Vincent Renaud

Lynda Driad

Marwan Al Heib

J F Noirel

BACK-ANALYSIS OF THE ROCKBURST IN FRIEDA 5 AND DESIGN PROGRAM FOR PANEL IN SEAM 6 (HBL -FRANCE)

This paper develops the studies carried out after the fatal accident occurred during the mining of Frieda5-1250 panel at Merlebach mine section of HBL (East France). On the 21 June 2001, a violent rockburst (local magnitude of 3.6) has affected the panel 1250 of Frieda5 coal seam. An investigation program was carried out to determine the causes of this accident and to understand the induced phenomenon. The program include: seismic investigations, geological and geotechnical analysis and numerical modeling. The objective is to determine eventually the presence of particular geological structures and there role in the initiation of the accident. After the accident, the HBL intended to mine the Seam 6. In order to monitor the behavior of the ground during the mining of the Seam 6, a geotechnical program comprising a complete study of the natural conditions and a seismic analysis was carried out. The environment of the panel is marked by the presence certain number of faults. Stress disturbed by the exploitation of the old panels and the presence of faults are likely to modify the equilibrium of strata before the exploitation and can induced seismicity. To study their influence, the INERIS carried out a numerical modeling in the scale of mine area (SUIT3D and FAULT3D), for the determination of the stress zone induced by old and new panels.

INTRODUCTION AND OBJECTIVE

On June 21 2001, a fatal rock burst affected the main gate of the Frieda 5-1250 panel in the Merlebach Mine (HBL). The rock burst was extremely violent (magnitude 3.6). The heave exceeded 2 m (figures 1 and 2, HBL note number 384/01-JPA/ND [START_REF]CREEP » -Frieda 5 bottom track[END_REF]). The damaged zone was 250 m long and was located at between 100 m and 350 m in front of the face (between the sides). Based on prior knowledge, the accident zone was not considered to be subject to the risk of dynamic phenomena. Therefore, the mechanism causing this phenomenon had to be known so that a methodology for monitoring future panels could be defined. This methodology will assure safety in the mine with regard to the risk of dynamic phenomena. A search program was started up for this purpose, including the following aspects: a detailed geotechnical exploration: exploration boreholes, geomechanical tests and stress measurements;

a series of numerical simulations at the scale of the entire deposit (using the SUIT3D software) and a model on a local scale (using the UDEC code [START_REF] Itasca | UDEC 3.01 User's Guide and Command Reference[END_REF]); a seismic and seismo-acoustic retro-analysis were carried out to understand the cause of the advance and the seismic activity that preceded and followed the event.

The results obtained were submitted to a college of international experts. Their opinions were used as a basis firstly to determine the causes of the accident in the Frieda 5 seam, and secondly on Seam 6 operations, then on the Dora seam.

location of recorded seismic events (1) 21 h 38 ; magnitude 2.3 (2) 21 h 55 ; magnitude 3.6 zone affected by the event

2.1

Fig. 2: View of the damaged zone and location of seismic events.

ANALYSIS OF THE FRIEDA 5 ROCK BURST

SEISMIC ANALYSIS OF THE FRIEDA 5 ACCIDENT

A detailed analysis of the seismic activity during this day was carried out to determine the location in space of the seismic focuses and the mechanism involved at the source.

LOCATION OF THE SEISMIC EVENT

Seismic monitoring of mining operations at HBL is done using a listening network composed of about twelve recording stations on the surface and at the bottom of the mine [START_REF] Wahed | Etude de la microsismicité induite à partir des sismogrammes tridirectionnels[END_REF][START_REF] Senfaute | La surveillance micro sismique de l'exploitation souterraine du charbon aux Houillères de Provence[END_REF]. One of the important phases of the seismic analysis is to estimate the position of event focuses, determined from the source directions of the P waves (the first to arrive) recorded by the sensors.

The seismic focus of the 21/06/01 event is located on the Frieda5 main gate 100 m in front of the face (figure 1). The location in depth shows that the depth of the focus of this event is about 1400 m (estimated depth below the surface, figure 2) at 200 m from the Frieda5 cutting face.

MECHANISM AT THE FOCUS

The rupture mechanism was identified starting from an interpretation of the direction of the first movement of the seismic wave from the event. The technique consists of positioning the direction of the first movement of the wave recorded by each sensor on a Schmidt geodetic reference network. This type of analysis determines the mechanism of the source that induced the event. Implosive mechanisms (total expansion) associated with rupture of the rock mass (for example block caving) are very frequently observed in coal mines. When fault type geological structures are involved, the resulting mechanism is shear (expansion + compression). It is possible that the mechanism is combined and therefore difficult to interpret, particularly when the seismic listening network is not sufficiently dense in terms of the number of recording stations.

In the case of the June 21 2001 event, the analysis of the direction of the first movement demonstrated an explosive rupture mechanism. More explicitly, the movement of the wave was a pure compression. Events recorded before the accident, and particularly the 21/06/2001 event at 19h 38 (~ 20 min before the accident), show a shear mechanism (figure 2).

A study was carried out on the history of the seismic activity at cutting faces located in the Frieda5 sector. This retro-analysis identified four events with a magnitude greater than 2.0 located exactly in the area of the accident. These events induced by operation of the Irma-1250 face in January 1999 were analyzed, considering the direction of the first movement of the wave. The identified rupture mechanism is explosive, and the same as the 21/06/01 mechanism associated with Frieda5 (figure 3). This observation shows that the anomaly in the accident area did exist, but was less active before Frieda5 was put into operation.

ROCKBURST LOCAL (STATIC LOADINGS) INDUCED (DYNAMIC LOADINGS)

Fig. 3: The identified rupture mechanism

CONCLUSION

The conclusions of the seismic analysis of the 21/06/01 event that occurred during operation of the Frieda5 face are as follows:

1. the local magnitude of the recorded seismic event was 3.6, and it was located on the main gate of the Frieda5 face at a depth of about 1400 m;

2. the rupture mechanism involved is of the explosive type (total compression), this mechanism induced by a high horizontal stress agrees with field observations (heave);

3. the explosive rupture mechanism was observed in the same zone in 1999 during operation of the Irma-1250 cutting face;

4. the Frieda5 event is probably related to a local geological anomaly at about 25 m from the Frieda5 face (solid and very strong sandstone bed)

SUMMARY OF THE GEOTECHNICAL EXPLORATION

An analysis of the boreholes confirms that there is a relatively solid sandstone bed at the level face, starting from a depth of 23 to 24 m. Its thickness varies between 15 and 20 m. A very thin layer of schist passes through the middle.

An analysis of the fracture and particularly the RQD^ shows that the sandstone bed is solid and that the value of the RQD is close to 100%.

The geomechanical properties show that the coal analyzed at the Frieda 5 face is fairly strong. The sandstone is highly anisotropic. The Young's modulus of the sandstone in particularly is not as high as Young's modulus for most sandstones tested in the HBL field [START_REF] Mery | Synthesis of the geomechanical characterizations[END_REF].

Stress measurements show that horizontal stresses are very high and the vertical stress is less than the weight of the ground (figure 4).

ANALYSIS BASED ON NUMERICAL SIMULATIONS

In this complex configuration, the numerical model provided some elements of a response attempting to reproduce the physical mechanism responsible for this accident. Two model types were made, one at large scale and the other at local scale (the main gate).

LARGE SCALE MODEL

The model using the SUIT3D program ([8], figure 5) was intended to achieve two objectives, firstly to find the initial tensor before mining started, and variations in the stress tensor close to the area of the accident [START_REF] Heib | Les nouvelles méthodes de modélisations numériques et le volume d'influence des exploitations minières en conditions complexes[END_REF]. The calculation was made assuming linear elasticity, with an isotropic stress tensor over the scale of the deposit. The seams considered in the calculation are shown in table 1.

Seam

MODEL AT THE SCALE OF THE MAIN GATE

Two models were made. One was for a vertical cut perpendicular to the center line of the gate. It is loaded by the stress condition measured at 800 elevation. The other model is a transverse model. The first model did not produce any results that could explain the origin of the rockburst.

The second model corresponds to a north south section and only includes the level face. The considerable extent of the level face implied a prior mechanism consisting of a long beam in bending built at its ends by zones with vertical overstresses.

The geometry of the model is built up using boreholes results made by HBL, at 50 m from the zone affected by the rock burst. Loads applied to the model were determined using the SUIT3D code, we have a stress field demonstrating the situation before the accident for the entire Frieda 5-1250 panel (figure 3);

A parametric analysis was carried out to simulate the bench bending mechanism, to find cases and sub-cases dependent on eight geometric and mechanical factors [START_REF]Attempted retro-analysis on the creep that occurred on June 21 2001 at the bottom level of the Frieda 5-1250 panel[END_REF].

The results of this model determined operating and environmental conditions to reproduce the rockburst. These conditions are illustrated in figure 4 and are summarized below:

1) a bottom case with a variable thickness sandstone bed (between 0 m and 20 m) and a 125 m wide zone in which vertical stresses were relieved;

2) a bottom case mode with N-S horizontal stresses increased by 25 %;

3) same mode as above (case 2) but with the Frieda 5 face advanced by 400 m, which is close to the accident configuration.

The mechanism studied would appear to be capable of reproducing total deconsolidation of the bed under some conditions (figure 6):

• the vertical stress relief in the accident zone must be less than or equal to 4 MPa;

« the North-South horizontal stresses must be high (higher than 60 MPa and therefore higher than the measurements made in situ);

• East-West horizontal stresses, and particularly shear stresses, are design parameters that are less sensitive than the previous two;

• the compression strength (Re) of the sandstone bed had to be halved so that the second sandstone bench could be fractured.

This fourth condition would imply heterogeneous mechanical properties of the sandstone bed (particularly the compression strength).

