ML19316A130
| ML19316A130 | |
| Person / Time | |
|---|---|
| Site: | Oconee  |
| Issue date: | 09/02/1977 |
| From: | Sams C, Sowers G LAW ENGINEERING TESTING CO. |
| To: | |
| References | |
| NUDOCS 7911280661 | |
| Download: ML19316A130 (51) | |
Text
{{#Wiki_filter:. . - - _ _ - - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ o V. . DUKE POWER CO31PANY JOCASSEE DA31 NORTHWESTERN SOUTil CAROLINA RiTllRii10 RIESllATORY WW RODE 01B ESTI3 TATE OF EXISTING STRAIN AND CRACKING POTENTIAL FRO 31 HYPOTHETICAL FOUNDATION DISPLACE 3 TENTS BY LAW ENGINEERING TESTING CO31PANY Charlotte. North Carolina REPORT PREPARED BY: Q 1914 - hy E. Sams. P.E. I Geotechnical Consultant Registered, South Carolina 3667 p George F. SowersM. AW Docketd 70'267 Senior Geo,ethnifal Consultant Registered, South Carolina 6231 CoWrol $ - D#e F,2/gp / af Dactimefit: REGutATORY DOCKET RLE
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o 9 INTRODUCTION i J This report is at the request of the Nuclear Regulatory Commission in their letter to Duke Power Company dated May 20, 1977. Duke Power has prepared a report to NRC presenting data relating to the safety of Jocassee Dam. However, the Duke report did not contain the following work requested by the NRC letter: l "The ability of the foundation of the dam to resist the effects ~ of potential fault movements should be assessed and reported. Past measurements of the settling, displacement and crachng of the dam should be interpreted to estimate the existing state of strain, particularly in the core of the dam. The additional l strain which can safely be tolerated should be estimated and related to the magnitude of potential fault movement; The tolerance of the abutment material to strain and cracking resulting from fault movement should be estimated based on the properties of the saprofites and the magnitude of potential I fault movement. If abutment cracking cannot be ruled out then the piping and erosional resistance of the weathered rock should be assessed." The purpose of this report is to respond to the above requests for information by NRC. 1
__ __ __ = _ . _ _ . _ . __ _ _ _ _ _ _ _ . _ _ _ . _ _ _ . . _ . 3 l i Summary and Conclusions L 1 l Using available measurements of settlement from the crossarm type vertical i ) settlement devices within the core of Jocassee dam and from reference points on the surface of the dam, and making reasonable assumptions for internal horizontal displacements, the existing state of strai.) within the core of the dam has been conservatively computed. The total strain capacity of the core and of the residual soll and weathered rock materials forming the abutments has been estimated from consolidated-undrained triaxial compression tests. ' An estimate was then made of the remaining settlement movements required to reach the total strain capacity. For a stretching condition that would cause tension cracking of the core, estimates of the extensional strain capacity at a location near the top of
- the core have been made using stress-strain characteristics deduced from the 4
consolidated-undrained triaxial compression tests. The results of the computation were in overall agreement with published values of extensional strain that apparently caused tension cracking with rapid straining. Based on .he triaxial tests and consolidation tests of the abutment materials, it was then deduced that the same approximate extensional strain limits could be I 1 applied to the residual soils as to the core, as a first approximation. Expressions were then presented for relating the average extension s'. rain to the hypothetical displacements of the rock foundatien for the full range c f heights of soil above the rock. Computations were then made of the tolerable rock displacements which, if exceeded, would cause tension cracking. l Finally, for a comparison with the previously computed disp;acements, the results of existing studies of sel'smicity, along with appropriate 2
e 1 1 results from the technical literature, were used to estimate a hypothetical foundation displacement (tectonic movement) associated with the maximum earthquake already calculated at Lake Jocassee. Conclusions available from the study may be tabulated as follows: (1) There is no evidence of any load transfer from the soil core i onto the granular shells during construction of the dam. Instead, the reverse has been true. Therefore, there is no indication of an existing potential weakness of the core that makes it susceptible to hydraulic fracturing on horizcntal or subhorizontal planes running longitudinally with the core. (2) Conditions in the core during the reservoir filling period were such as to inhibit the formation of tension cracks transverse to the dam and due to differential settlement. (3) Despite settlement strains the core still has ample capacity to absorb shear strains before the onset of shear fracture or plastic flow. (4) The appropriate ecde for foundation displacement surfaces i potentially most detrimental to the core in terms of extension strains near the crest was seen to be a plane striking transversely to the core. Reasonable assumptions for the dip of the plane from horizontal were determined to be 45 and 90 . 3
- s j
s - (5) The appropriate hypothetical mode of foundation movement was determined to be either normal (extensional) or reverse (compressive),wi th the maximum earthquake thought more likely to be associated with the normal displacement. The more critical normal movement was presumed. (6) It is estimated that tension cracking near the crest of the core or in the residual soils of the right abutment might occur if extensional strains from rapid foundation displace-ments exceed 0.5%. For the higher parts of the dam, about 350 ft or more, it is estimated that foundation displacements on a 45 dipping plane up to 34 in. could be tolerated before i cracking. The corresponding computed foundation displacements i on a vertically dipping plane, such as the N45 W Joint set, ) are 56 in before cracking. i j l (7) The computed tolerable movements before cracking of the 120 ft of soil and weathered rock of the right abutment are 12 inches for' shear displacements on a 45 dipping plane in the rock and 10 inches for displacements on a vertical plane within the rock, such as the rock joints. (8) The computed tolerable rock movements before cracking of a height of 45 ft of soll above rock, such as on the left abutment, are 4 to 5 inches for rock displacement on a 45 4 dipping plane and 7 inches for displacement on a vertical plane. However, the hard rock is only about 30 f t below full pool elevation for such a condition. 4
