SANCOLD Conference 2017: “Management of Dams and Reservoirs in Southern Africa”

Centurion, Tshwane, South Africa, 15 to 17 November 2017 © SANCOLD, ISBN 978-0-620-76981-5

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SPILLWAY MODIFICATION REQUIRES AMENDED ENERGY DISSIPATION AT THE TOE OF A DAM

A Roos 1, D Cameron-Ellis 1, HJ Wright 1 1. ARQ Consulting Engineers (Ltd) Pty. Dams and Hydro, Pretoria, South Africa

PRESENTER: DAVID CAMERON-ELLIS

ABSTRACT

Discharge over an ogee crested gravity dam ultimately flows as a uniform sheet at the toe of the dam from where a stilling basin, roller bucket of similar energy dissipation structure serves to manage residual energy and safely return the flow to the river. However, when an ogee is replaced with a labyrinth or PK weir during a raising, the flow regime at the toe of the dam is changed significantly. Conventional energy dissipation systems may be inadequate to manage this flow, especially when concentrated jets impact beyond the toe of the dam. In this paper, the authors discuss the physical model study investigations undertaken for the raising of Tzaneen Dam, with specific reference to the energy dissipation requirements associated with the changed crest configuration, and provide some insight into the robustness of conventional slotted roller bucket energy dissipators.

1. INTRODUCTION

The proposed Raising of Tzaneen Dam was defined at preliminary design level as part of the Groot Letaba River Water Development Project (GLeWaP). A raising of 3 m was recommended to augment water supply to the area and taken to implementation phase.

The dam was originally constructed with a conventional uncontrolled 90.47 m long ogee spillway and a slotted roller bucket to accommodate energy dissipation at the toe of the dam. The raising requires that the upper 4 m of the existing ogee crest be demolished. This will create a stable platform for the raising, which initially was envisioned to be a modified Piano Key (PK) Weir. Through a series of physical scale model investigations, the original PK weir concept was optimised as a composite spillway utilising 4 cycles of a 6.1 m high labyrinth with ogee portions on the outer monoliths. The raising includes an additional fill of 1.5 m on the Non-Overspill Crest (NOC).

While the design of the hybrid ogee/labyrinth spillway is not dealt with in this paper, the authors discuss the investigations undertaken to address the modified flow regime and energy dissipation configuration at the toe of the dam, by comparing the functioning of an impact slab with that of the existing slotted roller bucket.

2. SPILLWAY LAYOUT

The raised spillway plan layout is shown in Figure 1 and typical sections through the raised spillway in Figure 2.

3. DESCRIPTION OF THE HYDRAULIC SCALE MODEL

3.1 Physical Scale Model

The hydraulic model study was conducted at the Department of Water Affairs and Sanitation’s (DWS’s) Hydraulics Laboratory in Pretoria West. The model was constructed by personnel of the DWS Hydraulic Laboratory, who also assisted with the modelling.

The model was constructed to a 1:44 linear scale and, taking advantage of symmetry, only half the spillway was modelled. The maximum prototype flow of 2 950 m3/s corresponds to 103 l/s in the model.

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Flow stabilisers were provided to ensure a smooth approach to the spillway structure. The ogee crest was shaped using concrete and wood with a smooth finish. The labyrinth was constructed using plywood and the curved outlet channels were formed by shaped Perspex panels. Perspex was used at the sides of the model to ensure that the flow nappe against the sidewall could be observed and measured.

Figure 1. Spillway Plan Layout

Figure 2. Typical Raised Spillway Sections

The impact slab at the toe was constructed with plywood panels with slots for pressure sensors to measure pressure fluctuations on the impact slab. A Perspex sidewall next to the energy dissipation structure was used to visually inspect the flow characteristics of the overflowing water. Later the impact slab was replaced with a model of the existing roller bucket configuration to evaluate the performance of the roller bucket with the final spillway configuration. The complexity of the roller bucket made the task of constructing it out of wood difficult and it was therefore 3D printed, using the original design drawing as input. A 3D CAD model was scaled down and printed with a layer height of 0.25 mm using Polylactic Acid (PLA) as the print medium. Thereafter the parts were finished with XTC-3D paint to produce a smooth finish.

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The downstream channel was modelled with a combination of wood and bricks. A sluice gate was used to adjust tail water levels and a crump weir was provided to confirm the flow over the spillway. Figure 3 below shows a perspective view of the model setup in the DWS laboratory.

Figure 3. Perspective View of Model

3.2 Pressure Testing Equipment

Special equipment was required to measure the pressure fluctuations caused by the impact of overflowing water on the dam toe. Three pressure transducers linked to a custom PicoLog 1216 logger were selected. The data logger had multichannel data acquisition capabilities of up to 16 inputs from various sensors and a sampling rate of up to 1MS/s (1 000Hz), which was acceptable for the tests.