Furthermore, the results of this study show that as the Frieda5 face advances, the risk of the bed fracturing in the accident zone increases.

Furthermore, the length of this stress reduction zone also has an influence on the results; if the length of the pressure relief zone at 4 MPa decreases from 125 m to 80 m, the number of plastic points reduces by 30 %. The total length of the panel is 400 meters and the width of the cutting face is 260 meters.

GEOLOGY AND ENVIRONMENT OF THE PANEL

Seam 6 is located in the Cocheren North field, which is a series of Lower Coal Group. Boreholes made show that the roof and the walls are composed of beds of sandy schist and sandstone beds. These sandstone beds are usually associated with fine carbonaceous inclusions. According to boreholes observations, there is a sandstone bed, at least 15 m thick at the Seam 8 wall. Note that there is no disking in any of the boreholes.

The RQD values suggest that the roof is not as sound and more fractured than the face of Seam 6. Above the seam, note the presence of Seam 5 that is 10 m above seam 6, and the presence of seam 8 above the face at depth about 12 m. The face of seam 6 is different and almost homogeneous along the main and tail gates. Unlike the Frieda5 seam, for which the face is composed of alternating coal and schist seams overlying a 20 m thick sandstone bed.

Stress measurements confirm stress reduction in Seams 1 and 2, horizontal stresses are strongly anisotropic and their ratio is greater than 3.

The panel is surrounded by four faults including the St. Nicolas fault which is by far the most important fault, but is furthest from the panel (figure 5).

LARGE SCALE NUMERICAL MODEL

We made a numerical model at the scale of the Cocheren North field to determine stresses in the operations zone and the influence of previous operations using the SUIT3D code on the tracks and also to study the influence of operations on the faults that surround the future panel using the FAULT3D program.

The advance of the panel between 200 m and 470 m (maximum length) is simulated in 12 steps (figure 5). The model includes all mined panels of the Reumaux Mines (HBL).

The load on the full scale model was derived from:

• a vertical stress equal to the weight of the ground, namely (1250 m x 0.025 MN/m 3 = 31 MPa) ;

• an east-west horizontal stress equal to 0.43 times the vertical stress (equal to 13.5 MPa) ;

• a north-south horizontal stress equal to 1.3 times the vertical stress (namely 40 MPa).

The results were analyzed considering the variation of the state of stress with the advance of panel 1140-1250 (figure 5). The comments can be made on this model:

• operation of the Seam 6 panel will have an influence ahead of the faces of the Seam 1 and 2 panels starting from an advance of 325 m. This is a very small modification compared with the initial state before operation of the panel;

• starting from 400 m, the influence becomes more significant. It corresponds to an increase in the normal stress (zones close to the roof track increase by 2 MPa, the 28 MPa contours move by 110 m ahead of the face on the bottom level) and an increase in the east-west horizontal stress at the face of the 1st panel in Seam 1 (increase of 2 MPa);

• in the ventilation zone, the greatest increase concerns the east-west horizontal stress, which is 15% of the initial value. The normal stress only increases by 5%. No modification is observed for the north-south horizontal stress;

• the front ends of Seams 1, 2 and 6 do not intersect until the advance reaches 450 m (the planned length).

Monitoring measurements (deformation measurements) were reinforced starting from a 400 m advance.

HISTORY OF SEISMIC ACTIVITY IN THE SECTOR OF THE CUT IN SEAM 6

A study of the history of seismic activity in the sector of the Cocheren North field was carried out as part of the prevention plan for the cut on Seam 6-1250. In particular, we considered the history of seismic activity related to the operation of the cuts closest to Seam 6. These are Seam 1 (panels 1140 and 1250) and Seam 2 (panel 1140). Seismic activity related to all mining operations carried out between 1995-2001 is represented by the spatial distribution of focuses of events recorded in the HBL regional listening network. Based on the distribution of focuses in the zone limited to Seams 1 (panels 1140, 1250) and 2 (panel 1250), it is quite obvious that seismic activity is much lower (209 events) than was observed in the West sector (Irma, Erna and Frieda, 1500 events). A distinction is made between three activity distributions (figure 6):

MINING PERIOD AND SEISMIC ACTIVITY

1. minor rock bursts, most with a magnitude of less than 1.5, related to ruptures close to the St Nicolas fault;

2. a cluster of focuses located at about a hundred meters to the East of the panels in the preferred NS direction. A single rock burst with a magnitude of more than 2 is identified;

3. throughout the entire sector, only two rock bursts with a magnitude equal to 2.7 have been recorded and were located in the Seam 1-1250 panel.

To determine the mechanism, we considered all events with a magnitude of more than 2.0 recorded during mining panels in the Lower Gas Coal series.

In general, the mechanism identified in this sector is a shear failure of the fault type structure.

The results of the historic study on the seismic activity of Seams 1 and 2 (panels 1140, 1250) may be summarized as follows:

1. the seismic activity related to mining operations on Seams 1 and 2 is particularly weak compared with the activity in the sector of the Frieda5 cut;

2. only 7 of the 209 recorded events have a magnitude of more than 2.0.

3. identified mechanisms at the focus associate these events with shear type failures.

SEISMIC ACTIVITY IN SEAM 6

Mining operations at the front on Seam 6 took place between October and March 2002. Seismic monitoring was done using regional and local networks (seismoacoustic). The seismoacoustic network is composed of high frequency jugs (14 Hz), and monitors operations on the scale of the front.

The seismic activity related to mining operations of Seam 6 was very weak ; The regional monitoring network did not record any rock bursts that could be assigned to the front. The number of events recorded in the seismoacoustic activity is greater, but the events are low energy.

Figure 7 shows curves of the energy dissipated during operation of the Irma-1250, Frieda5-1250 and Seam 6 fronts as a function of the advance. For the three fronts, the energy is calculated starting from the seismoacoustic data. Irma-1250 was exploited over the longest length (1300 m), and advance of the Frieda5 and Seam 6 fronts was 300 m and 444 m. respectively. Thus, a comparison can be made between the three energy variation curves for an advance of up to 444 meters. An analysis of the three curves shown in figure 6 shows that the energy level released by the Frieda5 front is the highest and that the relief speed increases quickly after 180 m advance. For the Irma-1250 front, the dissipated energy is low until 420 m and then increases gradually. Note that the energy released by the front in Seam 6 is the lowest throughout the entire mining operation period.

CONCLUSIONS

After carrying out the analysis of the accident in the Frieda 5 seam (1140-1250), we introduced a new monitoring method in the Seam 6 panel (1140-1250). It was observed that seismicity was very low, and that the precautions taken did not cause any nuisance.

The same approach was used for the Dora seam.

The conclusion in this case was also favorable, and a monitoring and prevention plan was defined.

Experience with the previous panel suggests that the measures that we impose on ourselves will not have any significant repercussions on operation of the mine.

Fig. 1 :

 1 Fig. 1 : Diagram showing bottom level blow off.

Fig. 4 :

 4 Fig. 4: Stress diagram in accident zone

Fig. 5 :

 5 Fig. 5: Modeled seams with SUIT3D

Fig. 6 : 6 A

 66 Fig. 6: Overview: distribution of plasticity

Fig. 5 :

 5 Fig. 5: Different steps in modeling the advance on the Seam 6 panel.

Seams 1 and 2

 2 are located in the Cocheren North field, which is part of the Lower series in Merlebach mine. The Seam 6-1250 front is 120 and 100 m above Seams 1 and 2 respectively. Mining operations in this sector began on seam 1-1140 in 1995 and terminated at Seam 1-1250 in February 2001. Eight other cuts were operated in the Upper Coal series during this same period.

Fig. 6 :

 6 Fig. 6: Seismic activity of Seams 1 sector (1140, 1250 panels), Seam 2 (1140 panel), 1995-2001 period

Fig. 7 :

 7 Fig. 7: Curves of energy dissipated during mining operations of the Frieda5, Seam6 and Irma-1250 cuts

Table 1 :

 1 Characteristics of modeled seamsBased on these assumptions, the results obtained were as follows: mining operations in the Irma-1250 panel considerably modified the distribution of normal, horizontal and axial stresses. The influence of Erna3 and Frieda5 operations is lower and never exceeds 20%; overstress areas correspond to the Louise TB zone and the zone ahead of the Irma-1250 face; in the intermediate zone (where the part of the Frieda5-1250 panel on which mining operations are taking place is located, and the part of the bottom level affected during the accident); the vertical stress in these zones is relieved; furthermore, the influence of mining operations in the Frieda5-1250 panel is limited to zones close to the part of the panel from which coal has been extracted, and in particular it appears weak in the part of the level affected by the accident.

	north
	-70.00: ERNA3P
	0.00 :F5P
	50.00: BEORSETTE
	100.00: IRMAP
	324.00 : LOUISE

Ph. D., Eng. -I.N.E.R.I.S. (France)

Eng. -H.B.L. Coal Mine (France)

RQD : Rock Quality Design corresponding to the sum of the cores over a length exceeding 10 cm compared with the total length of the borehole


SALVIOlivier Hourtoloudavid
email: david.hourtolou@ineris.fr

ARAMIS PROJECT Accidental Risk Assessment Methodology for Industries in the framework of SEVESO II directive

This paper presents the ARAMIS project accepted for funding in the 5 th Framework Programme of the European Commission, which started on January 2002. The ARAMIS project aims at developing a new risk assessment methodology which allows to evaluate the risk level of an industrial plant by taking into account prevention measures against major accidents and the vulnerability of its environment. The methodology will support the harmonised implementation of the SEVESO II Directive. The project is built to result in the composition of an integrated risk index based on the definition of Reference Scenarios and combining the evaluation of : 1. Scenario consequence severity 2. Safety management effectiveness affecting the probability of occurrence of major accidents 3. Environment vulnerability. The methodology will be validated with case studies. Efforts are given to disseminate the methodology to decision-makers in charge of the control of major accident hazards. Thus the project development will be continuously monitored by a review team gathering risk experts from industry and EU competent authorities in order to ensure the widest acceptance of the approach.

