Model studies of the Chucás hydroelectric project

Transcription

1 Model studies of the Chucás hydroelectric project L. Capuozzo (1), M. Gil Flores (1), O. Jiménez (2) & D. Bacchiega (3) (1) ENEL Green Power PH Chucás, (2) Carbon Ingeniería S.A., Costa Rica, & (3) INA, Argentina ABSTRACT The Chucás hydroelectric project under construction by Enel Green Power Costa Rica, is located about 40 km west of San José, the capital. It consists of damming the Tárcoles river by a 63 m height dam, then the water is transported about 400 m downstream by a 6.5 m penstock, up to a powerhouse where 2 Francis type units will generate a total output of 50 MW. For flood control, the dam is equipped with four radial gates, with dimension 15x12,4 m, capable of discharging 5,680 m 3 /s, under design conditions and 8,100 m 3 /s, as a verification flood. This large discharge is handled through a sky-jump spillway chute discharging in the rock bed downstream of the dam. A critical point of this kind of spillway is the plunge pool, which ineffably will form because rock scour under the free falling jet. A big deal of research has been done to assess the extension of this scour, first by empirical formulae and more recently, by a detailed experimentation of the dynamic pressures in the rock joints. Despite of all these advances, for important dams like Chucas, a physical model study is still indispensable. This model, on a scale 1:65, was tested at the hydraulic laboratory of INA, Argentina. The model includes a good deal of the approaching reservoir, the complete dam and spillway, and a large section downstream, covering the zone of the plunge pool plus a zone where the powerhouse discharges back to the river. This paper presents the results regarding the necessity of a pre-excavated plunge pool with the aim of avoiding future uncontrolled scouring. Also, the effects of the spillway discharge into the powerhouse area were tested, requiring some modification to the original project. INTRODUCTION The Chucás hydroelectric project, owned by ENEL Costa Rica, is located about 40 km west of San José, the capital. The project, now is under construction, dams the Tárcoles River by a 63 m height dam, then the water is transported about 400 m downstream by a 6.5 m penstock, up to a powerhouse where 2 Francis type units will generate a total output of 50 MW. Fig. 1 shows the general layout of the project. For flood control, the dam is equipped with four radial gates, with dimensions 15x12,4 m, capable of discharging 5,400 m 3 /s under design conditions and up to 8,100 m 3 /s, as a verification flood. These large discharges are handled through a sky-jump spillway chute discharging in the rock bed downstream of the dam. A critical consideration in this kind of spillway is the plunge pool, which ineffably will form because rock scour under the free falling jet. In this particular case, because of the exposed penstock layout, there is a risk that the lateral expansion of plunge pool could affect this structure. Additional, there is always the risk that a backward erosion of the plunge pool may affect the dam foundations. As may be seen in Fig. 1, the power house is located about 200 m downstream of the plunge pool, and that location coincides with an abrupt 90 bend in the river. This particular condition means that the powerhouse will be very exposed to the residual energy of the sky jump, as well as the effect of the strong flow curvature. These conditions for the plunge pool formation and risk of erosion were very difficult to assess by theoretically or by numerical means, and therefore, a physical model study was warranted. HYDROLOGICAL AND GEOLOGICAL CHARACTERISTICS The drainage area up to the dam site is about 1683 km2. The basin covers the main population centres of Costa Rica, including the capital San José, with elevation from about 2800 masl to 250 masl at the dam site. The rainfall is characterized by a defined rainy season between the months of May to November with