The WIKA model S-20 pressure transducer was accepted as the most appropriate for the required measurements. The S-20 transducer is available in a number of versions and the IP68 rated version with an absolute working pressure range of 0 Bar to 1.6 Bar was chosen. This enabled the measurement of any sub atmospheric pressures up to 6 m of water head.

4. IMPACT SLAB EVALUATION

4.1 General

Through a process of optimisation, various configurations of the spillway were investigated, commencing with a form of PK weir initially suggested in earlier model studies undertaken for the raising of the dam. All of the configurations investigated included significant overhangs of labyrinth bucket or PK weir bay towards the downstream side of the structure, with certain configurations projecting the discharge nappe beyond the toe of the dam. Consequently, the spillway discharge no longer entered the slotted roller bucket as a stream of uniform depth, substantially compromising its energy dissipation function. An impact slab was provided downstream of the dam toe to counter the potential erosion.

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4.2 Functioning of the Impact Slab

Impact slabs function by resisting pressure fluctuations through their mass and supporting dowels, which mobilise a large enough portion of rock to resist the uplift pressures. The configuration of the initially proposed impact slab is shown in Figure 4 below: note that the dowels are not shown for clarity.

Figure 4. Impact Slab Configuration and Details

The rock mass beneath the impact slab would be exposed to pulsing forces during a flood event due to water spilling over the spillway, and plummeting about 40 m into the tail water, which would be about 8 m deep for the Recommended Design Discharge (RDD) event. This corresponds to a head difference between the reservoir level and the impact slab of 48 m. Pressures were recorded at strategic locations on the impact slab, with the sensors in terms of prototype dimensions in rows at approximately 5 m spacings in a downstream direction and 15 m along the length of spillway, with the rows lined up with the centre of the outgoing bays of the labyrinth.

4.3 HYDRAULIC MODEL TEST RESULTS

It was found that the critical flows that produced the highest peak pressure fluctuations were the floods lower than a Recommended Design Flow (RDF) event. The critical pressures recorded during testing indicated peak pressures of between 97 m and - 32 m (in terms of prototype pressures however the negative pressure would be limited to about -9,8 m in the prototype) as recorded during an event corresponding to 66% of the RDD. This was due to the tail water level at that specific flow providing less dissipation. Figure 5 below shows the recorded pressures for the critical sensor, together with the peak 95th, 98th and 99th percentile negative pressures.

Figure 5. Pressures measured on the impact slab

5. SPILLWAY OPTIMISATION AND INVESTIGATION OF ALTERNATIVE SOLUTIONS

5.1 Revised Crest Configuration

The significant negative and potential uplift pressures for the initial spillway configuration considered would require significant anchorage into the foundation to mobilise enough rock mass along with the impact slab itself to resist the fluctuating forces, even considering the distribution of resistance provided by the concrete slab.

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The pressures measured at the toe of the dam were also utilised in the optimisation of the spillway crest arrangements. The configuration of the spillway was tending towards a labyrinth with a concrete infill on the outgoing bay which deflected the flows away from the dam, as shown in the left hand section in Figure 6. While this provided additional mass of concrete and improved the aesthetics of the structure, the shaping concentrated the flows on the impact slab at the toe, leading to significant negative pressures. The labyrinth bays were accordingly smoothed as shown in the right hand section in Figure 6, to retain the water on the spillway face and reduce flow concentrations at the toe.

Figure 6. Modification made to the Outgoing Bay of the Labyrinth

The change reduced the maximum measured uplift pressure beneath the impact slab from the 95th percentile 15.4 m to 5.5 m. This improvement was significant, but resulted in reduced energy dissipation and consequently increased potential scour for a significant distance downstream. Figure 7 shows the functioning of the impact slab at the design flow following this optimisation.

Figure 7. Impact Slab Functioning at the Design Flow

Accordingly, while an impact slab proved viable, it represented a costly option due to the extended scour protection require and it was decided to evaluate the existing roller bucket design to determine it’s functioning with the proposed labyrinth spillway.

5.2 Functioning of the Roller Bucket

5.2.1 Roller Bucket Theory

USBR Engineering Monograph EM24 (USBR, 1984) describes the various types of flow conditions, in relation to tail water levels, for discharges returned to a river channel from a steep-sloped spillway, as illustrated in Figure 8. When the tail water is too low to allow for proper bucket performance, sweepout occurs (Tsweep). A tail water depth just safely more than the sweepout depth for the bucket still to function satisfactorily, is known as the minimum tail water limit (Tmin). On the other hand, if the tail water is too high, submergence of the roller bucket occurs. At this condition excessive turbulence is normally visible

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and flow diving from the apron scours the downstream channel. The upper limit, where diving does not yet occur, is known as the maximum tail water level (Tmax).