Introduction

The ARAMIS project was submitted for funding in the 5 1 Framework Programme of the European Commission in February 2001 under the programme "Environment and Sustainable Development", in the chapter untitled "The fight against major natural and technological hazards" of the Work Programme. This 3-years project started in January 2002.

The ARAMIS methodology builds further on methods studied in the 4 th Framework Programme such as in the ASSURANCE a project, a benchmark exercise on the uncertainties in risk analysis, and developed in the I-RISK project, which provides a methodology for indepth judgement of safety management requirements for the design, operation and maintenance of major hazards plants.

The development of ARAMIS is justified by the need of the elaboration of a methodology giving consistent rules for the identification of scenarios that take into account mitigation devices and some aspects of safety management, and being recognised by risk experts from Competent Authorities and Industry. Beside, there is an need to establish a method that is capable to assess the risk level of an installation by integrating the preventive measures implemented by the operators.

a Assessment of the Uncertainties in Risk ANalysis of Chemical Establishments, n°ENV4970627 b Development of an integrated technical and management risk control and monitoring methodoloj managing and quantifying on-site and off-site risks, n°ENV4960243.

Such a method is a prerequisite in order to reach the goals of the SEVESO II Directive, that are to improve the prevention linked in particular with the safety management. So, the ARAMIS methodology propose to characterise the risk level with an integrated risk index composed with independent parameters related to the consequence severity evaluation of scenarios, the prevention management effectiveness and the environment vulnerability estimation describing the sensitivity of the potential targets located in the vicinity of the SEVESO II establishments. The application of this method will result in a more consistent and harmonised risk evaluation and safety management strategy in all European Countries.

The paper starts with the presentation of the context of the major-accident hazards control and prevention in the EU. Then, the objectives of the project and the work contents are described in details. Finally, the expected impacts of such a methodology are addressed.

Context of major-accident hazards control and prevention

The 1999 annual report from the European Environment Agency [START_REF]Environment in the European Union at the turn of the century[END_REF] indicates that the trend in accidents shows that many of the often seemingly simple 'lessons learned' from accidents

have not yet been sufficiently implemented in industry's standards. There is no doubt that disasters will continue to occur throughout the EU. Some of these will be due to technology, some to the forces of nature. Inevitably there will be loss of life and environmental damage. However, hazards can be managed to reduce risks. The problem of low-probability, highconsequence events is likely to remain a key issue in terms of risk management.

The most significant EU Directive to help protect people and the environment from major accident hazards is the SEVESO II Directive. This Directive applies to those industries that use significant amounts of hazardous substances. Their operators must demonstrate that they apply a policy for the prevention of major accidents using appropriate measures related to both "hardware" and "software" aspects, such as safety management systems. This is likely to reduce risk levels, not only from high-probability, low-consequence accidents, but also from low-probability, high-consequence events, although these are by nature difficult to address.

In the SEVESO II Directive, the objectives in terms of risk management are very clear, but the remaining question is : How to reach them ? For example, there is no harmonised definition of the scenarios that have to be considered for risk assessment. Typically, the chosen scenarios (BLEVE, total loss of containment, fire in the largest tank, explosion of the largest mass of explosive, etc.) can be different according to the specific risk analysts and according to the deterministic or risk-based approach of the country applying the Directive. This situation is confirmed by the results of the EC project ASSURANCE, in which 6 European organisations perform a benchmark exercise for the risk analysis of a specific plant. The partners use various hazard analysis techniques and arrive at quite different conclusions with respect to the scenarios that are relevant for the safety assessment. Moreover, sometimes, according to reference [START_REF] Cassidy | Special issue : The SEVESO II Directive (96/82/EC) on the control of major accident hazards involving dangerous substances[END_REF], land-use planning constraints urge the operators to consider reduction of the safety distances. Then, it may be proposed to choose 'realistic' scenarios by taking account of the effectiveness of mitigation devices. In fact, because of the lack of rules for identifying scenarios including safety measure effectiveness, the expert's job is tricky and often involves large subjective elements.

Not only risk assessment experts, but also decision-makers are confronted with a variety of approaches to assess and manage industrial risk. The difference of cultures in the Member States results in a multiplicity of methods for the evaluation of major accident hazards [START_REF] Kirchsteiger | Special Issue on International Trends in Major Accidents and Activities by the European Commission towards Accident Prevention[END_REF]. This fact makes the comparison of risk studies performed by different analysts a difficult task and has significantly hampered the widespread use of risk assessment for decision making purposes. At the recent EC-JRC International Workshop on Promotion of Technical Harmonisation on Risk-Based Decision Making, held in Italy in May 2000 [START_REF] Kirchsteiger | Promotion of technical harmonization on risk-based decision-making[END_REF], most participants agreed that comparative risk assessment along harmonised procedures would significantly help the decision understanding. A harmonised risk assessment methodology would thus ensure that risk-based decision making provides the necessary transparency and strikes the right balance between scientific understanding and precaution.

To propose a harmonised methodology for risk assessment is difficult. However, some aspects of the different approaches can be put in common such as scenario identification, severity evaluation and the integration of the effectiveness of the safety management that affects the major accident probability of occurrence. Because of these reasons there is a real need to establish common rules to identify scenarios integrating the prevention management achieved by the operator and to propose a harmonised method for their evaluation [START_REF] Salvi | Why changing the way to measure the risk ?[END_REF].

Objectives

The objective of the ARAMIS project is to create a new integrated risk assessment methodology by combining the strong points from the different methods currently used in risk assessment in European Countries. The methodology will be used as a supportive tool to promote safety in the process industry.

In particular, it will contribute to speed-up the harmonised implementation of the Council Directive 96/82/EC of 9 December 1996 on the control of major-accident hazards involving dangerous substances known as SEVESO II Directive. Accordingly, this tool should be flexible enough to take into account the different national cultures in industrial risk assessment like deterministic or risk-based approaches, so that the new methodology could become a recommended tool used by risk experts and endorsed by the risk decision-makers in the whole EU.

In technical terms, ARAMIS project aims at :

1. Establishing a methodology for accident scenarios identification taking into account the prevention process carried out by the operators of SEVESO establishments, and endorsed by the Competent Authorities and the decision-makers in charge of risk control,

Composing an integrated risk index which takes into account:

• Consequence severity evaluation of scenarios,

• Prevention management effectiveness (preventative, protective and mitigation measures) on initiating and aggravating events, thereby reducing the probability of occurrence of major accidents, • Environment vulnerability estimation.

4.

Project work description

Introduction

The objective of ARAMIS is to develop a risk assessment methodology to evaluate the risk level of establishments by taking into account the prevention measures implemented by the operators and the vulnerability of its environment. The project work plan is built to result in the characterisation of the risk level which is based on the determination of Reference Accident Scenarios and integrates :

• Effects severity evaluation of scenarios,

• Prevention management effectiveness,

• Environment vulnerability estimation. The end-users of the methodology are both the industrial companies and the Competent Authorities in charge of the application of the SEVESO II Directive. Thus, the valorisation and dissemination plan start at the beginning of the project with large exchanges with the endusers partners in the consortium and in a Review Team.

Project description

This paragraph describes the three main phases of the project which are : 1. Development of the methodology ; 2. Finalising and testing the methodology ; 3. Valorisation and dissemination plan.

Development of the methodology

The development of the methodology starts with the identification of reference scenarios, that are evaluated. Then the prevention management effectiveness and the environment vulnerability of the establishment are characterised. All these results are integrated to assess the risk level of a given establishment. The various phases of the methodology are described in details hereunder.

• Scenarios identification The objective of this phase is to propose a methodology for the identification of Accident Scenarios. For industrial installations, the Major Accident Hazards will be first identified with an algorithm based on the labelling of the substances (Directive 67/548/EEC) and the conditions of their use (pressure, temperature, flow, etc.). Then, the Reference Accident Scenarios will be determined from the Major Accident Hazards and from the review of accidents which occurred on similar units. The Reference Accident Scenarios will take into account the current practices (state of the art) mentioned in the legal requirements with regard to design, operation and control, and mitigation. Therefore, the Reference Accident Scenarios will use results from the work on the prevention management effectiveness as described in Figure 1. Reference Accident Scenarios define realistic scenarios, considering an installation operated today. They will be used to evaluate the effects (severity) of the major accident and describe the hazard potential. • Evaluation of effects severity of scenarios The objective of this task is to define a severity index S depending only on physical parameters and characterising the possible effects of scenarios. It is intended to study the physical characteristics of the phenomena involved in accidents (dispersion, explosion, fire), and to take them into account to evaluate the severity of the scenarios. The parameters to be considered are :

SCENARIOS IDENTIFICATION

Industrial

-the effect area A concerned with the phenomenon : for instance, a disc in case of an explosion, the projection of a plume for gas dispersion ; -the phenomenon kinetics K : rapid for explosions, slower for dispersion and fires ; -the capability of intervention / to mitigate the disaster : possible for fire and gas dispersion, but possible only at the design step for explosion ; -the potential of domino effects D : fragment emission, interlocking of delayed phenomena. A severity index S is therefore a function of parameters only associated with physical phenomena. All scenarios identified can then be evaluated and ranked with this severity index according to the calculation of S o for the Major Accident Hazards and S re f for the Reference Accident Scenarios.