2 a relative decrease in the months of July and August; and a defined dry season in the months of December to April. The annual rainfall varies from 2000 to 3500 mm. Figure 1. General Layout of the Chucás Hydroelectric Project The flood discharge were calculated based on about 50 years of data from the Balsa hydrometric station, located few kilometres upstream of the project site. Based on standard statistical flood discharge methods, the flood discharges were estimated: the design discharge (1:1,000 yrs.) 5,400 m 3 /s, and the safety check flood (1:10,000 yrs.) 8,100 m 3 /s. The area where the Chucás Project is located consists mainly of volcanic rocks as andesitic lava, ignimbrite pyroclastic flows and occasionally thin layers of tuff and ash deposits. Good outcrops are observed on both banks of the river where the plunge pool is placed, however, at the river bed rock is not observed due to the constant water flow. The rock mass consists of aphanitic (fine grain) and aphanitic porphyritic (coarse grain) andesitic lava flow. The aphanitic mass is hard, highly fissured, and has good weathering resistance. The aphanitic porphyritic contains abundant calcite as joints filling giving a brecciated texture appearance (called the "coarse grain"), it is soft to moderately hard, moderately fissured andesite, susceptible to weathering (ISRM. 1978). The rock mass has been affected by hydrothermal process that makes it very susceptible to accelerated weathering when exposed to environmental conditions of the area (rain and sun). Local site investigation with boreholes indicates a variable RQD, however, at the left abutment rock is observed more fissured. The RQD percentage at the left abutment ranges from 0-76% with an overall average of 40%, which qualifies the rock mass as poor quality. At the right abutment rock conditions slightly improves, the RQD varies between 21 and 70%, with an average of 56%, although it is noted that the rock is more fissured between elevations to masl (RQD < 47%). The joints spacing produced rock blocks from 0.04 to 0.5 m length. These results would indicate that a plunge pool formation is very likely to occur relatively rapid; however, on the other hand, the hydrological

3 characteristics of the basin indicate that floods normally have relative short duration, in the order of few hours. LITERATURE REVIEW Energy dissipation downstream of high dams is a very delicate problem. According to Mason (1993), for medium heads from 30 to 50 m stilling basins are usually provided. For larger heads, cavitation becomes a problem and trajectory basin or ski jumps are currently widely used because they are a technically sound and economical way to control large quantities of water during flood season (Vischer & Hager 1998). This device is used mainly to increase the distance from the spillway to the place where the high velocity jet hits the channel bed, thus avoiding the danger of excessive scour immediately downstream of the structure. Due to the throw of the jet in the shape of a trajectory, energy dissipation takes place by internal friction within the jet, by the interaction between the jet and surrounding air, by the diffusion of the jet in the tailwater, and mainly by the impact on the channel bed. Inevitable, even in the hardest materials, rock erosion takes place creating natural plunge pools. In many cases, it is necessary to preexcavate the terrain in order to ensure the stability of the resulting pool. A deal of research has been done to assess the depth and extension of the scour caused by jets, particularly sky jumps. In order to evaluate the ultimate scour depth, many empirical and semi-empirical expressions have been developed, based on physical model tests or in prototype measurements. In general, scour formulae valid for plunging jet impact express the ultimate scour depth Y [m], defined as the scour depth beyond the original bed level, t [m], plus the tailwater depth, h [m], according to the specific discharge, q [m 2 /s], the fall height, H [m], and the characteristic particle diameter of the downstream riverbed, d [m], see Figure 2. Mason and Arumugam (1985) compared the application of 25 such formulas to 26 sets of scour data from prototypes and 47 sets of scour data from model tests. Their best fit of both model and to prototype conditions resulted in the following general form (A): Where K= *H 0.1, x=0.60-h/300, y=0.05+h/200,v=0.30, w=0.15, and z=0.10. It gives results with a standard deviation of 25% for model test conditions and 30% for prototype test conditions, with median values of 1.01 and Other formulas that were tried in relation to Chucás H.P. are: the Veronese (Mason 1993) which was developed mainly for alluvial material, and it is include here mainly for historical reasons. The Martins formula (Mason 1993) which was developed from data of 18 prototypes in rock beds. The Mason & Arumugan formula (B) (Mason 1993) which was quoted as the best fit for all model data. The Mason formula [10] that was developed taking into account the effect of air entrainment into the plunge pool, as represented by the air/water ratio β. The Damle formula (Khatsuria 2005) based on model and prototype data from Indian dams; it is an upper bound of the data. Finally, the INCYTH (12) formula that was tested against 50 prototype data, with a 26% standard deviation; it is also an upper bound of all the data but one case. Table 1 shows the coefficients and exponents for each one of these. In any case, most empirical equations are applicable to the specific case for which they were developed, and therefore very often physical models are used to evaluate the problem. However, laboratory tests are very sensitive, requiring to simulate the rock foundation by a material that adequately represents the dynamic behaviour of jointed rock. For this reason, most scour tests assume that the rock mass is already broken up and make use of crushed granular material to represent the scaled broken-up rock. However, such test conditions favour the formation of a downstream material bar, which generally results in an