1 1 I (9) The amount of hypothetical foundation displacement associated with the maximum earthquake was computed to be 3 to 4 inches maximum, f (10) The computed maximum foundation displacements associated with the maximum earthquake, only 3 to 4 in., are thus substantially less than the calculated tolerable movements in ! the main part of the dam, and still less than the calculated ' t tolerable movements in the abutments. I l (11) At the core-foundation contact, the high confining strasses should prevent open cracking despite foundation displacement. 2 (12) Even if open cracking of the core could occur, the 8 ft wide filter zones will protect the core from erosion. l Jocassee Embankment and Reservoir Filling Rates 2 As shown on Figure 1, the diversion tunnel closure at Jocassee was on 4-8-71, and the water level was at elevation 800.09, about 95 ft below the embankment f surface at that time. From April,1971 until the end of January,1972, there t was a period of elatively rapid rise in the reservoir level to elevation 970, t i with embankment construction staying 65 to 75 ft ahead of the rising water. l Then came a period of slower rate of reservoir rise through October, 1972 with embankment construction staying 100 to 120 ft above the rising water. By
September, the core had reached its planned elevation 1120. Shortly after topping out the embankment to core level, another period of fralatively rapid 5
rise in reservoir level occurred, coming from about elevation 1015 to eleva-tion 1085 by July, 1973, 35 ft below the top of the core. Thereafter, the reservoir rose more slowly, reaching full pond elevation 1110 in July,1974. During this period there we4e drawdowns to about elevation 1080. i l The core of the dam became saturated as the reservoir was raised. Beginning in early 1972, the level of saturation was measured in the open tubes of the four vertical settlement devices. The water levels thus measured within the _ core are shown on Figure 1. The data show the core saturation level rose at about the same rate as did the reservoir level, in the upstream-down-stream direction, the saturation line dropped steeply from pond level at the { upstream slope of the " impervious" core to progressively lower elevations moving downstream through the core. In the tube of vertical settlement device 1, located 35 ft upstream of the dam centerline, the saturation line stayed about 15 to 20 ft below reservoir level. In vertical settlement devices 2 and 4, located 30 ft upstre== of the dam centerline, the saturation line stayed about 40 ft below the rising reservoir level. In vertical settlement device 6, located only 20 f t upstream of the dam centerline,the saturation line stayed 40 to 50 ft below the rising reservoir. Af ter the reservoir was filled in July, 1974, the water level in VSD-6 stabilized about 50 ft below reservoir level, but then between May and June 1975. It assumed a higher level about 15 to 30 ft below the reservoir level. Measurements of Movement of Dam i The movements of Jocassee Dam were measured over varying periods of time by two methods: (1) six vertical settlement devices (USBR crossarm type) were 6
Installed during construction, four in the core and two in the downstream rockfill and random rockt ll shell; (2) seventeen surface reference points ! established in June and July 1973 af ter the dam was et leted. All the above are shown on Figure 2. This report utilizes the data from the vertical settlement devices (VSD's) and the 9 surface reference points along the crest of the dam, i Not shown on Figure 2 are the locations of: (1) three temporary surface reference points during early construction on the upstream rockfill shell at elevation 1051; (2) temporary surface pins on the ccmpleted crest; (3) the four additional monuments irstalled on the upstream side of the crest in July, 1975; and (4) two inclinometers and surface gage points installed In November 1973 at stations 14+46 and 14+92 offset 17.5 ft and 19.5 ft upstream of the dam centerline, respectively. These inclinometers, extending 4 40-45 f t below the crest of the dam, were installed to evaluate the depth of 4 movement causing longitudinal cracking along the dam crest. (The Consulting Board concluded that the cause of the longitudinal crack was unequal strains
- between the core and the upstream shell produced by the normal deflection and settlement of the dam. The crack did not involve the core material, based on data from the inclinometers).
Figure 3 is a longitudinal section through the dam core, showing prepared foundation line and the weathering profile on the abutments.
;
Only VSD's 1, 4 and 6 provided continuing measurements of dam core displace- . I ments during construction and until 9 to 10 months after the completion of core construction. VSD 2 became inoperable below elevation 1000 after 7
1 August 2, 1972, before core construction was completed; it was read above J elevation 1000 for the last time in February, 1973 The last reading taken on VSD 4 was July 16, 1973. VSD 1 became inoperable below elevation 1050 after November 10, 1975 but readings were taken above elevation 1050 on { 5-12-76. VSD 6 remained functional throughout its depth and its last j reading was taken 5-12-76. i
- There were mechanical problems with the sounding tool (torpedo) resulting in
~
l ) missed readings at vertical settlement devices between August and December 2 1971. i VSD 5 in the downstream shell was read through 6-26-71, some 3 months Sefore the shell reached its finished elevation at that location. VSD 5 was
- inoperable by the time the sounding torpedo was available again in sanuary,
1972. VSD 3 was not read af ter June 28, 1971 except for one reading on January 24, 1972. The surface monuments were estabilshed on the dam crest in June, 1973, and have been read for settlement since then to the present time. However, monument 9 was destroyed after the May, 1976 reading. The 8 monuments on the downstream benches in the slope were established in July,1973 and read until April 1975. The readings were discontinued because those on the elevation 810 berm had shown only a maximum of .02 f t settlement in the preceding year and those on the 925 berm had shown only about a maximum of .09 f t settlement in the preceding 6 months and about a maximum of 0.09 ft settlement in the 6 months preceding that. b i
Measurements of horizontal movements of the nine crest monuments were begun in July 1974, and have been continued at intervals to the present time (April, i 1977 is the latest reading). Settlecent During Construction and Associated Vertical Strains
- Figure 4 shows measured settlements during construction of the core and shell at VSD's 2, 3, 4 and 5 As of late June,1971 when embankment construction had progressed to about elevation 915-920 in the downstream shell, and to about elevation 940-950 in the core, the core settlements (at VSD's 2 and 4) had reached a maximum of about I ft, whereas those of the shell were between 1.8 ft and 2.2 ft (at VSD's 3 and 5, respectively) .