• Prevention management effectiveness

The objective of this task is to define an index M characterising the prevention management effectiveness. Because technical and organisational factors are key issues to prevent major accident, this task consists in developing a methodology to evaluate the management effectiveness. Safety management applied in a Major Accident Prevention Policy leads to define actions to manage technical, human and organisational factors.

The operational goal of safety management is to strengthen the barriers and lines of defence against accidents (safety equipment or human operation). Safety management contains a large number of responsibilities, tasks and functions that are difficult to disentangle. A way of discriminating different levels in safety management is as follows :

-Policy : The implicit or explicit statement of a company's intentions with respect to plant safety, the objectives and goals for safety management and the way safety is prioritised and incorporated in the company's daily management. -Organisation : Organisation of safety management requires allocation of resources, definition of tasks, and scheduling activities. -Operation and maintenance : An important part of safety management is maintaining the reliability of the safety-critical technical, human and organisational components. This activity/responsibility includes :

• training, education and competence of personnel,

• maintenance of technical systems and introduction of new safety devices,

• maintenance of procedures,

• keeping up hazard awareness, e.g. by updating risk assessments. -Leadership : Implementation of safety management requires leadership, showing consistency between stated policies, intentions and objectives and decision-making in daily plant management, setting examples, creating common values and attitudes.

Leadership has important impact on safety culture, safety awareness and prevention of "unsafe acts".

The evaluation methodology will be built on the use of several research approaches : -Analysis of the effectiveness of safety devices providing physical safety barriers and lines of defence according to their characteristics (nature, availability, reliability, maintainability, testability...). This analysis follow the principle of the norms IEC61508 and draft IEC61511 (Functional safety : safety instrumented systems for the process sector) and lead to general methods to improve safety barriers and some results will be used for the scenario identification. -Analysis and comparison of specific safety management systems (e.g. application of standards) and analysis of how safety policies are embedded in the company's overall management system. -Development and use of theoretical modelling of management tasks, with Structured Analysis and Design Techniques (SADT) or function oriented modelling. This will be built on the work carried out in earlier EU projects, like I-RISK which established different ways of linking technical risk analyses with organisational influences. -Expert judgement, in particular to prioritise the management factors for assessment purposes. -Identification of safety performance indicators using audit techniques, questionnaire techniques and analysis of incident reports. -Development and validation of audit techniques.

Safety management affects the probability of occurrence of the scenarios. Therefore the objective of this work will be :

• To assess the effectiveness of various forms and aspects of safety management in preventing accidents.

• To develop reliable indicators that are a good measure of the effectiveness of a plant safety management.

This information will be used to define a multidimensional index M characterising the prevention management effectiveness.

• Environment vulnerability estimation This phase aim at defining an index V characterising the spatial vulnerability of the environment of an hazardous establishment by characterising potential targets (population, natural and man-made environment) and to estimate their sensitivity. To reach this objective, the area of interest in the vicinity of a plant will be divided into meshes : the potential targets belonging to each class (population, natural and man-made environment) will be identified and localised with the support of Geographic Information Systems (GIS). The sensitivity of the targets (workers of the plant, residents, surface and underground waters, public buildings...) will be characterised and ranked, using a multicriteria ranking method (SAATY), determining a scale of vulnerability levels. Vulnerability maps will be obtained by calculating and combining the vulnerability of all the targets falling in the same mesh.

Finalising and testing the methodology • Characterisation of the Risk Level RL

The severity index S can be combined with the management effectiveness index M and the vulnerability index F to define a risk level RL index of an installation in its environment (See Figure 2). The objective of this phase is to study the relation between S, M and V to characterise the risk level. It will be studied whether the risk level should remain characterised by the 3 indexes or whether the 3 indexes could be aggregated to form a multidimensional index. The ARAMIS method enables ranking the hazards only in terms of severity with the calculation of the severity index S for each scenario. Then the scenarios identified in several units are comparable. It also enables taking into account the efforts (preventive measures) made by the company with the estimation of the prevention management effectiveness M. The result also makes it possible to compare the risk level between two or more units of an industrial group to define the priorities for the investment for safety.

• Case studies To validate and to improve the ARAMIS methodology, case studies will be carried out with the collaboration of Companies and Competent Authorities in several SEVES O establishments in Europe. For the selection of the test sites, it will be assured that countries with a consequence-based and probabilistic approaches are represented. After this exercise, the definition of the indexes will be modulated to improve applicability and validity of the procedure.

ARAMIS METHODOLOGY

RL^F(S r e f ;M;V)

Valorisation and dissemination

In the valorisation and dissemination plan, efforts will be made to transfer the methodology to risk assessors and decision-makers, who are the end-users of the methodology. Industrial end-users are represented in the consortium through an association of European industrial companies. It will help the consortium to relay information about the project and its progress, to find plants for case studies and to disseminate the methodology at the end of the project. For dissemination, a web site was built (http://aramis.jrc.it) aiming at promoting the project and disseminating the public results. An electronic newsletter will also be released by the project management on the web site, after the progress meetings. Moreover, a intermediate workshop is set to provide the end-users with some partial results of the project and to collect comments to improve the relevancy of the further work. And a final workshop is also planned at the end of the project to disseminate the main results to all relevant stakeholders. The two workshops will be open to third-parties not involved in the consortium and workshop proceedings will be issued and made available on the web-site. Besides, in connection with the workshops, to ensure the widest possible dissemination of the results during the project, the participants will publish papers in scientific international journals and conferences.

Consortium description and involvement

The consortium consists often organisations involved in the risk analysis of major accidents. They are presented in Table 1. INERIS, the co-ordinator of the project, has an international expertise in the field of major accident prevention. It works as technical support for the national Competent Authority in charge of the application of the SEVESO II Directive. INERIS will lead with the steering committee the aggregation of the works for the risk level index composition and validation. INERIS will also provide support for valorisation and dissemination and for the Parallel Review.

EC-JRC-IPSC and especially MAHB has a recognised international expertise in the field of major accident prevention. It has animated EU Working Groups dealing with the application of the SEVESO I and II Directives and is also experienced in the development and use of accident databases and GIS tools at European level. MAHB acts as leader of the activities related to valorisation and dissemination of the results as well as leader of the parallel Review Team.

FPMs-MRRC has a great experience in the application of the SEVESO II Directive, and already developed methodologies on the choice of accident scenarios to study domino effects. It acts as the leader of work on Scenario Identification. In addition, the MRRC also brings its experience about domino effects and accident consequences modelling in the Severity Evaluation.

UPC (through CERTEC) has a recognised expertise in the evaluation of the accident consequences for SEVESO plants (dispersion, explosion, fire modelling). UPC will develop research on the Severity Evaluation as task leader.

The "Pôle Cîndyniques" of ARMINES has built a methodology to formalise the development of accidents as a series of "particles of experience" which are collected and documented from the analysis of reports, debriefing sessions and interviews. Using this methodology, it will contribute to the scenario identification and the work related to the prevention management effectiveness.

The SITE department of ARMINES has a long experience in environmental system management characterisation. It will mainly focus on the prevention management effectiveness and environment vulnerability characterisation. The LGEI of ARMINES has competencies in using both multi-criteria ranking methods (SAATY) and GIS. It has developed a methodology based on these two aspects for studying risks in transportation of hazardous substances., and will contribute to provide a methodology to rank the vulnerability of targets (human, environmental, equipment) in the vicinity of plants to obtain a vulnerability cartography used to characterise the spatial vulnerability.

RISOE is experienced with drawing up and evaluating safety reports for hazardous installations. It has special experience with applying function-oriented modelling to analyse the effectiveness of the organisation of safety procedures and using questionnaire techniques for assessment of safety culture. RISOE will be leader of the work on the Prevention management effectiveness.

UROM is experienced in methodologies and software tools, including GIS systems, to carry out risk analysis and area risk studies. Its activities will be mainly devoted to the development of the methodology for characterising the potential targets and their vulnerability. It will prepare a software tool for determining the environmental vulnerability index basing on GIS information.

Due to its experience, on one hand in fire and explosion, and, on the other hand in safety management and risk assessment, CMI will carry out research respectively for work related to the Severity evaluation and the Management effectiveness by analysing the implementation of management standards and guidelines.

Bringing its expertise in safety management modelling and risk assessment, TUD will carry out a major effort in the ARAMIS research project in work on Prevention management effectiveness with expert judgement and audit tools developed at TUD.

IChemE-EPSC will participate in the dissemination of the results to the industrial companies which are members or associates of the EPSC. It is important to notice that a lot of EPSC members are end-users of the ARAMIS methodology. In the project, EPSC will circulate to the members information related to the project and its results, and care about the Review Team participation.

In addition to the consortium, a Review Team is indeed built up. It has an essential role for the dissemination of the results through decision-makers involved in the control of major accidents. The Review Team has a role in the management and will comment on the applicability and usefulness of the results achieved. The involvement in the methodology development of risk experts both from the Competent Authorities and industrial companies will ensure that the methodology will be known and recognised at a European level.

6.

Contribution to the implementation of the SEVESO Directive

The project supports the European Research Area concerning the improvement of the knowledge, encouragement of the Science-Industry dialogue and harmonisation in decisionmaking process related to hazardous establishments.

The ARAMIS method will indeed be proposed as a recommended and harmonised tool used by risk experts and recognised by the risk decision-makers in the EU. Harmonising industrial risk assessments in Europe would significantly contribute to the European Commission's overall efforts to establish harmonised policies following the SEVESO II Directive. Such a harmonised risk assessment procedure would be of significant interest for both Competent Authorities and Industry :

• It would constitute a risk evaluation and comparison tool for industrial sites, which integrates the strengths of probabilistic and deterministic approaches. • The procedure would enable definition of progress plans within the framework of a safety management system. • It would enable to moderate the selection of scenarios by taking into account realistic data and preventive measures. • It will enable the evaluation and consideration of plant-specific safety devices and safety management effectiveness, as required in the Safety Reports.

