  assumes a collapse of the cohesive soil layer in blocks when subjected to a concentrated load at the soil surface. The principle of this model is then to transfer the weight of the collapsed soil block and the surface overload to the GR layer by means of two vertical forces (2P) defined per meter of width of the geosynthetic sheet 18 HASSOUN et al., March 29, 2018

  describes in a reasonable way the experimental results. The strong hypotheses made on the geometry of the deformed GR layer (straight segments) and on load transfer explain the gap on the maximum tensile strain and deflection of the GR layer which presents a slight curvature in the experiment. The application of the load via two localized forces can also be questioned, in particular when the intensity of the overload applied at the soil surface becomes important compared to the weight of the collapsed soil blocks and the stress becomes larger near the centre of the cavity. 19 HASSOUN et al., March 29, 2018 CONCLUSION A series of experimental tests using classical trapdoor apparatus have been carried out to investigate soil arching effect, load transfer mechanism and soil -geosynthetic reinforcement interaction due to local subsidence for both granular and cohesive backfills. A specific methodology was developed in this work based on the treatment of the experimental results obtained by photogrammetry technique (DIC) to determine the interaction forces between GR and soil.

  Ladanyi et al. (1969) K = K E = (1 -sin 2 φ)/(1 + sin 2 φ) Aubertin et al. (2003) K = K p = tan 2 θ Huckert et al. (2014) K = 1.3 Chen R.X et al. (2010) K = (cos 2 θ + K p sin 2 θ)/(sin 2 θ + K p cos 2 θ) φ is the soil friction angle and θ = π/4 + φ/2 31 HASSOUN et al., March 29, 2018

8B:

  is the cavity s diameter T h : is the horizontal component of the tension force in the GR (Equations are given for positive horizontal coordinates x, x = 0 at the center of the cavity) 32 HASSOUN et al., March 29, 2018TABLE 7. Analytical approach of the experimental results in Tests 4-6 in terms of maximum GR deflection and strain GR displacement at the edge of the cavity U A = 0 GR displacement at the edge of the cavity U A granular embankment with H/B = 0.

Fig 1

 1 Fig 1

Fig 19 FigureFig. 1 .Fig. 2 .Fig. 3 .Fig. 4 .

 191234 Fig 19

Fig. 5 .Fig. 6 .Fig. 7 .Fig. 8 .Fig. 9 .Fig. 10 .Fig. 11 .

 567891011 Fig. 5. Experimental validation results

Fig. 12 .Fig. 13 .Fig. 14 .Fig. 15 .Fig. 16 .Fig. 16a .Fig. 16b .Fig. 17 .Fig. 18 .Fig. 19 .

 121314151616a16b171819 Fig. 12. Evolution with loading of vertical stress distribution transmitted to GR for Test 4

Table 2

 2

	). In order to obtain the stress levels during the triaxial tests
	5	HASSOUN et al., March 29, 2018

TABLE 1 .

 1 Scaling factors of model parameters

	Parameter	Symbol Unit	Scaling factor
	Length	l	m	1/n
	Gravity	g	m/s 2	1
	Stress	q	k Pa	1/n
	Deformation	ǫ	-	1
	Specific weight	γ	k N/m 3 1
	Friction angle	φ	°1
	Young s modulus	E	Pa	1/n
	Poisson s ratio	v	-	1
	Tensile stiffness of geosynthetic J	k N/m	1/n 2

TABLE 2 .

 2 Mechanical properties of the backfill soils

	Soil		Shear test		Grading characteristics
	γ Gravel 15.2	-	53.6/40.1	0	12.5	5	1.61
	Sandy-Clay 13.8	15	35	5.5	2	< 0.002	-

d (k N/m 3 ) w(%) φ pic / φ r es (°) C (k Pa) d max (mm) d min (mm) C u

TABLE 3 .

 3 Test program

	Backfill material Test H/B Description
		1	0.25 Unreinforced
		2	0.5	Unreinforced
	Gravel	3 4	0.75 Unreinforced 0.25 Reinforced
		5	0.5	Reinforced
		6	0.75 Reinforced
	Sandy-Clay	7 8	0.25 Unreinforced 0.25 Reinforced

TABLE 4 .