By the time the shell had reached near elevation 980 on 1-24-72 maximum j settlements had increased to almost 2 ft at VSD 3 location. The closest available reading in time at the deepest core data point, VSD 4, is 4-17-72, where maximum settlements were 1.8 ft with the core built to about elevation 1055 Interpolation between the 4-17-72 and 6-26-71 readings for VSD 4
- suggests core settlements were less than 1.5 f t when the core surface was at about elevation 980, compared to the nearly 2 ft in the shell at VSD 3 4
Figure 5 shows typical vertical settlement device data expressed as vertical strains versus embankment height above the middle of the layer depicted. Changes in thickness of the variouslayers on Figure 5 are divided by the initial thickness (typically 20 f t) to obtain the vertical strain. Variations of the curves for a given material may be caused by variations in the material and/or changes in the state of imposed stresses with elevation and location 9
of the layers. Figure SA shows that the rock and random zones of the shell apparently underwent similar amounts of vertical strain during the construction period, and both experienced greater strains than did the core (Figure SB). This cendition of less core settlement than shell settlement suggests that load transfer was occurring from the shell to the core during construction. In other words, the core was compressed by the greater shell settlement. Thus, there is no Indication of potential for tensile strains or hydraulic fracturing through the core on horizontal planes running perpendicular to the cross-section. Subsequent calculations in this report will deal with strains oriented within the longitudinal plane through the core (Figure 3), for potential fracturing on planes transverse to the dam axis. Settlement of Dam Crest During and Following Reservoir Filling The surface reference points were available in June and July, 1973 The vertical settlement devices that remained functional (1, 4 and 6) were
! available for measurement of settlements of the upper part of the dam after the core was completed and until July, 1973 The data are shown on Figure 6A as settlement versus elevation within the core. As the reservoir rose from elevation 1010 to elevation 1085, the uppermost crossarm within the core at VSD's 1, 4 and 6 showed approximate settlements of 0.40,1.10 (estimated) and 1.15 ft, respectively. These settlements correspond to average vertical strain increments (for the full core depth to rock below the crossarm at each vertical settlement device) of 0.24%, 0.32% and 0.44% for devices 1, 4 and 6, respecti"ely. It is not shown on Figure 6A, but data available to February, 1973, at VSD 2 shows a vertical strain increment at VSD-2 compatible with that of VSD 4 in the same time. It is logical to presume that, had VSD 2 continued i
i to function, it should have Indicated a. vertical strain increment on the order l 10
5 of 0.30% by July 1973. Some of the above dif ferences in measured core compression can be attributed to variation in de stress conditions due to the previously described varying locations upstream with respect to the dam crest. Figure 68 shows incremental settlements versus elevation within the core at - VSD's 1 and 6 during theperiod July 1973 to July 1974 and for VSD 1 durin g 1 the period July 1974 to July 1975 - Using the information on Figure 3 about the depth to prepared foundation along the entire length of the core, the above strain increments were Interpolated and extrapolated to prepare the plot of total crest settlements shown on the uppermost part of Figure 7 for the period End of Construction to July,1973 In the middle of Figure 7 are the settlements measured af ter July,1973, for the surface monuments 1-9 These latter data since July, 1973, show larger increments of settlement at points 7, 8 and 9 approaching the left abutment, which apparently continues to trend of greater settlement relative to core height previously established by vertical settlement device 6. 4 The maximum increment of settlement measured on the surface monuments during the period June / July 1973 until April 1977 was at point 8, station 17+00, where about I ft of settlement was measured. This is less than 0.3% of the dam height. Compared with other rockfill dams it is small. The total crest settlements since completion of the dam, obtained by adding i ! the extrapolated settlements before July 1973 to the measured settlements since 1 11
l i I I
July 1973, are shown at the bottom of Figure 7. The dam crest has settled nearly 2 feet between about stations 11+00 and 16+00. This is within the amount of settlement anticipated in the design, as the crest in this area was overbuilt typically 2 to 3 ft above nominal crest elevation 1125 This is about 0.5% of the maximum height. Figure 8 shows the post-construction continuing settlement-time data for selected crossarms from VSD's 1 and 6. These selected crossarms occupy _ positions near the top of the core and a few feet below the top of core, near the bottom of the core, and at a point within the core. These data show the , expected decreasing rate of core settlement with time, with the major settle-ment near the top of the core being conpleted by 1974. The data show the core at VSD 6 was compressing more than the core at VSD 1 during this period. As previously stated, it would thus appear this correlates with the larger settlements seen at points 7, 8 and 9, in tne middle part of Figure 7. There are no highly unusual aspects of the settlement data on Figure 8. Between May 6, 1975 and June 24, 1975, the data show that crossarms at VSD 6 rose in elevation by 0.11 to 0.13 feet. Af ter the " rise", settlement continued at the previous time-rate. The " rise" in elevation appears to correspond in time with the rise in saturation line at VSO-6 shown on Figure 1. No anomalous behavior is indicated during or after that time period at VSD 1. It is possible the indicated " rise" at VSD 6 is due to a change-over in personnel reading the data. The settlement versus time data for the surface monuments along the crest are shown on Figure 9, begianing in June 1973 These data show the same trends 12
l
I l l i 9 I mentioned previously in that by 1974 the crest settlement rate had reduced to
- a fairly constant value, appearing to be nearly linear with time.