The partnership in the consortium and in the Review Team ensures that the ARAMIS project will contribute on a very practical level to the EC research objectives built to support the further development and consistent implementation of European policies.

7.

Figure 1 :

 1 Figure 1 : Scenarios identification procedure

Figure 2 :

 2 Figure 2 : ARAMIS methodology representation

Table 1 :

 1 Description of partner organisation

	Organisation name	Short name	Country
	1. Institut National de l'Environnement Industriel et des Risques	INERIS	FRANCE
	Accidental Risk Division
	2. European Commission -Joint Research Centre -Institute for the	EC-JRC-	ITALY
	Protection and Security of the Citizen -Major Accident Hazard Bureau	IPSC-MAHB
	3. Faculté Polytechnique de Mons	FPMs-MRRC	BELGIUM
	Major Risk Research Center
	4. Universitat Politecnica de Catalunya	UPC	SPAIN
	Centre for Studies on Technological Risk (CERTEC)
	5. Association pour la Recherche et le Développement des Méthodes	ARMINES	FRANCE
	et Processus Industriels -ARMINES
	6. RiS0 National Laboratory	RISOE	DENMARK
	System Analysis Department
	7. Universita di Roma	UROM	ITALY
	Dipartimento Ingegneria Chimica
	8. Central Mining Institute	CMI	POLAND
	Safety Management and Technical Hazards
	9. Delft University of Technology	TUD	THE
	Safety Science Group		NETHERLANDS
	10. Institution of Chemical Engineers European Process Safety Centre	IChemE-	UNITED
		EPSC	KINGDOM

Acknowledgement

The ideas presented in this publication have been elaborated in the frame of the EU project ARAMIS " Accidental Risk Assessment Methodology for Industries ", contract no EVG1-CT-2001-00036, co-ordinated by INERIS (F) and including EC -JRC -IPSC -MAHB (I), Faculté Polytechnique de Mons -Major Risk Research Centre (B), Universitat Politecnica de Catalunya -CERTEC (SP), ARMINES (F),. Ris0 National Laboratory (DK), Universita di Roma -Dipartimento Ingegneria Chimica (I), CMI -Safety Management and Technical Hazards (PL), Delft University of Technology -Safety Science Group (NL), European Process Safety Centre (UK), Ecole des Mines de Paris -Pôle Cindyniques (F), Ecole des Mines de Saint Etienne -SITE (F), Ecole des Mines d'Aies -LGEI (F). The programme is organised within the Energy, Environment and Sustainable Development Programme in the 5 th Framework Programme for Science Research and Technological Development" of the European Commission." 8.


The objective of this paper is to present, based on the example of vinyl Chloride , the french procedure to set acute toxicity thresholds in the context of controlling urban development or landuse planning. Vinylchloride is a good example because it has never been until now examined for its acute toxicity and it is a carcinogenic chemical for humans chronically exposed at low concentration.

First it must be noted that these acute toxicity thresholds are established in a regulatory context (European Seveso II Directive , 1996) A risk study uses accident scenarios combined with threshold effects in human beings exposed to chemicals agents in order to be able to calculate the extent of:

• zones of hazardous effects before and after reducing the risk at source,

• isolation zones for controlling urban development.

The risk and safety study leads to determine the distances of effects centred on the storage unit, namely:

• the "distance of lethal effects"

• the "distance of irreversible effects"

The definitions of the toxicity thresholds were laid down at a consultation meeting on 4 June 1998, between representatives of the government, INERIS and the chemical industry. (transparent) "Lethal effects" are those which cause death in most individuals.

"Irreversible effects" correspond to a persistent injury or functional damage as a direct result of exposure in an accident situation (single, short-duration exposure resulting in disability).

"Reversible effects'" are when there is a return to the state of health prior to the accident.

The "lethal effects threshold" is the maximum concentration of pollutant in the air for a given exposure time, below which no deaths are observed in most individuals.

The "irreversible effects threshold" is the maximum concentration of pollutant in the air for a given exposure time, below which no irreversible effects are observed in most individuals.

These thresholds were drawn up by a "groupe de consensus" following the "Methodology for determining lethal effects and irreversible effects thresholds" in the case of accidental release of a chemical substance into the atmosphere, which was adopted on 3 r May 2001 and may be consulted on the INERIS Internet site (www.ineris.fr).

The composition of the « groupe de consensus « includes persons coming form government, public organisations and industry 3. Then the qualitative analysis of the lethal data has permitted to select two studies taking into account quality criteria.

• Mastromatteo(1960)

• Prodan (1978) The quantitative analysis was carried out based on the studies selected .

The statistical model used is the probit model. The probit analysis enables the proportion of effects (here mortality) to be linked to the level of exposure, characterised by a concentration and a time period.

Statistics software (MCSim®) was used to obtain probit equations. The LC50 and LCoi.were calculated.

The value n in Haber's equation (C n .t = k) has also been calculated from the analysed and selected data.

For each animal species, the probit equation established and this value n are as follows:

(transparent) Concentration is expressed in ppm and time in minutes.

For determining the non-lethal effects threshold according to the various studies analysed, the expert group examined the studies and critical effects from various studies (transparent) 4. In the review of the results it was considered that the data available for humans and animals show that the toxic mechanisms of this substance are the same and the pulmonary kinetics route do not differ significantly. This enables the animal data to be extrapolated easily to human beings.

> Definition of the lethal effects thresholds in human beings

Analysis of the results on rat, mouse and guinea pig shows that the guinea pig is the most sensitive animal species, (transparent) Thus, examining these results for these species combined with the data available in the literature enables us to adopt the data relating to mice for determining lethal effects thresholds (better quality of data). These thresholds are based on the values of LC o i for exposure times of 1, 10, 20, 30 and 60 minutes.

The choice of data relating to mice is supported by the results of Mastromatteo's study (1960), which shows no mortality whatever the animal species (rat, mouse, guinea pig) at 100,000 ppm for 20 minutes.

Furthermore, Belej's study (1974) shows no toxicity in primates for exposures of five minutes to 200,000 ppm. There is therefore little or no difference between species, which justifies not applying a safety factor in extrapolating the data from animal to humans.

Finally, we should remember that the lethal effect induced by vinyl chloride is characterised by a narcotic effect depressing the central nervous system causing cerebral anoxia.

The values adopted are set out in the It should be underlined that these lethal effects thresholds are in the high flammability area of vinyl chloride between 4 and 22%, that is, 40,000 and 200,000 ppm excluding the first few minutes of exposure.

> Definition of the irreversible effects thresholds

• Non-Carcinogenic Effects Thresholds

The analysis of non-carcinogenic effects described in the literature emphasizes that the main toxic effect of vinyl chloride is localised in the central nervous system and is characterised by a narcotic effect. This narcotic effect induces depression of the central nervous system (by altering neurone cell communications) causing cerebral anoxia whose degree of intensity determines lethality. There is no means of assessing or quantifying the extent of these lesions causing either irreversible effects or lethal effects. Moreover, in the case of early treatment (oxygen therapy) at the time of vinyl chloride poisoning, the exposed individuals show no disorders in the following days, nor in the long term. The current state of knowledge is therefore insufficient to be able to set non-carcinogenic irreversible effects thresholds.

• Carcinogenic Effects Thresholds

Vinyl chloride is classified as a carcinogen in the case of chronic exposure by different organisms (category 1 of the European Union and the I ARC).

The toxicological experts of the consensus group have therefore considered the question of adopting the carcinogenic effect as an irreversible critical effect following a single exposure of short duration.

The experimental data relating to prove a carcinogenic effect under these exposure conditions are quite scarce. Only the works of Hehir (1981) and Tatraï (1981) have been identified in the literature.

The toxicological experts of the consensus group decided not to adopt these studies since the hepatic and pulmonary tumours described in the animals are isolated cases whose prevalence corresponds to the spontaneous rate for this kind of tumour in the species concerned. Moreover, they are specific to animals and have not been described in humans as a result of a single exposure to vinyl chloride. Only cases of hepatic angiosarcomas are known for chronic exposures in humans.

The group also pondered the relevance of extrapolating toxicological reference values (TRV) established for whole life exposures to short exposures.

Within the context of assessing health risks for a repeated exposure over a long period to a carcinogenic substance without a threshold (genotoxic) the concept of unit excess risk (ERU) is adopted. ERUs are estimated from epidemiological or experimental studies on animals.

For vinyl chloride, International organisations have published toxicological reference values for chronic exposures over a whole lifetime. In France, the circular of the Ministry of Ecology and Sustainable Development of 2 nd August 2001 relating to classified installations using vinyl chloride monomer adopts that of the WHO guideline (2000), that is, an ERU of 10" 6(lug/m 3 )" 1 for acceptance risk of 10" 5 . (transparent).

If a relationship is assumed where the product of the exposure concentration multiplied by the time of exposure is a constant, one can calculate a given acceptable Individual excess risk (for example 10' 5 ) which corresponds to a wholelife (613200 hours) exposure concentration of 10 p.g/m3 , the concentration corresponding to a one or eight hours exposure. The values are only given as an example, since knowledge of cancer-triggering mechanisms is insufficient to confirm this extrapolation over several orders of magnitude of exposure concentration. The existing data rather lead to reject this approach, since the genotoxic properties of vinyl chloride are linked to its main metabolite (chloroethylene oxide) and not to vinyl chloride itself. The genotoxic action of this metabolite is expressed in the formation of DNA adducts. This kind of genie lesion is characterised by a low probability of tumour development.

Furthermore, examination of the toxicokinetics of this substance also shows that the formation of this carcinogenic metabolite is a saturable phenomenon for concentrations between 3000 and 7000 ppm (Withey, 1976).