 4 Derived expressions for the calculation of the stress reduction on an underground structure caused by arching

	Reference	Equation
	Stress reduction at maximum arching
	Engesser (1882)	p p 0 = B H (	K a 2 tan φ B H K a	+ tan φ 6 )
	Bierbaumer (1913)	p p 0 = B 4H tan φ
	Evans (1983)	p p 0 = 1 -H B tan φ
	Iglesia et al. (2014) p p 0 = B H (	K E 2 cot φ+ B H K E	+ cot φ 6 )
	Guido et al. (1987) Carlson (1987)	p p 0 = B 4H p p 0 = B 2 4H tan φ

TABLE 5 .

 5 Literature expressions of K

	Reference	Expression of K
	Marston and Anderson (1913

TABLE 6 .

 6 Equilibrium equations used in analytical models for different load distributions

		Inverted triangular distribution shape Parabolic distribution shape
	Stress equation	q

  25 (Test 4) : with a non zero stress value at the center of the cavity, determined based on experimental results for load distribution in Tests 4-6

	f max (mm)	10.54	12.59	13.48	14.24	14.46	16.76	17.89	18.84	18.8 ± 0.30
	ε max (%)	1.32	1.45	1.46	1.55	1.02	1.14	1.15	1.22	1.10 ± 0.08
				Reinforced and unloaded granular embankment with H/B = 0.5 (Test 5)
	f max (mm)	11.88	14.19	14.40	15.54	14.47	16.94	17.26	18.41	17.9 ± 0.60
	ε max (%)	1.69	1.86	1.8	1.94	1.45	1.60	1.56	1.68	1.73 ± 0.10
				Reinforced and unloaded granular embankment with H/B = 0.75 (Test 6)
	f max (mm)	12.36	14.76	14.77	16.03	14.92	17.49	17.56	18.85	18.30 ± 0.75
	ε max (%)	1.84	2.01	1.94	2.09	1.59	1.75	1.69	1.83	1.96 ± 0.11
	NZ

* 33 HASSOUN et al., March 29, 2018

TABLE 8 .

 8 Comparison of experimental and analytical results inTest 8 (H/B = 0.25) , ε B T A , T B f ε A , ε B T A , T B

			Loading Q 1	Loading Q 2
	Results	f ε A mm %	kN/m mm %	kN/m
	Experimental 14.1 1	0.133 22.3 1.75	0.210
	Analytical	11	1.35	0.183 18.6 2.11	0.254

 HASSOUN et al., March 29, 2018

NOTATION

The following symbols are used in this paper: T j = total tensile force defined per unit width at any point j of the GR (N/m); T v, j , T h, j = vertical and horizontal components of the tensile force at any point j of the GR (N/m); U A = horizontal displacement of the GR at the edge of the cavity (mm); W L , W P ,I P = liquid limit, plastic limit, plasticity index of sandy-clay soil respectively (%); W = moisture content in soil (%);


Philippe Gombert
email: philippe.gombert@ineris.fr

Ondra Sracek

Nikolaos Koukouzas

Grzegorz Gzyl

Susana Tunon Valladares

Robert Fraczek

Christoph Klinger

Arkadiusz Bauerek

Juan Enrique

Alvarez Areces

Sinead Chamberlain

Susana Tuñon Valladares

Robert Frączek

Álvarez Areces

Krzysztof Paw

Łukasz Pierzchała

An Overview of Priority Pollutants in Selected Coal Mine Discharges in Europe

Keywords: coal basin, pollution, database, emission limits

établissements d'enseignement et de recherche français ou étrangers, des laboratoires publics ou privés.

Introduction

During coal mining, groundwater is pumped in order to lower the water table and permit exploitation. After abandonment and cessation of pumping, the natural rise of the water table inundates or floods the mined voids [START_REF] Banks | Modeling and monitoring of mine water rebound in an abandoned coal mine complex: Siersza Mine, Upper Silesian Coal Basin, Poland[END_REF]Wolkersdorfer and Bowell 2004). This flooding, also called "rebound", can create a mine aquifer, as connections between the subsurface and surface waters, often used for drinking water, are reestablished (Younger et al. 2002). Also, during mining, the strata are exposed to an oxidizing environment, but the return to pre-mining water table levels restores low-oxygen conditions. Flushing of oxidized minerals can then release contaminants [START_REF] Kopřiva | High arsenic concentrations in mining waters at Kanˇk, Czech Republic[END_REF], which may pollute the local groundwater and surface waters [START_REF] Gzyl | Verification of the "first flush" phenomenon in mine water from coal mines in the Upper Silesian Coal Basin, Poland[END_REF]. Concentrations of dissolved major ions in untreated mine discharge are frequently very high (Cravotta 2008a;[START_REF] Cravotta | Priority pollutants and associated constituents in untreated and treated discharges from coal mining or processing facilities in Pennsylvania, USA[END_REF].