l There is no indication of any sudden anomalous behavior displayed in the data l on Figure 9 There is a slight hint of larger rate of settlement at monument 4 thru 8 (stations 9 thru 17) during the period between April and October, 1975 However, there are similar increases in rate for other periods of time, such as between May and August 1974, that show in all the monuments. This
" stepped" behavior in the secondary compression range has been observed in long term laboratory consolidation tests on many soils and rockfill samples from other sites.
it has been assumed in the calculations that the top of the core underwent approximately the same deflections as those measured on surface monuments 1-9, which were located a few feet downstream of the crest shoulder. The validity of this assumption for the vertical deflections is confirmed by comparing the settlements measured af ter July,1973, of those crossarms situated at or a few feet below the surface of the core with the settlements measured on the surface monuments (Figures 8 and 9, respect vely). Horizontal Movements of the Dam Crest Along the Axis (Toward the Abutments) and Associated Strain Measurements of the horizontal cross-valley movements along the crest began in July, 1974. Calculations have been made for the earlier horizontal movements caused by differential settlement for a period af ter core completion. Using the estimated differential crest settlements, the observation that 2/3 of the 13
core crest settlement was from compression below about elevation 1050 (from Figure 6A) and using an empirical analogy from beam theory, the horizontal crest movement pattern, parallel to the axis, during the period September 1972 until July 1973 has been computed. It is shown on Figure 10A. The estimated maximum horizontal movement during that time was about 0.16 ft over the right abutment between stations 3+00 and 9+00. All horizontal movement is estimated to have been toward the deepest part of the dam, causing compression parallel to the axis between about stations 9+00 and 17+00, and extension elsewhere. The estimated maximum extensio-11 strains from the - data on Figure 10A for the period September 1972 to July 1973 are about 0.08 percent to 0.11 percent averaged over 100 to 150 ft horizontal distances between stations 1+55 (assumed as no movement) and station 3+00, and between stations 7+00 and 8+00. i Measurements of horizental chord distance between the surface reference points, in the direction of the axis of the dam, are available beginning July,1974. The resulting pattern of horizontal movement is shown on Figure 10B and the data are on Figure 11. The data points on Figure 11 show scatter, consistent with the precision obtainable in making such measurements by triangulation from reference points on each abutment. The pattern of movem6nt on Figure 108 was estimated from the data on Figure 11 by distributing the compressive change in length between stations 11+00 and 13+00 equally between surface reference points 5 and 6, and then adding the measured changes in length between other surface reference points. The horizontal movement pattern on Figure 10B is consistent with the measured settlement patterns in tne middle of Figure 7 The largest horizontal 14
movenents were taking place in the vicinity of points 7, 8 and 9, where larger movement (settlement) were measured. The estimated maximum extensional strains from the data on Figure 108 for the period July 1974 until April 13, 1977 are about .03 percent averaged over the distance between stations 1+55 and 3+00, and .2 percent averaged between stations 19+00 and 20+25 These amounts are inconsequential, considering that they occurred slowly, over a nearly 3 year period, during which time vertical compressive strains were also occurring and hindering any adverse effects (cracking). A Downstream Horizontal Deflections of the Dam Crest and Associated Strains Measurements of horizontal movements downstream have been made at time intervals from an initial position beginning July, 1974 through April 13,1977 (except for monument 9, which was destroyed after May, 1976). The deflection versus time data are shown on Figure 12 and the deflection pattern is on Figure 13 for the time span July 12, 1974 to April 13, 1977. The downstream deflection versus time data on Figure 12 show no sudden devia-tions or other anomalous behavior. The pattern of the downstream deflections on Figure 13 show the arched crest of the dam was moved downstream during the period July, 1974 to April, 1977. This means the length of the dam was shortening, owing to its curvature, and compressive strains were being induced parallel to the axis from this movement. The deflection was greatest above the deepest part of the dam: 0.85 ft at stations 11+00 and 15+00, and 0.91 ft at station 13+00. The change in strain (compressive) parallel to the axis caused by the 0.91 f t deflection, based on the radius of curvature of the dam (5730 ft) 15