Taking these factors into account, the toxicological experts of the consensus group finally concluded that the present state of knowledge precluded setting threshold values for irreversible carcinogenic effects associated with a single exposure.

In conclusion :

This is the point of view of the French "group de consensus" which has been validated within INERIS organisation.

It must be emphasised that French Acute toxicity thresholds are set :

-in a regulatory context to make prevention for land use planning.

-taking into account two acute thresholds whose definitions are exact to select critical studies and critical effects. For example : for lethality effects, the group examine lethality studies and no other.

-Regarding the uncertainty factors with a less conservative approach than that of the AEGL Committee but a protective approach for the general population.

  IN THE CONTEXT OF CONTROLLING URBAN DEVELOPMENT OR LAND USE PLANNING Dr Annick Pichard, INERIS, Parc technologique ALATA, B.P2, 60550 Verneuil en Halatte-France AEGL Committee, september,2002

  The main steps of the demarch are the following (transparent) :-Research of available Official value Research and examination of the toxicity data in human beings -Research and examination of the toxicity data in animals Analysis of toxicity data Analysis of lethality data -Analysis of the non-lethal-effect -Review of the results and definition of the acute thresholds levels 1. Concerning the Official data TEELS (Temporary Exposure Emergency Limits ) are defined by the US Ministry of Transport and used when ERPGs are not available. They are intended for assessing the effects on a general population in the case of accidental exposure for a period of sixty minutes. They are defined without a safety factor and characterised (transparent) For vinyl chloride, these thresholds are: • TEEL-0: lOppm • TEEL-l:50ppm • TEEL-2: 50 ppm • TEEL-3: 75 ppm 2. A large review of the littérature has been made to find the toxicity data for humans and animals.

  The concentrations thus obtained are listed in the following table(transparent)

		Exposure time	Exposure time
		1 hour		8 hours
	ERI	mg/m 3	ppm	mg/m 3	ppm
	io-5	6,132	2,360	767	295


Tikou Belem

Françoise Homand

Mountaka Souley

New shear strength criteria for rock joints taking into account anisotropy of surface morphology

Three new peak shear stress criteria were proposed to predict the Variation of shear strength with normal stress for dilatant (irregulär and regulär surfaces) and non-dilatant to slightly dilatant (planar rough surfaces or regularly undulated surfaces ) rock joints under both constant normal stress (CNS) and constant normal stiffaess (CNK) loading. The planar rough surfaces exhibit an isotropic nature whereas the undulated surfaces are anisotropic. These modeis take into account the progressive degradation of surface roughness during the course of naonotonous or cyclic shearing. The third proposed model is a generalized new shear strength criterion which predict suecessfully the shear strength behavior of rock joints under CNS and CNK loading.

l. INTRODUCTION

The behavior of rock joints under either constant normal stress (CNS) or constant normal stifmess (CNK) loading conditions depends mainly on the wall surfaces roughness characteristics, the degree of matching and/or interlocking, the presence of filling materials and on the rock material properties. According to [l], the dominant factor influencing the mechanical behavior of rock joints is their morphology and roughness, and the most difficult thing is their characterization and modeling. The presence of discontinuities strongty affects the stability of rock engineering stmctures such äs tunnels, Underground mine excavations and open pits. The knowledge of the shear strength of rock joints is necessary when deterroming the stability of jointed rock masses. The Variation of peak shear stress with normal stress has been modeled by many authors [START_REF] Patton | Multiple modes of shear failurs in rode[END_REF][START_REF] Ladanyi | Simulation of the shear behaviour of a jointed rock mass[END_REF][START_REF] Jaegger | Friction of rocks and stability of rock slopes[END_REF][START_REF] Barton | Review of a new shear strength criterion for rock joint deformation[END_REF][START_REF] Swan | Prediction of shear behaviour of joints using profiles[END_REF][START_REF] Saeb | A variance on Ladanyi and Archambault's shear strength criterion[END_REF], and more recently by [8j. Although these modeis have substantially improved our knowledge on rock joints behavior, their limitations must be recognized [l]. In addition, few works have been devoted to take into account the progressive degradation of joint wall surface asperities during the course of monotonous or cyclic shearing [START_REF] Plesha | Constitutive modeis for rock discontinuities with dilatancy and surface degradation[END_REF][START_REF] Hutson | Joint asperity degradation during cyclic shear[END_REF][START_REF] Benjelloun | Experimentai and numerical investigation on rock joints[END_REF][START_REF] Jing | Study of rock joints under cyclic loading conditions[END_REF][START_REF] Homand-Etienne | Rock joints behavior under cyclic direct shear tests[END_REF].

The main purpose of the present study is to predict the peak shear stress of natural or artificial joints by taking into account the evolution of secondary and primary roughness through new proposed roughness Parameters based on 3D laser profilometry [START_REF] Belem | Quantitative parameters for rock joint surface roughness[END_REF][START_REF] Belem | Fractal analysis of shear joint roughness[END_REF]. In the concept of primary and secondary asperities, roughness is subdivided into two components: the secondary (or second-order) and the primary (or first-order) roughness [START_REF] Jing | Study of rock joints under cyclic loading conditions[END_REF][START_REF] Kana | Interlock/friction model for dynamic shear response in natural jointed rock[END_REF][START_REF] Belem | Quantitative parameters for rock joint surface roughness[END_REF]. The secondary asperities are defined by the surface heights distribution (which correspond to the sensu stricto roughness) while the primary asperities are defined by the overaIl geometry of the surface (which describe the surface anisotropy).

Based on joint initial morphological properties such äs 2 nd order asperities (DR r ), surface angularity (9 s ), surface anisotropy/isotropy (k n ), maximum asperity height (ß o ) and roaximum shear displacement («""), two new shear strength (r?) criteria were proposed. This approach takes into account the matching/non matching and the dilatant/non dilatant nature of sheared joints. The strength behavior predicted by the proposed new shear strength criteria is compared to the results of some laboratory direct shear tests and to the shear strength criteria proposed by Ladanyi and Archambault [START_REF] Ladanyi | Simulation of the shear behaviour of a jointed rock mass[END_REF], Saeb [START_REF] Saeb | A variance on Ladanyi and Archambault's shear strength criterion[END_REF] and Barton [START_REF] Barton | Review of a new shear strength criterion for rock joint deformation[END_REF].

NEW ROCK JOINT SURFACE ROUGHNESS PARAMETERS

In order to better characterize joint surface roughness which includes characteristics such äs magnitude, angularity, undularity and anisotropy, numerous parameters have been proposed by Belem et al. [START_REF] Belem | Quantitative parameters for rock joint surface roughness[END_REF] to Supplement and overcome the inefHciency or the subjectivity of some of the already existing linear parameters (e.g. Zi, R L , P s and JRC). Some of these new parameters are succinctiy given beiow, but more details should be consulted in Belem et al. [START_REF] Belem | Quantitative parameters for rock joint surface roughness[END_REF].

Profile mean angle, 9 p

The profile mean angle was defmed for describing its angularity. Por a given joint topographic profile, the arithmetic mean angie of profile inclinations, 9 p , is caiculafed along x-or y-axis by: -tan

N -\ -z, ( l
)
where k denotes x-or y-axis; z" the discrete values of profile heights; (Nk-1), the number of intervals used for slopes caiculation along Ä:-axis; Ak, the sampling Step along ife-axis (Fig. l). For the whole joint surface involving non-'identical profile lengths (e.g. circular sections), the pseudosurfacial mean angle, defined äs the weighted mean of the arithmetic mean profile angies in Ä-direction, [öp ) k=x y , is given äs follow:

= atan ( 2
)
where Mi is the total number of profiles in kdirection (;cory); Nj[ the number of discrete points corresponding to the j* profile along Ä-axis; L [ the length of the j* profile along Ä-axis. When all the surface profiles have the same nominal length (L{ = Z*; i.e. identical L{ in /k-direction), Eq. ( 2) is reduced to an arithmetical inean.

Three-dimensional mean angle, 9 s

By considering that the joint wall surface is made up of an assembly of elementary flat surfaces defined by topographical data (X, y, z), the spatial orientation of each elementary surface is characterized by the azimuth and the inclination angle, 9, of its normal unit vector (Fig. 2). N e,lementary actual area (A,).

-^~ -\ elementary nominal Ax i .''hy Fig. 2. The i"' elementary surface showing the angle 9,.

The three-dünensional mean angle is caiculated from the angies 6; of normal vectors of all the elementary mean planes äs follow:

1 m m ,=i (3)
where m is the total number of elementary surfaces. Angle Qs is considered äs the mean angle of surface asperities or surface angularity.

Degree of surface apparent anisotropy, k y

The pseudo-surfacial mean angle caiculated along both x -and y-directions,0 p , can well describe the apparent structural anisotropy of surfaces in the xy plane. The degree of surface apparent anisotropy, k u , is defined by the ratio of the anisotropy ellipse small half-axis b to its big half-axis a by the following relationships :

b mm a (0<* (4)
Subscript x or y denotes respectively the direction in which the individual profiles arithmetic mean angle, 6p, have been caiculated. When 9 < K < 9.25, surface is anisotropic (e.g. k a = 9 corresponds to surfaces with saw teeth, undulating surfaces, etc.); when 9.25 < K < 0.5, surface is more anisotropic than isotropic; when 9.5 < £ a < 9.75, surface is more isotropic than anisotropic; when 9.75 < k a < l, surface is considered isotropic (e.g. k i = l corresponds to a perfectiy isotropic surface).