Contamination of mine water has been studied as part of the European MANAGER project, which was partly funded by the Research Fund for Coal and Steel (RFCS), to assess "Management of mine water discharges to mitigate environmental risks during the post-mining period". Mitigation of the environmental risks of mine water on aquatic ecosystems involves decreasing the discharge of contaminants. The main aim of the MANAGER project was to develop and implement cost-effective, sustainable, and innovative mine water treatment technologies to mitigate environmental risks after mine closure. This paper presents the results of the first part of the project, which focused on identifying mine water contaminants that cause the greatest environmental risks.

The main goals of this stage of the project were: a)

to gather information about mine water contaminants in various European coalfields in a coherent and comprehensive database; b)

to select mine water contaminants that pose risks to aquatic ecosystems, taking into consideration local conditions and the objectives of the Water Framework Directive.

The study included numerous European coalfields from the 6 countries participating in the MANAGER project, i.e. France, Poland, Germany, Spain, Greece, and the United Kingdom (UK) (Fig. 1). A total of 97 representative coal mine waters were investigated and more than 26,000 analysis or physicochemical measurements were gathered.

Figure 1 will be placed near here during the printing process

Review of Selected Coal Mine Waters in Europe

FRANCE

INERIS, the French national institute for risk assessment, prepared a water composition database for the French coalfields, using data supplied by "Charbonnages de France" (the former French coal authority), GEODERIS (the French public interest group for risk management related to post-mining), INERIS files, and from scientific publications. The database contains 81 discharges with 450 flow measurements and 274 water analyses.

Coal mining began in France in the 13 th century and ended in 2004 [START_REF] Bouzenot | Post-mining in France[END_REF]. During this period, 4,700 Mt of coal were extracted from nine major coal basins: most of them were in the Nord-Pas de Calais and the Lorraine basins, which respectively extend to Belgium and Germany. About half of all the French coal basins are now flooded and most of them discharge to surface waters. The main pollutants are sulphate and iron [START_REF] Fabriol | La gestion de l'eau des mines en phase post-extractive[END_REF].

The coal mine aquifers contain about 3×10 9 m 3 of mine water and their total outflow is estimated to exceed 120 Mm 3 /a, which is equivalent to about 6% of the total river flow in France. 95% of the coal mine waters flow into rivers, while 4% flows into lakes, and 1% flows directly into the sea.

GERMANY

Coal mining in Germany is very diverse, comprising both lignite and hard (anthracite and bituminous) coal. Whereas opencast lignite mining (not discussed in this paper) will continue post-2018, the last active hard coal mining will cease in 2018. The German hard coal measures extend east to west, following the Rhenohercynian and Saxothuringian strike, and extend into Belgium and France in the west and Poland in the east.

The remaining active mines can be found in the Ruhr Area (Rheinisch-Westfälisches Steinkohlerevier) and at Ibbenbüren. As in many other coal mining areas, the mines have varied hydraulic status; mines currently being flooded are located close to mines that have already been flooded and are discharging water. There is still no discharge during the flooding because pumping is temporarily being maintained. Furthermore, the groundwater level is not uniform; many abandoned and partially flooded mines in the Ruhr Area have a temporary water level below -600 m b.s.l. This means that most of the German coal mines aquifers have not reached equilibrium conditions, which would allow for extrapolation of long term discharge flow rates and quality. Instead, an "unsaturated zone" exists above the mining water table, where oxidative reactions play an important role, and the potential for post-flooding mobilization of oxidation products is still considerable.

The Ruhr Area is the largest of the German coal fields and its discharges exceed the sum of all of the other German hard coal mining areas. Pumping rates in this area has ranged between 48 and 54 Mm³•a -1 during the last decade. However, its main environmental impact is related to the flow rate of the receiving water course. Therefore, future dewatering of the Ruhr Area plans within in frame of mine closure and cessation of pumping activities for withdrawal from some actual discharges into local rivers (Emscher, Ruhr, and Lippe), and installation of an underground drainage system via mining voids to a major discharge into the Rhine River, which can provide high dilution rates [START_REF] Fischer | Konzept zur langfristigen Optimierung der Grubenwasserhaltung im Ruhrrevier der Aktiengesellschaft -Proc[END_REF].