4 underestimation of the total scour depth. Nevertheless, reasonable results can be obtained in terms of the ultimate scour depth (Balloert & Schleiss 2003), but the extension of the scour hole is often overestimated. Table 1. Coefficients and exponent used in formula [1] Formula K x y w V Z 1+β Veronese (1937) N/A 0 N/A Martins (1975) N/A 0 N/A Mason & Arumugam (1985) A Given in the text N/A Mason & Arumugam (1985) B N/A Mason (1993) N/A Damle (1966) N/A 0 N/A INCYTH (1992) N/A 0 N/A N/A: not applicable. Figure 2. Main Parameters in the plunge pool problem (H=fall height, h=tailwater depth, t=scour, α=angle of flip-bucket) More recently, the physical-mechanical processes that govern the phenomenon have been investigated. In short (Balloert & Schleiss 2003), scouring is a highly dynamic process, governed by the interaction of three phases: air-water-rock, intervening in a consecutive series of physical-mechanical processes: (1) aerated jet impact, (2) turbulent shear-layer diffusion in plunge pool, (3) fluctuating dynamic pressures at the water-rock interface, (4) propagation of these pressures into underlying rock joints and hydraulic fracturing of the rock, (5) dynamic uplift of single rock blocks, and finally (6) downstream displacement and/or deposition (mounding) of the broken-up material. On this line of research are worth mentioning the works of Spurr's (Spurr, 1985) method and Annandale's (1995) method, both corresponding to the erodibility index (EI) methods; Fiorotto and Rinaldo (1992); and Bollaert & Scheiss (2003) which aimed to include the dynamic pressures acting at a plunge pool bottom, and the resulting significant transient amplification when transferred into underlying rock joints. One important conclusion is that despite all the advances in the understanding of the physical phenomenon, these methods require extensive data about the rock condition, and require calibration of several parameters. If, in addition, there exist constrains in the geometry of the scour hole, clearly a physical model study is still indispensable, as it was the case for the Chucás sky jump. INITIAL ASSESSMENT OF THE PLUNGE POOL FORMATION

5 The total head of the Chucás dam from the upstream reservoir to the bottom of downstream river bed is 49 m. The velocity at the end of the flip bucket was estimated at about 20 m/s, and therefore no air intakes were deemed necessary. The location of the flip bucket insures that there would be no interference with the tailwater depth, up to a discharge of 5,400 m 3 /s, which is the design discharge of the spillway. The spillway comprises 4 tainter gates, 15 m wide by 12.4 m height. For the design flood the theoretical water level profiles resulted in a water height of 3.76 m and a velocity of 21.9 m/s at the end of the flip bucket; the corresponding Froude number was 3.6. A flip bucket radio of 16 m and a take-off angle of 40 were selected based on several recommendations given in the literature (Vischer & Hager 1998). Jet trajectory calculation indicates a trajectory length of about 54 m when hitting the downstream water elevation, and 72 m when reaching the original river bed (see Figure 2). In principle, this was considered good enough to protect the dam foundation. Table 2 lists results obtained with several formulas for estimating the scour depth. This if further illustrated in Figure 3. Figure 2. Spillway and initial plunge pool geometry Table 2 and Figure 3 show results for three conditions: the 200, 500 and 1000 years floods. As mentioned before, the floods in Chucás are of relative short duration (few hours), therefore it is estimated that the scour process will developed along several years, and not in a single event. Mason (1993) suggests that it is uneconomical to proportion the plunge pool for their optimum performance at the maximum flood; in fact, it can be just as important to look at much more frequent floods, for returns periods from 200 to 1000 years. Based on the above, and in consideration of the rock conditions a value of maximum scour of around t 20 m was adopted as a feasible result. For the different formulas, only two of them slightly exceed this value. Then, using relationships to estimate the development of the longitudinal erosion on the scour pool (Khatsuria 2005), the geometry shown in Fig. 2 was calculated. Very difficult to assess is the lateral extent of the pit scour, because the few available relationship do not exhibit adequate validity. This is very critical for Chucás scour pit since the right hand margin is a very steep rock scarp, and the left margin, although not as steep, supports the 6.5 m diameter steel penstock. Fig. 4 shows the initial