Degree of surface roughness, DR r

Joint surface roughness coefficient, R s , has been defined for a single joint wall by El Soudani [START_REF] El Soudani | Profilometric analysis of fractures[END_REF] äs the ratio of the actual joint wall area, Ai, and the nominal surface area. A n (see Fig. 2) äs follow :

R.= A A, (5)
Actual surface area can be caiculated either by sarface triangulation or by integral method [START_REF] Belem | Quantitative parameters for rock joint surface roughness[END_REF]. Tbe upper lirrdt value of R s = 2 was suggested by El Soudani [START_REF] El Soudani | Profilometric analysis of fractures[END_REF] and is applicable to brittle fractures without recovering. The degree of joint surface roughness, DR r , was defined from Eq. ( 5) in order to relate the evolution of surface roughness to its initial state by the following relationship:

DR A A,-A R R. (0<DR r <l) (6)

NEW SHEAR STRENGTH CRITERIA

Most of the existing shear strength criteria in the literature are developed to predict the peak shear stress of üütially mated and interlocked joints displaying some dilatant behavior. If the majority of these modeis takes into account some aspects of joint initial roughness [START_REF] Ladanyi | Simulation of the shear behaviour of a jointed rock mass[END_REF][START_REF] Jaegger | Friction of rocks and stability of rock slopes[END_REF][START_REF] Barton | Review of a new shear strength criterion for rock joint deformation[END_REF][START_REF] Swan | Prediction of shear behaviour of joints using profiles[END_REF][START_REF] Saeb | A variance on Ladanyi and Archambault's shear strength criterion[END_REF], very few modeis take into account the effect of both surface anisotropy [START_REF] Kulatilake | New peak shear strength criteria for anisotropic rock joints[END_REF] and progressive degradation of joint surfaces during the course of shearing [START_REF] Homand-Etienne | Rock joints behavior under cyclic direct shear tests[END_REF]. In the present approach, authors suppose that there niainly exist two types of joint:

(i.) non-dilatant joints (isotropic surfaces) (ii.) dilatant joints (isotropic/anisotropic surfaces)

We assume that dilatant joint surfaces (regulär or irregulär joints) are often anisotropic or more or less anisotropic where the primary roughness component plays a preponderant role. We also assume that non-dilatant joint surfaces (generaüy irregulär) are more or less isotropic and where only the secondary roughness goverhs the shear strength behavior. In addition,. we «uppose that a dilatant joint will tend to be much less degraded than a nondilatant joint which must be much degraded during the course of shearing. Dilatant joints are usuaüy mated and/or interlocked while non-dilatant joints can be non-mated and non-interlocked (e.g. soilstructure, rock-structure, soil-rock, concrete-rock or mine backfill-rock interfaces).

Two shear strength criteria are proposed for the prediction of peak shear stress of non-dilatant joints exhibiting a strong potential of degradation and that of dilatant joints which exhibit a iow potential of degradation. These criteria were formulated in the more general case of cyclic shearing where the shear stress-displacement curve (U s , T) is broken up into four portions (Fig. 3). For each cycle of shearing the model will predict the peak shear stress (T P ) on the first portion (Portion 0) of the (,U s , t) curve äs shown on Fig. 3. Shear displacement (« S ) on the Portion 0 will be u°a nd the shear displacement for one cycle of shearing will be of (4M°). With n, the number of cycles of shearing, the maximum cumulated shear displacement on the Portion 0, K™" = u°(4n -3).

For a monotonous shearing, n = l and M"™' = H°. Another parameter to be taken into account in the formulation of the modeis is the maximum amplitude of surface heights, ßo (difference between maximum and minimum height of asperities), which is a constant.

Shear strength criterion for non-dilatant

joints It was found from previous works [START_REF] Belem | Morphology and mechecanical behavior of rock discontinuities[END_REF] that the roughness of planar unmated and non-interlocked joints decreases a tot with the increase of normal stress and number of cycles of shearing (no significant dilatancy was observed). In addition, almost all the asperities have been sheared off after 5 cycles of shearing. It was concluded that the surface asperities degrade during the course of shearing by a progressive wear äs well äs by a collapse and leading to an increase in contact area and interfacial friction. Let ß n be the strength component in the shear behavior of rough, planar, almost isotropic and non-dilatant joint. The peak shear stress (T P ) criterion for this type of joint which takes into account sliding through asperities and progressive degradation of asperities during the course of monotonous and cyclic shearing is given äs follow :

where (j> r is the residual friction angle, ß n (subscript n denoting non-dilatant) is an angle describing the effect of joint roughness change due to asperities degradation. From our experimental investigations, it was found that angle ßn is angle B s -dependent and is given äs follow:

A=' 9 °(l + w,) (8)
w =-
and in tenn of number of cycles of shearing :

a a ' DR[ -xlog (9a) ( 9b
)
where 9° is the initial surface angularity; (l + W g ) is the progressive degradation factor; (T u , the normal stress; O c , the compressive strength of sample material; ao, the maximum amplitude of surface; u°, the shear displacement on the fast portion (Portion ff) of shear curve (see Fig. 3); u™', the cumulated maximum shear displacement on the Portion 0 (for n cycles of shearing, n > 0); k a , the degree of surface apparent anisotropy; DR°, the initial degree of surface roughness.

Por unmated and non-interiocked rough planar joints, it was experimentally observed [START_REF] Belem | Quantitative parameters for rock joint surface roughness[END_REF] that shear stress continuously increases with shear displacement and this means that angle ß n will increases with Y™" or the number cf eycles n. Substituting Eqs. (8 & 9b) in Eq. ( 7), the new peak shear stress criterion for non-dilatant joints exhibiting a high potential of degradation is given äs follow:

r. =cr, tanU +0 a . 1 + [START_REF] Hutson | Joint asperity degradation during cyclic shear[END_REF] This peak shear stress criterion takes into account surface angularity (8 5 ),' surface anisotropy/isotropy (Ä: a ), strength of sample material (O c ), surface second-order roughness (DR°) and first-order roughness (ßo) and finally, maximum cumulated shear displacement, u'™" (or number of cycles of shearing).

Shear strength criterion for dilatant joints

It was also found from previous investigations [START_REF] Homand-Etienne | Rock joints behavior under cyclic direct shear tests[END_REF][START_REF] Lefevre | Comportement mecanique et morphologique des discontinuites en cisaillement[END_REF] that for a moderately dilatant joint, cyclic shearing involved a progressive degradation of undulations which in turn increases contact areas and therefore the shear strength. Consequently, an efficient model must be able to take into account, the absence or the presence of pure dilatancy (sliding through asperities), the progressive degradation of asperities with shearing and the friction of joint walls on debris. Among the numerous shear strength criteria for regulär or irregulär dilatant joints proposed by many authors, only those proposed by Ladanyi and Archambault [START_REF] Ladanyi | Simulation of the shear behaviour of a jointed rock mass[END_REF], Jaegger [START_REF] Jaegger | Friction of rocks and stability of rock slopes[END_REF], Barfon [START_REF] Barton | Review of a new shear strength criterion for rock joint deformation[END_REF] or Saeb [START_REF] Saeb | A variance on Ladanyi and Archambault's shear strength criterion[END_REF] are more suitable to predict the shear strength of both regulär and/or irregulär joints. This is because these modeis take into account the change of initial roughness due to shearing. The common point between these different shear strength criteria is the significant contribution -even essential -..of peak dilatancy angle (;' p ) in the formulation of these modeis. In this study we start from the Barton's criterion in which we replace the peak dilatancy angle (i' p ) by a new dilatancy-degradation angle which differs from / p- According to Barton [5], the peak shear stress is related to peak dilatancy angle and joint initial roughness by the following equation: r^ff.tanfô+ij [START_REF] Benjelloun | Experimentai and numerical investigation on rock joints[END_REF] where (j)b is the basic friction angle, and the peak dilatancy angle, i p , is given by the relationship:

JCS ( 12
)
where JRC is the joint roughness coefficient; JCS the joint wall compressive strength.

In the first stage of this study authors attempted to reformulate Eq. ( 12) using the new previously proposed roughness parameters. It is well known from the numerous experimental studies in the literature and from our laboratory investigations that the peak dilatancy angle decreases non-linearly with the increase of normal stress. When comparing surface initial mean angle values, 0°, to those of experimental peak dialatancy angle, f p , it was found that no trivial and simple relation does exist between these two angles. Experimental peak dilatancy angle is always greater than angle 9°w hich was practically constant for the majority of sampies. For a marked dilatant joint, the peak dilatancy angle was successfülly predicted using an exponential model given by:

f p = h ex p -' 'a. DR0 (13) (14)
where io represents the maximum inclination angle of surface prior to shearing (i.e. under 0 n = 0); AI is a sample constant depending on the initial roughness (DR r ) and the degree of surface apparent anisotropy, k a . From our investigations the maximum inclination angle, i y , was estimated to be at least twice the 3D surface angle (i' o = 2(9°). Then, Eq. ( 13) become: [START_REF] Belem | Fractal analysis of shear joint roughness[END_REF] Eq. ( 15) predicts well peak dilatancy angle for marked dilatant joints. But, substituting Eq. ( 15) m Barton's shear strength criterion (Eq. 11) leads to an underestimation of experimental peak shear stress. To better predict peak shear stress, a new angle must be defined in order to take into account both dilatancy phenomenon and degradation of surface asperities during the course of monotonous or cyclic shearing. This new dilatancy-degradation angle must be greater than the observed peak dilatancy angle Q' p ). Now, let i a be this new angle (subscript d denotes dilatancy-degradation). Consequently, angle u will differs from the experimentally observed peak dilatancy angle (i'd > ! p ). Angle id describes the effect of dilatancy äs well äs the surface change due to asperities degradation which increases joint waü surface contact areas and hence the shear strength. Based on our experimental observations, angle ;d can be formulated äs follows:

i d = ( 16
)
where AT. is a constant depending on initial roughness (DR r °), degree of surface apparent anisotropy (k a ), maximum amplitude of surface (a o ) and maximum shear displacement. K"'" (or number of cycles, n). It was found that constant Az can be expressed äs follows:

(17a) (17b) [START_REF] Belem | Morphology and mechecanical behavior of rock discontinuities[END_REF] and in term of number of cycles of shearing :

A, 1

a o { 4«-3 2(V 3 DR°a nd finally, i d = i.
3 DR°B y substituting Eq. ( 18) in Eq. ( 11), the shear strength criterion for dilatant joints which includes the influence of dilatancy, asperities degradation and the number of cycles of shearing is given äs foilow:

(19)

Generalized

shear strength criterion for dilatant joints AU the existing shear strength criteria are specifically formulated to predict the peak shear stress under constant normal shear stress (CNS) loading conditions. The main difference between CNS and CNK loading paths is that normal stinhess K a = 0 for CNS loading while K » > 0 for CNK loading. This is because O n > 0 and is constant for CNS loading, while da varies continuously during the course of shearing for CNK loading. Normal stiffness, K a , is related to normal stress, O n , and normal displacement or dilatancy (M n ) äs foilow: du.