is .008 percent average over the full crest len,.h. This average compressive strain increment, though small, would offset some of the extensional movements at the crest being caused by differential settlement. Of course, this effect is already built into the actual measurements of horizontal movement parallel to the axis. Discussion of Existing Horizontal Extension Strains at Too of Core l lt is generally accepted by the geotechnical engineering profession that a possible cause of cracking i~n the upper part of some completed embankments appears to have been differential settlement within the fill as it becomes wet from the rising reservoir. (1), (2) Horizontal movements occur within the zone of material above the seat of settlement. I The data on Figure 6A show that, during the period when the Jocassee reservoir rose fairly rapidly from about elevation 1010 at the end of core construction (Sept., 1972) to elevation 1085 by July 1973, settlements in the core involved compression throughout all or a major part of the entire core depth. In other words, the upper part of the core was undergoing compression as well as the lower parts where saturation was occurring. Thus, the data for the vertical settlement devices do not indicate that significant additional core compression at Jocassee occurred as a result of wetting by the rising reservoir. Instead, j l (1) Leonards, G. A. and Narain, J., " Flexibility of Clay and Cracking in Earth i Dams", ASCE, Journal of the Soil Mechanics ind Foundations Division, V89, i SM2, March, 1o63 (2) Lee, Kenneth L. and Shen, C.K. , " Horizontal Ho 'ements Related to Subsidence", ASCE, Journal of the Soil Mechanics and Foundations Division, V95, SM1, January,1969 16 l l
I settlement continued slowly as adjustment to the self-weight of the embankment. The fact that compression was occurring above as well as below the rising reservoir would inhibit the formation of transverse cracks in the core during the period of reservoir rise. (The data on Figure 6A also suggest that much of the surface settlement at the vertical settlement device locations as the reservoir rose was due to movements in the granular materials of the upstream shell above the core). _ lt is thus concluded that the existing state of strain at the crest of the Impervious core caused by settlement, saturation or water load has not caused incipient near-crest transverse cracking. Calculations have been made for the state of strain developed in the core at depth, and the results will be presented and discussed in the following section. Existing Strains Within the Core Differential settlement of the core involves shear strains deep within the dam because the foundation of the dam is not level. Examination of the prepared foundation profile on Figure 3, the crest settlement pattern on Figure 7, and ! the axial movements at the crest on Figure 10 indicates that strains developed within the core of the dam are probably greatest in the abutment areas, particularly near the abrupt slope changes at stations 8+50 on the right abutment and near 16+25 on the left abutment. Calculations were made for the existing strains within the core of the dam at these two irregularities and also between stations 1+55 and 4+00 on the right abutment. Near 16+25, VSD 6 allows strains to be calculated in the soil at depth between the actual device 17
i 4 , location and the rock slope of the abutment. For the irregularity in the rock ! slope of the right abutment, near 8+50, VSD's 1, 4 and o were assumed to represent the settlement conditions above the irregularity at the points where l I the rock elevation corresponded to the rock elevation at the actual device i i location. VSD 1, af ter correcting for less depth to rock, was used to calculate the strains at stations 1+55 to 4+00 on the right abutment. 1 i Since no measurements were taken of horizontal movements within the core,
; such movements were estimated from data obtained by Wilson (Figure 57 in the chapter " Deformation of Earth and Rockfill Dams", in Embankment-Dam 1
-i Engineering, Casagrande Volume, John Wiley & Sons, 1973). This figure is consistent with the horizontal movements estimated and measured at the crest I
of Jocassee Dam, and should thus apply to movements at depth too. The calculated maximum shear strains within the core are shown on Figure 14. We believe that the results on Figure 14 represent the most highly strained 1 portions of the core. These results show that significant gradients of strain exist near the abutment irregularities, with highest strains closest to the sloping rock. The strains developed at shallow depth below the crest are relatively low. These computed values are conservative for two reasons: (1) they neglect the effects of downstream deflection of the curved dam under water load; (2) the relieving of stresses induced by strain through long term creep. It could alternatively be argued that the settlements associated with secondary compression are a form of creep, or deformation without associated changes 18 t
- - , - -r,, .- - - _ - - --,y _ , . - - - ,-- --w -
in stress. Therefore, associated settlement strains are not related to failure strain. Strains Recuired to Produce Plastic Flow or Critical State Shear of Abutment ! and Core Materials Abutment materials: saturated, consolidated undrained triaxial tests were made on Denison barrel samples of the soft weathered rock and hard residual soils obtained in borings on the abutments made following initial foundation ~ i ! stripping. Most of the tested samples were obtained from borings made near ! station 5+00 on the right abutment. The sampled material was subsequently , l removed at that exact location, but similar material remains above rock l right of station 3+30. Samples were also tested from a boring about 40 ft i downstream of station 15+00 on the lef t abutment, and another about 600 f t upstream of approximately station 8+50 in the right foundation.