6 assumption regarding the lateral extend of the pit. In any case, it was decided from early stage, that the only way to determine the area of influence of the erosion pit with an appropriate degree of accuracy, was by use of a hydraulic model. Table 3. Scour computation data Return period T years Discharge Q m3/s Total head H m Tailwater depth H m Head up to flip bucket Z m Width B m Unit discharge Q m2/s Material mean diameter Dm m Angle of take off Α Air/Water ratio Β Figure 3. Scour depth results Scour Depth (m) yr 500 yr 1000 yr Figure 4. Estimation of lateral scour

7 THE HYDRAULIC MODEL In processes of local erosion downstream of dissipation structures, such as the resulting from the operation of a ski jump, vortexes large dimensions and low frequencies are the most important, being the macro-turbulence responsible for the phenomenon. Since viscosity has no significant influence on these swirls, the situation is well simulated if the model reproduces the turbulence structure of the prototype. It is noteworthy to mention that the local erosion in general is not dependent, in the long term, of the diameter of the particles involved. In addition to the above, it is important to evaluate the behaviour of the bed downstream around the plunge pool area, this in order to ensure that the model will make a proper diagnosis of the maximum depths of erosion and establish an approximation of the behaviour of the material removed from the riverbed. The particles present in the bed may be defined by two main parameters, the geometrical dimensions and the specific weight. The parameter representative of the size is the average diameter of d 50 grading curve, as other features such as its shape, are very difficult to represent. The selection of the length scale of the model took into account several aspects: Availability of physical space for the implementation of the physical model and accessibility to the model. Availability of the required discharge of laboratory facilities, in order to represent the maximum discharge in the prototype: m 3 /s. Quantification of the number of Reynolds to be larger than 10 4, Features of water feeding circuit, and the available discharge capacity. Constructive aspects of the model. A scale 1:65 was selected. The model was built at the Hydraulics Laboratory of the Instituto Nacional del Agua, Argentina, in their Large Model Workshop, which counts with 10,000 m 2 of covered area. The model includes a good deal of the approaching reservoir, the complete dam and spillway, and a large section downstream, covering the zone of the plunge pool plus a zone where the powerhouse discharges back to the river. The maximum required discharge is 365 l/s. MODEL RESULTS

8 The investigation program covered the following topics: 1) Testing the behaviour of the approaching flow to the dam, discharge, rapid channel, and flip bucket. Several cases with one or two closed gates were also tested. 2) Testing the model with fixed bed and pre-excavated plunge pool, with the initial geometry shown before. With this, the adequacy of the initial assumption about the plunge pool may be tested. 3) Testing the model with fixed bed without a plunge pool. This helps to verify the consequences of delaying the construction of a pre-excavated pool by few years. 4) Testing the model with movable bottom but considering limits in lateral movement, under the assumption of construction of rock reinforcement in the plunge pool. This tests the effect of a lateral restriction in the plunge pool evolution. On each case, the behaviour of the flow around the powerhouse bend was tested, in order to evaluate the impact on the structures. As of this writing, the following aspects were measured or evaluated: Reservoir Trajectories and streamlines of approaching flow to the dam. Flow into each spillway span, and study of flow asymmetries. Qualitative influence of piers and abutments. Spillway and ski jump Discharge capacity. Influence of piles on the spillway water profiles. Water level profiles. Operation with 1, 2 or 3 gates. Pressures along the spillway chute. Performance of the sky jump. Minimum take-off discharges at the end of bucket. Plunge pool and downstream reach Jet trajectories Impact area on the river and on both side slopes. Flow configuration in the impact zone and energy dissipation in the plunge pool. Flow configuration in the downstream reach including powerhouse area. Affectations on the penstock. Water levels at the powerhouse and tailrace area Water velocity measurements. Results as very numerous to describe in this paper, but some of the main findings and conclusion will be briefly described: The spillway, approaching flow, and gates, work well for all range of discharges, except for the verification flood (8100 m3/s). For this discharge, the most left bank gate shows a separation profile, creating a recirculation zone (see Fig. 5). This situation is abnormal although it may accepted since occurs at very high discharge, above the design value.