GREECE

Many coal basins exist in Greece: the major coal mining areas are found in or near Ptolemais, Florina, Drama, Elassona, Megalopolis, Ioannina, and Komnina [START_REF] Koukouzas | Coals of Greece: Distribution, Quality and Reserves[END_REF]. Power generation in Greece relies mainly on lignite combustion (about 50% of power generation). The major generating stations are situated in Ptolemais-Kozani, Amyntaio, Florina, and Megalopoli, and are owned by the Public Power Corporation (PPC). Data concerning mine water discharges were obtained from the PPC Environmental Protection and Land Reclamation Sector Determination agency, and relate to the Ptolemais and Meliti opencast lignite mines. Water samples were collected and analysed from wells, pumping stations, and streams. The main contaminants were As, Cd, Cr, Ni, Zn, Pb, Hg, and SO4.

POLAND

Information on the impact of mine water discharges on the water environment in Poland is dispersed and not readily available. Under the framework of this project, multi-year data, including basic hydrochemical parameters, specific contaminants, and priority substances in 15 mine water discharges, were gathered by the Central Mining Institute in Katowice.

The Upper Silesian Coal Basin (USCB) is located in southern Poland and is one of the largest hard coal basins in the world [START_REF] Rozkowski | Impact of mining on groundwater chemistry in the Upper Silesian Coal Basin (Poland)[END_REF]. Hard coal has been mined in the USCB since about 1540. Recently (2014), there were 29 active hard coal mines, 25 abandoned coal mines still being dewatered, and one abandoned coal mine where dewatering has stopped and which consequently has been partially flooded. The amount of groundwater being pumped from the coal mines totals 209 Mm 3 /a. Mine dewatering is the main cause of the USCB's hydrochemical regime disturbances. Mine waters from these mines are typically discharged into tributaries of the upper Wisła (Vistula) and upper Odra (Oder) rivers [START_REF] Janson | The occurrence and quality of mine water in the Upper Silesian Coal Basin, Poland[END_REF]). These mine waters are generally alkaline, but have elevated concentrations of sodium, chloride, and sulphate [START_REF] Pluta | Origin of brines in the Upper Silesian Coal Basin (Poland) inferred from stable isotope and chemical data[END_REF]. Radium is also often present in waters with low sulphate concentrations [START_REF] Chalupnik | Radium removal from mine waters in underground treatment installations[END_REF].

In the USCB, about 480 million metric tons of mine waste is stored in approximately 160 waste heaps (Szczepańska and Twardowska 1999;GUS 2015). Research carried out within the MANAGER project revealed that some of them may be considered to be "hot spots," releasing acidic runoff waters with high concentrations of sulphate, iron, aluminium, zinc, and nickel into the groundwater [START_REF] Bauerek | The concept of the pre-treatment of water runoff generated on mine waste pile of ZG Janina in Libiaz[END_REF]. Quantitative data concerning this phenomenon is sparse, although the data from the eastern part of Silesia shows that the environmental impact of contaminated runoff waters is locally important, and that it increases substantially during warmer periods of the year, when heavy rains occur [START_REF] Bauerek | Chemical composition of surface runoff from flotation wastes of Zn-Pb ore formation of the Mississippi Valley-type, Olkusz, southern Poland[END_REF]. Some of the waste heap discharge waters are alkaline, but they frequently have high concentrations of sulphate, ammonia, and metals [START_REF] Sracek | Evaluation of the impacts of mine drainage from a coal waste pile on the surrounding environment at Smolnica, southern Poland[END_REF]).

SPAIN

In central Asturias, a large coal field has been mined since the early 19 th century in several exploitation units [START_REF] Ojeda | Asturias en la industrialización española, 1833-1907[END_REF]. Most of the strata of the Central Carboniferous River basin can be considered effectively impermeable, except for some intensely fractured siliciclastic and carbonate rocks, through which infiltration takes place [START_REF] Martos | Proc, XII Congreso Internacional de Energía y Recursos Minerales[END_REF]. There are more than 73 coal shafts in this area, and more than 300 mines and 2000 mine adits in this mountainous area [START_REF] Fandos | El yacimiento de HUNOSA en la Cuenca Carbonífera Central Asturiana. Informe inédito[END_REF]). Currently only four collieries remain in operation in two coal fields (San Nicolas and Santiago Colliery in the Caudal Field, and Maria Luisa and Carrio Colliery in the Nalón Field).