) The axial strain required to produce either plastic flow or critical state j shear was determined from stress-strain plots. This is referred to as
" strain capacity" in this report. It was plotted as a function of the effective ccnfining consolidation pressure in Figure 15A. The data show
- the critical axial strain of the weathered rock /hard soil increases some-i 1
- l what with effective confining pressure. At a mean confining pressure of 7 1
to 8 ksf, the failure strain is 10.7% to 13.6% ; at 28 to 29 ksf it is 12.9% to 16.8%. i i In addition to the consolidated-undrained triaxial shear tests used to prepare i Figure 15A, several consolidation tests were made on the samples of residual i soil and weathered rock from borings. These consolidation tests, combined i 19
with the stress-strain data from triaxial shear tests, yield the following average hyperbolic stress-strain parameters for the residual soils and weathered rock taken together: Poisson's ratio = 0.3 (assumed); K=123 (uncorrected for disturbance); n=0.64. The shear strength parameters assigned for the design calculations for the dam were c'=0, c '=38 , and, for consolidated-undrained conditions at normal stresses less than 10 ksf, C=400 psf, 0=37 . Core Materials: The saturated, consolidated-undrained triaxial tests made on undisturbed samples of field-compacted core soil show that the strain at failure of the core material increases with effective consolidating pressure, or the mean effective confining pressure before straining. The data are shown on Figure ISB. In the core, the coefficient of effective earth pressure at rest was assumed at about 0.5, (Poisson's Ratio = 0.35), thus relating the horizontal and vertical consolidating stresses before straining. The vertical overburden stress was assumed to be equal to the major normal stress for purposes of estimating the strain capacity at various elevations in the core. This conservatively neglects the additional confining stresses added by the downdrag of the shell which settled more than the core. In Figure ISB, the major effective normal stress before straining has been shown, rather than the mean confining stress, which is o'3 in the laboratory. The strain capacity at various elevations in the core, based on the mean of the data in Figure ISB and the above assumptions, is shown on Figure 16. The o' y value for transferring data from Figure 15B to Figure 16 was cal-culated by conservatively assuming the water level in the core to be at full pond level. Figure 16 shows strain capacities of the core materials ranging from 6.5% to 7% near the crest up to 10% to 11% in the deepest part of the core (below elevation 800). 20
;
Comparing Figure 16 with Figure 14 indicates the core materials at Jocassee, despite strains Induced by settlement, still have a capacity to absorb shear strain before plastic flow occurs. The soils immediately below the crest appear to have a remaining strain capacity in the range of 5% or more. The soils between elevation 1100 and 1000 ft have an estimated total strain capacity of 7% to 8%, with existing strains estimated to be less than 2% leaving 5% to 6% available strain capacity. Between elevations 1000 and 900, the estimated total strain capacity is 8% to 9.3%, with existing strains estimated at less than 4%, leaving 4% to 5.5% available strain capacity. ' Between elevations 900 and 800, the estimated total strain capacity is 9.3% to 10.3%, with existing strains estimated as less than 6.5%, leaving 3% to 4% available strain capacity. Below elevation 800, the estimated total strain capacity is 10.3% to 11%, and the existing strain is estimated as less than 4% to 6%, leaving at least 4% to 5% available strain capacity. As was mentioned previously, these values are conservative for two reasons: (1) they neglect the effects of downstream deflection of the curved dam under water load; (2) more significantly, they neglect the relieving effect of long term creep. It could be argued that the settlement associated with secondary compression (since about 1974) are in the form of creep; therefore, associated strains should not reduce the strain capacity of th'e soil. FOUNDATION DISPLACEMENTS--APPROPRIATE DIRECTIONS AND TYPE OF MOVEMENT in the discussion below, the following definitions are applied in describing the foundation displacements. Normal displacement implies vertical movement caused by horizontal extension (Figure 18C). In normal displacement the 21
. . _ - _ . ~_. _ - - . - __
distance between points on opposite sides of the zone of displacement is increased. Reverse displacement implies vertical movement caused by horizontal compression. The distance between points on opposite sides of the reverse displacement zone is thus decreased. Strike-slip motion is horizontal (see Figure 18A) . i The hypothetical foundation displacements to be considered at Jocassee Dam are complicated by a consistent reverse-faulting focal mechanism below the dam computed from the seismograph data (see the Law report of September, 1976, Appendix 4 of Duke's Summary Report) . This focal mechanism differs from
~
that assumed for the maximum earthauake. Consecuently, more than one direction j' of movement will be evaluated. , l (1) Direction and Dip of Hypothesized Movement Surfaces
- a. The maximum earthqtake was hypothesized to have a fault plane that strikes N84 E and dips 45 SE. For convenience in computation in this study,the strike was considered anywhere within 15 of the hypothetical (N81 W to N69 E) and the dip within 10 . Figure 17C & D shows the strike of this surface superimposed on a plan view of the outilne of the core section,
- b. An area of reverse focal mechanism has been computed based on the observed small events below the dam. Displace-ments for this stress orientation would be en planes striking l N50 to 70 W and dipping 40 to 55 (SW or NE). However, these mechanisms are fer events with shallow hypocenters i
only, and the maximum earthquake would probably not occur
1 22
with this orientation. Figure 17A & B shows the strike lines of this displacement superimposed on the core outline.
- c. Displacement on existing joint sets below the dam is a third possibility. The joints under the dam strike N45 E and N45 W and are nearly vertical. These directions a e respectively parallel and perpendicular to the dam axis.
(2) Modes of Displacement The two predominant modes of displacement computed based - on the observed small events were normal and reverse. Normal (or extensional) displacement predominated in the ~ focal plane solutions and was used for calculating the i maximum hypothetical earthquake. Strike slip displacement also was evident but almost always with a significant normal or reverse component. The computed displacement mode for the events directly under the dam is reverse or a compressional with very little strike slip component. i FOUNDATION DISPLACEMENT AND INDUCED EXTENSIONAL STRAINS NEAR CREST A potential for cracking and hydraulic fracturing would be presumed if the effective stress in the core were reduced to zero or a small negative value by foundation movements. Extensional foundation displacements beneath the core could cause the effective stress in the core to decrease. The upper ; i part of a dam has been shown by experience to be sensitive to extensional l r strain, because it is the. thinnest part; moreover, it sustains less internal pressure from its own weight that tends to close cracks. The following paragraphs describe an estimate of how much extensional strain could be tolerated by the upper part of the Jocassee core. 2s
-T
f N 4 it was assumed the initial effective vertical stress in the core is equal to the depth below the crest multipled by the unit weight ofthe core material,
, less the depth below full pond multiplied by the unit weight of water. This l Is conservative because the greater settlement of the shell would induce increased effective vertical stresses exceeding the weight of the soll column.