9 Figure 5. Spillway behaviour for 5400 m 3 /s (design discharge) The jump sky jet impinges slightly on the side valley walls. This situation needed the introduction of lateral deflectors in the end chutes, in order to concentrate the jet toward the plunge pool. They were tested, and a deflection angle of 15 was found adequate (see Fig. 6) Figure 6. Deflector of 15 ; impingement at downstream valley Q=5400 m 3 /s The pre-excavated pool is very effective in the dissipation of energy (see Fig. 7), for the range of operational discharges, but for the verification flood, where high residual energy is observed. As design criterion for the penstock, it had been accepted that the 1000 year flow (5400 m 3 /s) water levels should stay just below the bottom of the pipe. In this sense, what is seen in the model is acceptable, some flow waves overtop the level of the penstock terrace, but only slightly.

10 Figure 7. Behaviour of pre-excavated plunge pool=5400 m3/s For discharges above the 500 years flood, some water waves seem to affect the powerhouse terrace (see Fig. 8). As a design criterion for the powerhouse, no affectation should occur up to de 1000 years flood. Therefore, this situation will be addressed by means of a parapet to be built around the end of the penstock and powerhouse terraces. This parapet was also tested. Tests were carried out without the pre-built plunge pool, considering a fixed bottom as it is river now. This test is important in order to evaluate the possibility of delaying the construction of the plunge pool. In generally, the behaviour is good, except for increased turbulence around the downstream area. Testing is underway to measure pressure fluctuation along the river valley walls, using a mobile bed. This information will allow to evaluate the necessity of rock protection along the valley walls. Also, information about the maximum scour depth will be obtained. Although the testing program is still underway, the results clearly demonstrated the necessity to perform physical modelling in case of complex situation, as it is the case for the Chucás spillway. Also is clear the value of the physical model in order to understand complex flow problems, like the conditions around the penstock and powerhouse of the project. Figure 8. Behaviour of pre-excavated plunge pool=4,650 m3/s

11 REFERENCES Annandale, G.W., 1995 Erodibility. J. Hydr. Research, v.33, n. 4. Fiorotto V., Rinaldo A., 1992, Fluctuating Uplift and Lining Design in Spillway Stilling Basins, J. Hydr. Engrg., ASCE, 118(HY4) Balloert E., 2002, Transient water pressures in joints and formation of rock scour due to high-velocity jet impact. Communication 13. Laboratoire de Constructions Hydrauliques, Ecole Polytechnique Fédérale de Lausanne. Balloert E., Schleiss A., 2003, Scour of rock due to the impact of plunging high velocity jets Part I: A state of the art review, J. of Hyd. Research, v 41, n 5. Balloert E., Schleiss A., 2003, Scour of rock due to the impact of plunging high velocity jets Part II: Experimental results of dynamic pressures at pool bottoms and in one- and two-dimensional closed end rock joints, J. of Hyd. Research, v 41, n 5. Khatsuria R.M., 2005, Hydraulics of Spillways and Energy Dissipators, Marcel Dekker, N.Y. Lopardo R.A., Sly E., 1992, Constatación de la profundidad máxima de erosión aguas abajo de aliviaderos con salto esquí. Revista Latinoamericana de Hidráulica, no. 4. Mason P.J., Arumugan K., 1985, Free Jet Scour below dams and flip buckets, J. Hydr. Engrg, v.111, n. 2 Maison P.J., 1993, Practical guidelines for the design of flip buckets and plunge pools, Water Power & Dam Construction, Sep/Oct. Spurr, K.J.W, 1985, Energy Approach to Estimating Scour Downstream of a Large Dam, Water Power and Dam Construction, v 37, n11. Vischer, D.L., Hager W.H., 1998, Dam Hydraulics, Wiley.