Coal has been mined in Spain for more than 200 years, generating a system called "anthropogenic aquifer," which is still being dewatered by several wells [START_REF] Pendás | El agua en los procesos de cierre de minas en Asturias[END_REF]. The low sulphur content in the coal, as well as the abundant carbonate materials, which act as a buffer, reduces the problem of acidic and ferruginous water. However, concentrations of sulphate, bicarbonate, calcium, magnesium, and other ions can be elevated [START_REF] Martos | Proc, XII Congreso Internacional de Energía y Recursos Minerales[END_REF]). Hunosa, a state-owned coal company, is responsible for the mine water discharges at locations where pumping stations are installed. The total outflow exceeds 37 Mm 3 /a for 15 discharge points, all of which are being monitored. In addition, four old mine water discharges have been periodically monitored.

UNITED KINGDOM

Underground and opencast mining is ongoing in the UK, but on a considerably reduced scale when compared with historic mining rates in the 19 th and 20 th centuries. The Coal Authority manages the effects of past coal mining, including mine water pollution, and has been regularly monitoring mine water levels and quality across the UK for over 20 years [START_REF] Crooks | A sustainable approach to managing the treatment of mine waters associated with historic mining[END_REF]. Most of the mine waters in the UK are circumneutral in pH and ferruginous, and range from fresh to hypersaline, depending on coal field geology, geometry, and mining history. Iron and, to a lesser extent, manganese are the principal metal contaminants in UK mine waters. Other metal and metalloid contaminants are typically only present at relatively low concentrations.

The Coal Authority operates 70 mine water treatment sites throughout Britain, of which 53 remediate existing mine water discharges and 17 treat water on-site to prevent new discharges of contaminated water [START_REF] Lawrence | The Coal Authority annual report and accounts 2014-15[END_REF]. Together, the 70 mine water treatment sites prevent over 4,000 tons per year of iron from discharging into watercourses, and so improve and protect the UK's rivers and streams. Some water treatment operations also exist to protect important drinking water aquifers.

Materials and Methods

The first step for selecting priority substances for mine water was to develop a coherent and comprehensive methodology and the tools for gathering, storing, and analysing mine water contaminant data from different coalfields in Europe. This consisted of the following activities:

• Gathering available information, from all partners, about contaminants in rivers and mine water discharges from different coalfields in Europe, and collecting additional samples where data were lacking; • Listing of 58 major and minor ions, metallic trace elements, and physicochemical parameters, based on all of the measured or analysed factors; • assigning the following attributes to each discharge point:

o individual name, o information about geographical coordinates, country, coal district, name of mine, type of mining and mine status, and quantity and altitude of the discharges, o name, code, and mile of river or water body where a discharge was localized.

• Standardization of units for each parameter (e.g. nitrogen compounds can be expressed in mgN/L or mgNO3/L, depending on the country); • Replacement of concentrations under limits of detection (LOD) by systematically applying LOD/2; this was necessary because the studied countries had different LOD and/or the LOD may have varied over time; • Performing a basic statistical evaluation based on multi-year observations for each mine water discharge point; • Gathering information about emission limit values (ELV) and acceptable status ranges, according to local regulations introduced by the Water Framework Directive (WFD) as environmental quality standards (EQS).

The next step was to analyse the information gathered from the different coalfields to identify the important contaminants from the point of view of the WFD objectives. Water parameters that could cause environmental risks under specific local conditions, and contaminant concentrations in discharges that exceeded local pollutant emission limits, were also selected as mine water priority substances (MWPS). Selection of MWPS was based on two multi-year average contaminant concentrations (MACC) criteria, where a mine water discharge was causing a significant deterioration in water body status:

• The MACC of a specific contaminant in mine water discharges was exceeding an established ELV or EQS, based on local regulations or the WFD, regardless of initial water quality; • The MACC of a specific contaminant in mine water discharges was not exceeding established ELV or EQS, but the water body quality exceeded one or more of these thresholds downstream of the mine water discharge point, based on local regulations or the WFD.