l Also, the head loss across the core has been ignored. It was assumed that the horizontal effective soll pressure is related to the vertical effective 2 pressure by a coefficient of earth pressure of 0.54 (which assumes elastic _ l l conditions and a Poisson's ratio of 0.35). To estimate the changes in effective stress acting in a horizontal direction, (acting on a transverse - plane through the embankment) caused by horizontal extensional strains near the crest, a condition of plane strain deformation with no changes in either the vertical stress or pore pressure was assumed, and the generalized Hooke's law was applied. Small increments of horizontal strains were selected, and af ter calculating the change in stress, the tangent modulus E for the next in-
; crement of strain was approximated using the relationship:
1
- _ 2
[8 % 3 Rf(aj -o ) (1-s i ne ) 3 E = Kp*I --- --
! (p,) _2c cose + 203 * "#_
i i i The tangent Young's modulus was estimated from the results of the previously discussed record consolidated-undrained triaxial tests on core materials, expressed by the parameters K=350, n=0.4 and R =0.75 f It was assumed that the modulus in unloading, expressed through K , was 1.5 times the modulus for loading. ., 24 i
, . - . +,m,. --_w -_ .--,-~r ,- =
i 3 L ! The results , for an element of soll under overburden stress conditions 4 j compatible with those estimated at elevation 1100 using the above assumptions, ! i ] are shown in Figure 19A. Figure 19A shows the horizontal effective stresses i in the core soils near elevation 1100 (20 ft below the crest of the core and 10 f t below the full pond elevation) decrease at a decreasing rate with exten- ! l sional (stretching) strain. When ex' tensional straining reaches 0.5%. the ! i
- effective horizontal stress has dropped to about 1/5 of its original value.
! The plot suggests rapidly increasing strains with little stress change at 1 about 0.5% extensional strain. At this point, cracking or plastic flow could t
- begin. This confirmed by reported cracking of other dams at about 0.1% to 0.3% strain.
Figure 18 shows two general cases (part B is a special case of part C) of how , foundation displacements could stretch the dam and cause extensional strains. in Figure 18A, foundation displacement having a predominance of strike-slip motion would stretch the dam. If the displacements remain small, the averaged extensional strain can be estimated by the expression shown in the figure. Figure 188 and C show how foundation displacements having a predominance of f i normal (extensional) mode would cause extensional strains due to both bending and stretching the core. A steeply dipping surface of displacement (compress-ional or extensional) would cause primarily bending deformations of the dam i (Figure 188). The expressions for extensional and compressional- strain caused ' by bending only in Figure 188 originate from experience with surface subsidence 3 , effects in the mining industry, in Figure 18C, an expression is given for estimating the effects of combined stretching and bending from normal displace-
- ment on a sloping surface.
.n. 3'
It will be assumed that the near-crest strains caused by the foundation displace-ments are half the average strains computed from the expressions in Figure 10B and C. (This is conservative for small displacements; St.Venants principle suggests that near-crest strains would be virtually zero because heights are orders of magnitude times the foundation displacements). Stri.ke-slip motion, i Figure 18A, will not be considered since it is not important in the focal mechanisms computed for the observed small events at Lake Jocassee. l FOUNDATION DISPLACEMENTS TO CRACK CORE NEAR CREST From examination of Figure 17 and from the previous discussion of directions for presumed foundation displacements, it is concluded that the most appro-priate conditions for analysis of extensional strains are shear foundation displacements on a surface striking approximately perpendicular to the dam axis and (1) steeply dipping (Figure 187), and (2) dipping about 45 from the horizontal (Figure 18C, angle #=45 ). For these two cases, Figure 190 has been prepared. it illustrates the amount of foundation displacement causing near-crest extensional strains of 0.5% for various heights of embankment core. From Figure 198, for the higher part of the dam, (between roughly stations 9+50 and 14+50, where the height above prepared foundation is in the range of 350 to 380 ft), foundation displacements on a 45 dipping plane striking trans-versely to the dam of greater than 34 inches would be required to cause near-crest strains exceeding 0.5 percent. For a steeply dipping plane, such as the N45 W Joint set, the corresponding displacement to cause 0.5% strain is i about 56 inches. 26
1 As the embankment height becomes less above prepared foundation in the sloping abutments, the estimated tolerable extensional displacement in the underlying foundation becomes less, as seen on Figure 198. FOUNDATION DISPLACEMENT AND REMAINING SHEAR STRAIN CAPACITY OF CORE The existing state of strain within the core of the dam at locations where straining to date appears to have been the most is shown on Figure 14. . According to Figure 16 compared with Figure 14, as discussed previously, the core between elevation 1100 and 1000 has more than 5% to 6% remaining out of . a total strain capacity of 7% to 8%, and the core between elevation 900 and 3 1 800 has more than 3% to 4% remaining out of a total strain capacity of 9 3% to 10.3%. Between approximately stations 1+55 and 4+00, the core below elevation 1065 has experienced settlement displacements to date on the order of 0.8 f t to 0.9 ft causing the existing strains shown on Figure 14. The remaining strain capacity is greater than (5 to 6) divided by (7 to 8) or 70 to 75 percent of the total. This infers ultimate settlement displacements on the order of (0.8 ft to 0 9 ft) divided by (.25 to .30) or 2.7 to 3.6 f t will be required to fully develop the strain capacity. This leaves 1.9 to 2.7 ft (23 to 32 in.) ; of settlement displacement to fully develop the strain capacity at the currently most strained locations between elevations 1100 and 1000. These values are very conservative, as was previously discussed. i Between approximately stations 7+50 and 9+50 and between approximately 16+00 and 16+50 the core below elevation 900 has experienced settlement displacements 27