Au " and (20a)

(20b)
where a^ is the initial normal stress.

Since we seek to predict the peak shear stress, then only the peak dilatancy (« np ) will be considered (Au n = An np ). Variation of peak dilatancy is related to dilatancy rate, tan(i' pcNK ), and incremental shear displacement. Au s , by the following relationship:

=Aw s xtan(i pCM: ) (21)
As only the fast portion (Portion 0) of the shear curve is considered (see Rg. 3) for the prediction of peak shear stress, AM S can reasonably be approximated by u°, the shear displacement on the fu-st portion (A« s = "°) -Hence, Eq. ( 21) become:

Au np =u°xtm(i pCNK )

Substituting Eq. ( 22) in Eq. (20b) leads to:

K, («° x tan(i p (22) (23)
where O np is the value of normal stress when the peak shear stress is reached during the course of CNK loading test.

From the proposed shear strength criterion for CNS loading path given by the relationship:

a generalized shear strength criterion can be derived for both CNS and CNK loading paths by substituting Eq. ( 23) in Eq. ( 24):

CHS.CNK (25a)
., = (a-. + K [u° x tan(; )])tan(A + ;, ) (

and finally where i p Ĩt was found from our experimental investigations that for a given Om, peak dilatancy angle observed for the CNS loading (; P _CNS) is always greater than the peak dilatancy angle for the CNK loading (;' P_CNK) due to the effect of the normal stiffhess K a . The CNK peak dilatancy angle can be predicted by the following relationship:

KX DR° (T",+( (27)
Eq. ( 27) is slight sünilar to Eq. ( 15) but different. In Eq. ( 26), the first term of the right hand side will correspond to the CNS loading condition when K n = 0. If K a > 0, Eq. ( 26) predicts the peak shear stress of CNK loading conditions.

MONOTONOUS AND CYCLIC SHEAR TESTS

Monotonous and cyclic shear tests were carried out on three types of joint: a man-made regularly undulated mortar joint (Fig. 4a), a non-interlocked and non-mated granite joint with hammered surfaces (Fig. 4b), and a natural rough and undulated schist joint mortar replica (Fig. 4c).

Direct shear tests program

The hammered joints underwent 5 cycles of shearing, the man-made regularly undulated joints underwent 10 cycles of shearing and finally, the natural schist joint replicas underwent monotonous shearing under CNS loading conditions with O n ranging from 0.3-6 MPa. The CNK tests were carried out on the regularly undulated joint and the schist joint replica (Figs. 4a &c) under K a ranging from 0.16-2 MPa/mm and um ranging from 0.4-2 MPa.

4.3.

Topography dato. acquisition In order to estimate the morphological parameters of test samples, surfaces topography was measured prior to and after each shear test using a laser sensor profilometer. The roughness parameters caiculated from these topographic data are listed in Table 2.

CNS test results

Tables 345summame the observed peak shear stress and peak dilatancy angle for the three types of test sample. The values are presented for the first cycle and the last cycle (5 11 1 cycle for the hammered joint and lO" 1 cycle for the schist joint replica). Eq. ( 19) 0.5 l 1.5 2 Normal stress a (MPa) 0.5 Fig. 7. Shear strength predicted by the new dilatant shear strength criterion when using dilatancy angle (i' p ) and the new proposed dilatancy-degradation angle (i'd) and compared to the experimental results for the natural schist joint mortar replica. for the undulated joint and the schist joint replica respectively. Exanunation of these figures shows that the proposed failure criterion and Barton's criterion are in good agreement with the experimental data.

CNK lest results

Tables 6 &7 present the predicted CNK peak shear stress by the new proposed generalized shear strength criterion (Eq. 26). A good agreement is observed for the marked dilatant schist joint replica (Table 6) and for the moderately dilatant joint with regularly undulated surface (Table 7).

CONCLUSIONS

Based on new surface roughness parameters previously defined by authors [START_REF] Belem | Quantitative parameters for rock joint surface roughness[END_REF], three new shear strength criteria wers proposed to predict nondilatant and dilatant shear strength behavior of rock joints. The first model (Eq. 10) was proposed for the prediction of the shear strength of unmated nondilatant joints (isotropic-surfaces). This shear strength model takes irito liccount the progressive degradation of surface asperities. The second model was proposed to predict the shear strength of mated dilatant joints (anisotropic surfaces) under CNS loading. The third model is a generalized shear strength criterion for both CNS and CNK loading. These three modeis successfülly predicted the shear strength of tested samples. Model parameters are the initial roughness and the shear displacement.

Fig. l .

 l Fig. l. Slopes and angies of a topographic profile along ; c-axis.

Fig. 3 .

 3 Fig. 3. Typical cyclic sbear curve showing the four portions.

Fig. 4 .

 4 Fig. 4. Test samples. a) regulär undulations b) hammered surface c) schist joint replica.

5 Fig. 5 .

 55 Fig. 5. Peak shear stress of the unmated hammered granite planar joint predicted by the new non-dilatant shear criterion.

Fig. 6 .

 6 Fig. 6. Peak shear stress of the regularly undulated mortar joint predicted by the new dilatant shear criterion.

Fig. 8 . 5 Fig. 9 .

 859 Fig.8. Peak shear stress predicted with various shear strength criteria and compared to the new dilatant shear strength criterion for the regularly undulated joint. g.5.3"

Figs. 8 &

 8 Figs.8 & 9 shows the peak shear strength predicted by the new dilatant shear strength criterion (Eq.[START_REF] Lefevre | Comportement mecanique et morphologique des discontinuites en cisaillement[END_REF] for the undulated joint and the schist joint replica respectively. Exanunation of these figures shows that the proposed failure criterion and Barton's criterion are in good agreement with the experimental data.

Table 2 .

 2 Test samples initial

				morphological	parameters
		o	DR,	K	(mm)	< (mm)	(mm)
	Hamrnered	12.7	0.043	0.98	1.742	10	130
	Undulated	10.3	0.023	0.21	2.000	10	370
	Rough	11.9	0.045	0.43	8.703	20	20

Table 3 .

 3 Mechanical properties or hanunered joint (Rg. 4b)

			MPa)	h
	(MPa)	l Bt cycle	5" 1 cycle	(°)
	0.3	0.220	0.251
	1.2	1.009	1.149
	4.0	3.003	3.613
	Table 4. Mechanical properties	of schist joint replica (Fig. 4c)
	cr n (MPa)		MPa)	ip 0
	0.4 0.8	0.637 1.147	19.70 15.50
	1.2	1.665	12.62
	1.8	2.356	9.20
	2.4	2.872	7.03

Table 5 .

 5 Mechanical properties of undulated joint (Fig.4a)

						3.5"
	(MPa)	I 1 * cycle	(MPa) 10 th cycle	h l ist cycie	Wcycle (°)	3"
	0.5	0.4	0.4	17	16	S 2.5-
	1.0	1.0	1.2	15	13
	2.0	2.3	2.3	17	11
	3.0	3.0	3.5	16	11
	4.0 •	3.9	4.4	15	9.9	.M 1.5-
	5.0	4.5	5.3	16	9.1
	6.0	5.32	6.0	16	7.3
	Fig. 5 shows the predicted peak shear stress by the
	new non-dialtant shear strength criterion (Eq. 10)
	for the hammered granite joint after the first and the
	last (5'1') cycle of sheanng and compared to
	experimental data. This figure shows that this new
	non-dilatant shear strength model is able to well
	predict the shear strength of non-dilatant or
	moderately dilatant joints. The peak shear strength
	predicted by the new dilatant shear strength model
	(Eq. 19) for the regularly undulated mortar joint and
	the schist joint replica are presented on Figs. 6 & 7.
	From these figures, good agreements between
	experirnents and predictions is observed.

Table 6 .

 6 Mechanical properties of schist joint replica (Fig.4c)

	(MPa)
	0.4
	0.4
	0.8
	1.2

Table 7 .

 7

		K "	'p CNK	Tp exp.	T p Eq.(26)
		(MPa/nim)	Eq -(27)	(MPa)	(MPa)
		0.16	O 10.42	1.40	1.54
		1.00	5.17	3.10	3.94
		0.32	9.13	2.35	2.49
		0.48	8.01	2.15	2.84
		Mechanical properties of undulated joint	(Fig. 4a)
	«ni	K,	'P one	TP exp.	i p Eq. (26)
	(MPa) 2.0	(MPa/nim) 0.0	n 10.20	, (MPa) 2.1	(MPa) 2.46
	2.0	1.0	7.85	3.8	-4.34
	2.0	2.0	6.04	5.2	'5.66