Results and Discussion

The main goals were achieved. A coherent and comprehensive database for storing and analysing mine water contaminant data from the different European coalfields was developed, and mine water priority substances that pose the risks to the aquatic ecosystems were identified. The database contains:

• Information about MACC at 97 mine water discharge points,

• ELV for 62 discharge points and EQS regarding the WFD,

• A list of mine water priority substances for each country, and • Substantiation of the selected mine water priorities.

Contaminants were identified that can cause environmental risks, such as deterioration of river ecosystems, when their concentrations exceed ELV for European coal mine waters (Table 1). Sulphate and iron were identified as priority substances in almost all of the countries, while several metallic and metalloid trace elements (e.g. As, Cr, Zn) were problematic only in some. High concentrations of chloride (and sodium, not shown) were linked to the Na-Cl brines associated with coal mining in Germany, Poland, and UK. High 226 Ra 2+ concentrations in Germany and Poland were associated with high Ba/low sulphate water.

Table 1 will be placed near here during the printing process Box/whisker plots for the principal priority substances shown in Table 1 can be observed in Figure 2. Concentrations values and basic statistical parameters are provided in Table 2. More than 26,000 analyses were gathered for 97 mines spread over the studied countries, ranging from 5 to 26 by country. These mines represent a total discharge of about 220 m 3 /s, i.e. almost 7,000 Mm 3 /a.

Figure 2 and Table 2 will be placed near here during the printing process

The pH was always higher or equal to 6.0, with an average of 6.50 (Table 2). Sulphate (Fig. 2a) averaged 713 mg/L and iron (Fig. 2b) averaged 20.9 mg/L, for all of the country's mine discharges. These elements are produced by the oxidation of sulphides like pyrite [START_REF] Blowes | The geochemistry of acid mine drainage[END_REF], which produces 1 mol of Fe 2+ and 2 mol of SO4 2-for each mol of pyrite that is oxidized. The reaction also produces two mol of H + , which can react with carbonate rock, such as calcite and dolomite, if any is present in in the nearby strata.

Furthermore, the released Fe 2+ is oxidized and precipitates at pH > 3.0 as ferric hydroxide, which is also an acidproducing reaction [START_REF] Langmuir | Aqueous Environmental Geochemistry[END_REF]. Meanwhile, the neutralization reactions can produce rather high concentrations of Ca 2+ (Fig. 2a, Table 2), which in this study averaged 384 mg L -1 , and Mg 2+ (Fig. 2a), which averaged 169 mg L -1 . The Ca 2+ may precipitate with sulphate in neutralized waters as gypsum, CaSO4•2H2O, but sulphate in neutral pH mine waters generally behave in a conservative way because saturation with respect to gypsum is not typically reached. Jarosite, KFe3(SO4)2(OH)6, precipitates in mine waters at a pH <3.0 [START_REF] Langmuir | Aqueous Environmental Geochemistry[END_REF][START_REF] Blowes | The geochemistry of acid mine drainage[END_REF] and, thus, cannot control sulphate concentrations in the generally circumneutral studied mine discharges.

Manganese was also present in most of the mine discharges (Fig. 2b), with an average value of 3.69 mg/L.

Chloride together with sodium (not shown), averaged 2,747 mg/L, and was most commonly associated with the coal deposits in Germany, Poland, and the UK (Fig. 2a). In Germany and Poland, these major ions come from fossil sea water but in the UK, they are probably due to sea water inflow into submarine mine adits (Younger et al. 2002). Chloride behaves conservatively in mine discharges and its concentrations generally decreases only by dilution and dispersion [START_REF] Motyka | Influence of contaminated Vistula River water on the groundwater entering the Zakrzowek limestone quarry, Cracow region, Poland[END_REF]. Boron (Fig. 2b), which averaged 0.96 mg/L, is also linked to fossil sea water in Germany and Poland.

Lead (Fig. 2c) was present in all of the mine discharges except for Spain at relatively high concentrations (averaging 54.2 μg/L), compared to its EQS (1.2 µg/L -1 ). Cadmium was also present in several of the mine discharges, presumably as a consequence of dissolution of carbonates. Its average concentration (5.5 µg/L) was also high relative to its EQS, which depends on water hardness; the EQS for Cd varies from less than 0.08 µg/L if the hardness (as CaCO3) is < 40 mg/L to 0.25 µg/L if the hardness exceeds 200 mg/L. Typically, there is no solubility control on these contaminant concentrations but they can be attenuated by adsorption on ferric precipitates (Cravotta et al. 2008).