4 . , I to date on the order of 2.6 ft to 2.8 ft causing the existing settle-ment strains shown on Figure 14. The remaining strain capacity is more than (3 to 4) divided by (9.3 to 10.3) or 30 to 40 percent of the total. This infers ultimate settlements on the order of (2.6 ft to 2.8 ft) divided by (.60 to .70) or 3.7 to 4.7 ft to fully develop the strain capacity. This leaves 1.1 to 1.9 ft (13 to 231n.) of estimated additional settlement to fully develop the strain i capacity at the currently most strained 1ccations below elevation 900. Based on the observed settlement rates and even discounting creep, these limits will - not be reached. DISPLACEMENTS. CRACKING AND STRAIN CAPACITY OF ABUTMENT I
; Figure 3 shows the weathering profile for the right and left abutments, based I
on c.ta from more borings and grouting holes. The left (east) abutment is formed of a shallow depth of residual soil over weathered rock, with hard rock begins at about elevation 1080, only 30 ft below full pool elevation and 45 f t below nominal dam crest elevation 1125 The right (west) acutment, however, includes a thicker zone of residual soil and weathered rock. For the right abutment proflie location on Figure 3, the elevation of rock is about 990, 120 f t below normal pool elevation. However, rock outcroppings exist at elevation 1100 or above on the upstream slope of the right abutment. The right abutment profile shown on Figure 3 is located on the downstream slope, where the natural and present ground surface elevation is 1125 to 1130. The now excavated crest of the hill forming the abutment reservoir-rim had a ground surface rising westward from elevation 1125 at the end of the dam to 1150+, and the hill crest was located 75 to 100 ft upstream of the right abutment proflie location on Figure 3 Hence, it is likely that the rock line 28
for the right abutment beyond the end of (west of) the embankment in Figure 3 is somewhat higher in elevation beneath the crest of the hill proper. The water levels measured in the observation wells on the abutment near the end of the dam range from elevation 1105 to 1090 and lower. The changes in effective stress from an assumed initial stress acting in a horizontal direction (normal to a transverse plane through the abutment reservol.r rim) caused by horizontal extension strains could be estimated assuming a condition , of plane strain deformation with no changes in the vertical stress. The modulus parameter, K, previously presented for the weathered rock and hard , residual soll should be multiplied by a correction factor on the order of 1.5 to 2 or more to account for sample disturbance. This correction would assign modulus values (at low strain) for the hard residual soils and weathered rock that, for the range of confining pressures less than 6000 to 8000 psf, are still somewhat lower than those used in similar computations for the core soils. Hence, it was assumed, as a first conservative approximation, that a horizontal i extensional strain limit of 0.5% would apply to an element of residual soil on the reservoir rim where the ground surface is at about elevation 1125. Thus, Figure 198 can be utilized as a conservative evaluation of displacement. to crack the abutments. With about 120 ft of weathered rock and soll overlying hard rock at elevation 1005, shear displacements on a surface slooing at 45 in tha rock of 12 inches would be reqstred to cause extensional streins of 0.5%. For displacements on a steeply dipping surface, such as the predominant joints, the corresponding shear displacement for 0.5% extensional strains would be 19 inches. For 45 f t of soil over rock, such as the lef t abutment, the computed displacements are 4 to 5 in. and about 7 in, for rock displacements on a 45 and vertical plane, respectively. 30
The displacements actually necessary to moblize the shear strain capacity of the weathered rock and hard residual soils of the right abutment should be considerably greater than those stated in the preceding paragraph, since there are no known shear strains of significance already existing beneath the crest of the abutment hill, and since the shear strain capacity of the material is in excess of 10% at low confining pressures, as shown on Figure 15A. Ah0UNT OF HYPOTHETICAL FOUNDATION DISPLACEMENT ASSOCI ATED WITH MAXIMUM EARTH-QUAKE As presented in the Law report of September 1976, the magnitude of the maximum ' earthquake was computed to be 5.6 based on a stress drop of 7 bars and a fault radius (r) of 3.54 Km. The maximum displacement U max., which is 1.5 times d the average displacement, can be computed f rom equations frcm Brune (1971)( ' a r 16 3 d
,g 2 75 2 where a = 7 bars r = 3.54 Km a = 2.67 gm/cm 3 0 =
3.45 Km/sec. The results give 8.6 cm (3.4 inches) of maximum displacement. This is in good agreement with observed results implied by data b shinnery (1969) . The implied average movement is 3.4 divided by ' f, c r 2.3 inches. (3) Brune, James N. (1971) Journal of Geophysical Resed h, 76(20) p. 5002. 5 (4) Chinnery (1969) Bulletin, Seismological Society of America 59(5) 1969-1982. 30
The maximum displacement is much less than the estimated extension movements necessary to cause tension cracking of the upper part of the dam. Therefore, hydraulic fracturing near the top of the core would not be initiated by shear in the foundation. The maximum displacement is also less than the calculated tolerable displacements of the abutments.
~
Shear in the foundation of the order of magnitude hypothesized, 3 to 4 in., would produce shear cracking and displacement of the core adjacent to the core-foundation con tact. However, open cracking is not likely because of the high confining or normal stresses in the core at this depth. Even if an open crack should develop in the core there is a 3 layer filter protecting the core from. erosion. Each layer is 8 ft wide. Hence, even strike slip shear displacement of 3 to 4 in, would not significantly impair f the filter integrity. Therefore we conclude that the dam core will not crack nor initiate seepage erosion failure should shear displacement of 3 to 4 in. occur. l l l 31
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EX! STING SHEAR STRAINS IN CORE L ,.. ....., _ ,, ,. ~ ,- ,.. . .., , , . ~ . . . . . m -- - - ..- FIGURE 14 1 W_
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" 45
- HYPOTHETICAL DISPLACEMENT SURFACES
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- --AMOUNT OF DISPL/CENmT _
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