Figure 8. Description of Flow Scenarios (Obtained from USBR, 1984)

5.2.2 Existing Roller Bucket Testing

The bucket configuration and invert level of RL 679.70 m was obtained from the original design drawings for Tzaneen Dam. Using this invert level, the USBR Monograph EM24 deterministic considerations indicated that the maximum tail water level would influence the bucket functioning above a flow of approximately 2 950 m3/s, which is equal to the routed Safety Evaluation Flow (SEF). This can be seen in Figure 9 which also shows a comparison of the modelled values. It can be seen as flows approach the SEF, the Tmax water level approaches and intersects the tail water level. This indicates that flow Scenarios C and D in Figure 8 would be expected at the maximum spillway discharge.

Figure 9. Comparison of Model Test Depths with Deterministic Depths

Figure 9 above shows that the model indicates slightly higher Tsweep and slightly lower Tmin depths than predicted by the Monograph EM24. This is due to the highly irregular crossflows and flow concentrations produced by the labyrinth spillway, causing water to come into contact with the roller bucket inconsistently and allowing recirculation onto the roller bucket. Figure 10 below shows sweepout conditions at the Recommended Design Discharge (RDD) of 954 m3/s, as well as the inconsistent flow in the roller bucket.

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Figure 10. Sweepout Conditions at RDD showing Recirculating and Concentrated Flow

5.2.3 Discussion of Observed Flow Patterns

The flow concentrations which are evident in Figure 10 are contrary to the design principles applicable to the original roller bucket, which was based on the assumption of uniform flow conditions over the width of the roller bucket. The variation implies that portions of the roller bucket operate at higher or lower flow rates than the mean uniform flow condition and accordingly the actual Tmin, Tmax and Tsweep will not correspond with the requirements applicable to uniform flow conditions over the full width of the spillway.

To evaluate the phenomenon, video footage taken during the tests was studied in 5 m discrete sections and the flow variation along the width of the structure estimated. This was plotted against the uniform flow condition as shown in Figure 11.

Figure 11. Flow Variation along the Roller Bucket for Half the Spillway

The maximum variation in flow for the RDD (954 m3/s) was estimated as approximately +25% more and -45% less than the average flow, while for a flow of 1 950 m3/s the deviations were +24% and -15% respectively.

Figure 12 below shows the prototype tailwater rating curve and the required Tmin, Tmax and Tsweep elevations calculated utilising Monograph EM24 formulations for the existing bucket configuration. Good operation of the bucket is ensured where the tailwater curve lies between the Tmin and Tmax curves.

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As example of the effect of the flow variation on the operation of the roller bucket, lines indicating a uniform flow of 1 950 m3/s (midway between RDD and SEF) and the associated local diffused (-15%) and concentrated (+25%) flows are superimposed on the design curves. The points where the latter flow concentrations intersect the Tmin, Tmax and Tsweep curves indicate the operational constraint conditions applicable to those flows. Projecting these points back to the uniform flow line provides an indication of the reduction in operational range that can be expected under non-uniform flow conditions. Specifically, the required Tmax value is significantly lower and the Tsweep value is higher than would be the case for a uniform flow condition.

Figure 12. Concentrated Flow Tmin, Tmax, Tsweep Determination

Further, while projection of Tmin from the diffused flow serves to lower the required value at the uniform flow point. The Tsweep and Tmin points accordingly move closer together, confirming the mechanism which causes the associated observed curves shown in Figure 9 to converge towards each other from the deterministic values predicted by EM 24.

At higher flows, the narrowing of the band between Tmin and Tmax will be exacerbated, with the tailwater falling outside the projected band, implying both sweepout and/or diving flow should be experienced.

5.3 Scour Downstream of Dam Toe

While geotechnical core drilling through the body of the dam and some distance downstream thereof confirmed sound granitic gneiss is present in the foundation of the dam, no direct evidence of the dolerite band under the spillway mentioned in earlier reports was obtained. This being said, bathometry undertaken at the toe of the dam indicated an approximate 3 m deep scour hole on the left side of the spillway return channel. Without dewatering the tailpond, it is accordingly uncertain whether this is as a result of scour of the dolerite intrusion or through continuous use of the outlet works, or both. It nevertheless indicates that the rock downstream of the dam may be prone to further erosion under the revised flow regime at the toe of the dam.

Scour measurements in the model were taken for different flood conditions at the associated expected tail water levels. The typical scour pattern associated with diving flow (USBR, 1984) is shown in Figure 13. This normally results in a scour hole downstream of the bucket lip, whilst deposition occurs further downstream.

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Figure 13. Typical Scour Patterns (USBR, 1984)

For scour prediction, a 6 mm aggregate, corresponding to rocks with a diameter of 264 mm in the prototype, was used. Very limited scour was detected at floods at or below the RDD over the bulk of the spillway width, but a clear disturbance was noticeable directly downstream of the bucket adjacent to the prototype sidewall.