Significant concentrations of Hg (Fig. 2c) occur in coal mine discharges in several countries, averaging 0.3 μg/L, which exceeds its EQS (0.05 µg/L), which means that many of the mine waters exceed this threshold.

High concentrations of Ni, As, and Cr (Fig. 2c) were less common and occur mostly in mine discharges in the UK, France, and Poland. Their respective average concentrations and EQS were 37.4 μg/L for Ni (EQS: 4.0 µg.L -1 ), 18.6 μg/L for As (EQS: 4.2 µg/L), and 8.3 µg/L for Cr (EQS: 3.4 µg/L -1 ). Generally, there is no solubility control on the concentrations of these elements, but they also are adsorbed on precipitated ferric phases (reaction 3) at about neutral pH (Cravotta 2008b;[START_REF] Sracek | Attenuation of dissolved metals in neutral mine drainage in the Zambian copper belt[END_REF]) because desorption of oxyanions occurs at higher, alkaline pH, i.e. for As at pH above 8.2 [START_REF] Ravenscroft | Arsenic Pollution. A Global Synthesis[END_REF]). These trace elements do not represent a global environmental problem for coal mine water.

Zn was also found in mine water in Poland and Germany with sometimes very high concentrations (respective maxima at 17,800 and 25,000 µg/L), but there is no EQS for this element. Just like Ni, Zn occurs as a cation and there is generally no solubility control on its concentrations in mine discharge.

Se can be an important issue in coal mine water (Vesper et al. 2008;Yudovich and Ketris 2006). However, it was only analysed in coal mine waters of Greece and Spain with an average concentration of 3.5 µg/L. Again, there is no EQS for this element.

Finally, 226 Ra 2+ (Fig. 2c), which averaged 1,190 µg/L, is of particular interest, but it was present only in mine water discharges in Germany and Poland with significant differences (respective maxima were 7,300 and 244 µg/L). In these waters, sulphate has been reduced by reactive organic matter (Appelo and Postma 2005) and so there is no solubility control of 226 Ra 2+ . In high sulphate waters, 226 Ra 2+ co-precipitates with gypsum and/or barite [START_REF] Chalupnik | Radium removal from mine waters in underground treatment installations[END_REF]. Relatively high concentrations of Ba (Fig. 2b), respectively averaging 26.1 and 5.1 mg/L, also occur in low sulphate waters. High Ra and Ba concentrations in low sulphate waters are consistent with findings of Cravotta (2008b) coal mine discharges in Pennsylvania.

Conclusions

More than 26,000 mine water analyses were gathered from 97 mines from the countries involved in the project, i.e. France, Greece, Germany, Poland, Spain, and the UK. In each country, the amount of water stored in coal mine aquifers totals over 10 9 m 3 , and the combined mine water outflow (natural in the case of abandoned mines or artificial, due to pumping, in the case of active mines) is around several hundred Mm 3 per year. All the studied mines represent a total discharge of about 7,000 Mm 3 /a.

The quantitative influence of coal mine water discharges is therefore substantial, and evaluating their environmental impact on surface and underground water is important. The coal mine water quality evaluation notes a first group of pollutants, represented by SO4 and Fe, which were shared by all countries; a second group of frequently observed pollutants (Mn, Cd, Pb); and a third group, that is specific to certain mine waters (Cl, B, Ba, As, Cr, Ni, Zn, Hg, Ra). Some discharges show substantial mineralization, frequently accompanied by high concentrations of metallic or metalloid trace elements, and sometimes by radium and barium, typically in low sulphate waters. Chloride and boron have a special status because they are linked to the impact of fossil sea water brines. The remaining trace metals generally were present at low concentrations, but since their EQS are very low, they often exceed these thresholds.

The selection of mine water priority substances was based on a comparison of river water quality standards with mine water quality data. This means that there is no certainty that the mine discharges of selected substances will cause deterioration of river status or adversely affect aquatic ecosystems. The mine water quality database that we developed identified the contaminants that are of primary concern from the point of view of achieving WFD objectives, and that should be considered during environmental risk assessment.