It was only at floods larger than the RDD that the diving of flow (Figure 13) and severe scour were observed, although the apparent flow pattern was very inconsistent, with scour occurring only periodically.

Unfortunately, the laboratory’s pump chambers flooded during February 2017 and significantly reduced the pumping capacity in the laboratory, preventing further testing at high flood conditions. An attempt was made to test the model at the maximum available flow to obtain some indication of the functioning of the roller bucket. The resulting scour from flows of 1 950 m3/s is shown in Figure 14. Scour was observed against the right sidewall to a distance of 25 m downstream of the toe, and the eroded material was deposited about 40 m downstream of the dam toe. In the centre of the spillway channel (left side of model due to symmetry) the respective distances were 17 m and 30 m. These observations confirmed that the slotted roller bucket operates in a narrower envelope than predicted by the USBR monographs due to the flow concentrations created by the proposed spillway configuration.

Figure 14. Scour and Deposition Patterns at 1 950 m3/s.

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Despite the flow concentrations, the roller bucket generally functions as originally designed at the lower flow concentrations. Under these conditions Tmax is clearly not a constraint, as can be seen in Figure 9. While the Tmin criterion is affected by the flow concentrations, the tail water depth is greater than the criteria and most of the energy will be safely dissipated under these conditions.

For higher flow concentrations, energy dissipation is less efficient and some scour may develop. Testing at the maximum available capacity of the laboratory showed that scour will occur for the higher order floods and therefore it was deemed prudent to include scour protection measures to counter this. To this end, a concrete slab will be included downstream of the roller bucket. This will simultaneously address the existing scour depression evident in the bathometry of the tail pond.

Anchorage of the concrete slab was based on the measurements taken for the original impact slab design. Figure 15 shows the scour protection slab downstream of the existing slotted roller bucket.

Figure 15: Slotted Roller Bucket with Scour Protection Slab (Anchorage not shown for clarity)

The above configuration is still to be tested at SEF to confirm that the proposed slab provides adequate scour protection.

6. CONCLUSIONS AND RECOMMENDATIONS

The hydraulic model study conducted for the Tzaneen Dam Raising assisted with the design of the spillway and energy dissipation measures necessary at the toe of the dam. Modelling led to the identification of several potential design flaws that were rectified by making certain modifications to the configuration at the toe of the dam. The following conclusions were drawn:

• Iterative changes to the model facilitated optimisation of both the labyrinth structure as well as the energy dissipation facilities at the toe of the dam.

• Irregular flow patterns consisting of flow concentrations and diffusions were observed downstream of the dam, as is generally expected with labyrinth spillways.

• As a result of these flow concentrations, high positive and negative pressures are induced at the toe, requiring protection against scour.

• While an impact slab to manage the varying pressures is viable, optimisation of the new spillway structure showed that the existing slotted roller bucket could be utilised together with a simple erosion protection slab.

• Due to the flow concentrations created by the labyrinth, the flows will vary along the slotted roller bucket. Investigation showed that the associated design criteria of Tsweep, Tmin and Tmax limit the structure to a narrower band than predicted using the USBR design monographs, with the flows above the RDF falling outside the criteria.

• Scour measurements downstream of the roller bucket indicated that scour is not expected for flows up to the RDF, but that it will occur for higher flows and therefore scour protection in the form of an extension of the concrete at the toe of the roller bucket is required. While, in terms of SANCOLD guidelines, a certain amount of damage can be allowed for flows in excess of the RDF the condition of the rock downstream of the dam is unknown. It was accordingly considered prudent to make provision for this scour protection, pending a more detailed evaluation of the rock conditions once the tailpond is dewatered, specifically in light of the scour that has already been identified.

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7. ACKNOWLEDGEMENTS

The authors thank DWS for their permission to publish this paper. The opinions and views presented in this paper are, however, those of the authors and do not necessarily reflect those of DWS.

The personnel from the DWS Hydraulics Laboratory, specifically Messrs Johannes Matlala, Patrick Mpyana and Freddy Khoza, are thanked for their assistance with the construction of the hydraulic model.

Messrs Johan Hattingh and Chris Hattingh of DWS are thanked for their assistance with the planning, coordination and testing of the Phase 1 testing. Messrs Stephen Anderson, Henry-John Wright and Alto Roos of ARQ Consulting Engineers were responsible for specific phases of testing.

The Approved Professional Person (APP), Dr Quentin Shaw is thanked for their input, critical comments and assistance during testing of the model.

8. REFERENCES

United States Bureau of Reclamation. 1984 Engineering Monograph No. 24. Hydraulic Design of Stilling Basins and Energy Dissipators. Eighth Printing: May 1984. Denver, USA.

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