Szczepańska J, Twardowska I (1999)

Figure Captions Fig. 1 :Fig. 2 :

 Captions12 Figure CaptionsFig. 1: Location of the studied coalfields in Europe: a) Major elements; b) Minor elements; c) Trace elements

  Fig 1

  Distribution and environmental impact of coal-mining wastes in Upper Silesia, Poland. Environ Geol 38(3): 249-258. Vesper DJ, Roy M, Rhoads CJ (2008) Selenium distribution and mode of occurrence in the Kanawha formation, southern West Virginia, USA. Int J Coal Geol73: 237-249 Wolkersdorfer C, Bowell R (2004) Contemporary reviews of mine water studies in Europe, Part 1, Mine Water

	Environ 23: 162-182
	Younger PL, Banwart SA, Hedin RH (2002) Mine Water, Hydrology, Pollution, Remediation. Kluwer Academic
	Publ, Dordrecht (NL)
	Yudovicvh YE, Ketris MP (2006) Selenium in coal: a review. Int J Coal Geol 67:112-126

Table 1 .

 1 Global statistics of priority pollutants. EQS (Environmental Quality Standards) are EQS-AA (Annual Average) for all trace elements but PNEC (predicted no effect concentrations) for As and Cr (under European Directive 2013/39/UE). No. = number of samples.

			Major ions (mg/L)	Minor ions (mg/L)	Trace elements (µg/L)	Bq/L
		pH Ca Mg Cl SO4 Fe Mn B Ba As Cd Cr Pb Ni	Zn Hg Se Ra
	No.	2583 2470 2419 2972 2848 2538 2507 1977 1637 245 1269 846 1702 1758 1348 236	30	91
	Min.	6.0	1	1	1	0 0.01 0.01 0.01 0.01 bdl bdl bdl bdl bdl	bdl 0.1	0.3 0.01
	Max.	8.9 5070 1320 71400 4680 330 77.0 13.0 5800 900 130 155 2280 421 25000 2.5	10 7.30
	Ave.	6.5 384 169 2756 713 20.9 3.69 0.96 9.77 18.6 4.5 8.3 54.2 37.4	440 0.3	3.5 1.19
	EQS	6-9				4.2 *	3.4 1.2 4.0	7.8 0.05
	* depending on water hardness: from <0.08 µg/L if hardness (as CaCO3) < 40 mg/L to 0.25 µg/L if hardness  200 mg/L; bdl
	= below detection limit

Table 2 .

 2 Statistics of priority pollutants by country; Number (No.) represents the number of water samples analysed in the mines referenced in each country; "bdl" is below detection limit)

	Country	Major ions (mg/L) Minor ions (mg/L) Ca Mg Cl SO4 Fe Mn B Ba As Cd Cr Pb Ni Zn Hg Se Ra Trace elements (µg/L) Bq/L
		No. 60 60 60 60 60 60 0	0	60 60 60 60 60	60	60	10	0
	Greece	Min. 44 15 Max. 184 49 78 467 1.21 0.10 6 0 0.03 0.01		0.2 0.2 1.7 4.0 5.0 0.5 3.2 1.3 28.0 12.0 24.0 40.0 bdl 10.0 bdl 10.0
		Ave. 94 28 16 144 0.18 0.03		1.1 1.0 3.9 5.1 8.7 6.8	bdl 10.0
		No. 67 67	0	65 65 63 40	0	40 40 40 40 40	0	40	20	0
	Spain	Min. 46 8 Max. 319 100		24 0.01 0.01 0.01 1747 9.05 0.92 0.19		0.3 0.1 0.3 0.3 bdl 9.4 0.1 0.6 4.8 2.9		bdl 0.3 bdl 0.4
		Ave. 129 53		416 1.87 0.26 0.07		0.6 0.1 0.3 0.4 1.4		bdl 0.3
		No. 231 231 606 606 243 213 21 44	0 198 198 200 199 201	15	0	10
	Poland	Min. 1 21 41 119 0.01 0.01 2.23 0.12 bdl bdl 0.6 0.4 0.3 0.3 0.1 Max. 1062 707 36072 2868 39.0 8.27 4.25 7.70 bdl 130 155 560 421 17800 2.5		0.060 0.244
		Ave. 263 162 7579 783 3.59 1.30 3.19 5.06 bdl 9.3 18.9 60.8 43.3 287 2.3		0.160
		No. 1427 1378 1427 1225 1313 1427 1340 986 53 385 0 848 900 484
	United
	Kingdo
	m

Acknowledgements: This paper was issued from the research project MANAGER (Management of mine water discharges to mitigate environmental risks for post-mining period). The authors thank the Research Fund for Coal and Steel (RFCS) of the European Commission for helping to fund this project (RFCR-CT-2013-00005).