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UPDATED ASSESSMENT OF CONCRETE GROWTH EFFECTS
ON HIWASSEE DAM
Dan D. Curtis1 Bradley J. Davis
2
Shahzad Rahman3 Rex C. Powell
4
ABSTRACT
Tennessee Valley Authority’s Hiwassee Dam is experiencing concrete growth caused by
Alkali-Aggregate Reaction (AAR). Extensive investigations and remedial actions were
undertaken in the early 1990s. The concrete growth investigations were recently updated
and this paper presents the results of the updated investigations.
Hiwassee Dam is a 307-ft high concrete gravity dam located near Murphy, North
Carolina. The project was completed in 1940 and consists of a non-overflow gravity
dam, a spillway with seven bays, an intake and a two-unit powerhouse.
A review of survey and instrumentation data is presented. The survey and
instrumentation data is processed to provide estimates of concrete growth rates in the
dam. The survey, instrumentation data and measurements from stress cells and concrete
overcoring are used to calibrate a global finite element model of the dam. The finite
element program ANSYS was modified using the user-programmable features to include
the nonlinear stress-dependent nature of concrete growth and enhanced creep behavior of
AAR-affected concrete. The finite element model was calibrated to field measurements
and then used to predict the future conditions in the dam and spillway. The finite element
model was used to assess the behavior of slots, which were cut in the dam in the early
1990s. The finite element model results were also used to assess the effects of concrete
growth on the sliding stability of the dam.
A local finite element model was used to investigate the effects of concrete growth of the
spillway piers. The spillway has seven radial gates and the gate anchorage steel is
embedded in the spillway pier concrete. The analysis showed that the anchorage steel
had become highly stressed due to the concrete growth and additional anchorage was
recommended and installed. A thorough review of instrumentation data confirmed the
conclusions from the finite element modeling.
1 Senior Civil Engineer, Acres International, 4342 Queen Street, P.O. Box 1001, Niagara Falls, Ontario,
Canada, L2E 6W1, Tel: (905) 374-5200, Fax: (905) 374-1157, [email protected] 2 Civil Engineer, Tennessee Valley Authority, 1101 Market Street, LP1F, Chattanooga, Tennessee 37402-
2801, Tel: (423) 751-4844, Fax: (423) 751-3548, [email protected] 3 Civil Engineer, Acres International, 4342 Queen Street, P.O. Box 1001, Niagara Falls, Ontario, Canada,
L2E 6W1, Tel: (905) 374-5200, Fax: (905) 374-1157, [email protected] 4 Senior Civil Engineer – Acres International, 100 Sylvan Parkway, Amherst, New York 14228-1146, Tel:
716-689-3737 • Fax: 716-689-3749, [email protected]
INTRODUCTION
Project Description
Hiwassee Dam is located on the Hiwassee River in Cherokee County, North Carolina.
Construction of Hiwassee Dam started in July 1936 and was completed in December
1940. The dam is a 307-ft high concrete gravity structure. The total length of the dam is
1285 feet and is composed of a 260-ft long central spillway section, and two non-
overflow sections on the left and right side of the spillway, measuring 597 and 428 ft,
respectively. The main spillway consists of seven radial gated bays. The powerplant has
two units, one of which is a pump-turbine. There is a roadway over the top of the dam
with bridge spans at the spillway bays. The dam is owned and operated by the Tennessee
Valley Authority.
The concrete structures at Hiwassee Dam are affected by Alkali-Aggregate Reaction
(AAR), which causes the concrete to expand over time. The AAR-induced deformations
can have adverse effects on the functioning of mechanical equipment such as spillway
gates and turbines. At Hiwassee, the AAR-induced stresses have caused cracking at
several locations, most noticeably along the horizontal construction joints at both ends of
the dam. The non-overflow blocks adjacent to the spillway have deflected into the
spillway opening, causing the spillway gates to bind. There was also concern that these
deflections may eventually overload the spillway bridge. As a result, TVA has in place an
extensive program consisting of monitoring by instrumentation, finite element based
investigations, and a slot cutting program aimed at extending the service life of the
project.
The objective of this paper is to present an updated assessment of the effects of concrete
growth due to AAR at Hiwassee Dam.
Previous AAR Assessments and Remedial Actions
In 1991 a three-dimensional, linear-elastic finite element model of the dam and its
foundation was developed by Stone & Webster Engineering Corporation using ANSYS, a
general-purpose finite element analysis program developed by ANSYS Inc. In these
analyses AAR-induced concrete growth was simulated by an equivalent thermal
expansion of concrete. The model was used as an aid for devising remedial measures to
control stresses and deflections resulting from concrete growth. From the analysis, it was
concluded that remediation would include the following: post-tensioning of the dam,
cutting of expansion joints in the bridge, and cutting of four transverse slots; two adjacent
to the spillway and two near the abutments. This would help control the stresses and
allow additional expansion of the concrete in the spillway area without causing binding
of the spillway gates. Figure 1 shows a downstream elevation of the dam and the location
of the four slots in the dam.
The remedial repairs were completed using a phased approach. The non-overflow
sections of the dam were post-tensioned to an average depth of 75 feet, using two rows of
multi-strand tendons. This was included in the repair program, primarily to improve
stability under seismic loading and during slot cutting. The spillway section was not
post-tensioned.
Figure 1. Downstream Elevation Showing Four Slots in Upper Portion of Dam
In 2002, Acres International was contracted to provide an updated assessment of the
condition of the dam. The updated assessment included: a review of instrumentation
data, global finite element analysis (FEA) of concrete growth effects, stability
assessment, local FEA of spillway piers and recommendations for remediation. The
following sections summarize the results of these updated assessments undertaken by
Acres.
REVIEW OF SURVEY AND INSTRUMENTATION DATA
Survey and instrumentation data includes the following:
• top of dam settlement and alignment surveys
• Electronic Distance Measurement (EDM) surveys
• spillway pier stationing surveys measuring tilt of end piers
• upstream-downstream tilt of spillway Block 14 plumb line
• growth meters and extensometers in Block 16 of the left non-overflow section
• joint meters at slots and adjacent block joints
• stress measurements from overcoring and stress cells
• temperature measurements from various instruments in the dam.
The top of dam survey data was processed in two ways. The vertical rise measurements
from the special survey of 1991 were used to estimate the total rise of the dam since the
end of construction (1940). The top of dam surveys after 1990 were used to estimate
rates of vertical rise using the 1990 survey as a baseline. The survey measurements
provided very useful information on the average rate of concrete expansion over the
height of the dam. The various surveys were quite consistent with respect to the rates of
concrete growth in the vertical direction.
The EDM surveys measure the longitudinal length changes of the right non-overflow
section, the spillway and the left non-overflow section. The measurements are taken at
the top of the dam. The measurements from 1975 onward provide useful information on
the change in the width of the spillway as concrete growth causes the spillway end piers
to tilt inwards.
The Block 14 plumb line provides upstream-downstream tilt measurements of spillway
Block 14. The plumb line was operational prior to reservoir impoundment in 1940,
therefore it provides useful data on the dam response to impounding and also the tilt in
the upstream direction caused by continued concrete growth.
Growth meters and extensometers were installed in Block 16 adjacent to the spillway.
The meters were installed at the top of the dam and in an access gallery near the base of
the dam. These instruments were installed in 1991; hence this new data was not
available in the previous evaluations. The measurements, together with the survey data,
were used to estimate the distribution of vertical concrete growth as shown in Figure 2.
Figure 2. Extensometers and Growth Meters in Block 16
Joint meters are installed across the two spillway and the two abutment slots. The joint
meters monitor the closure of the slots over time.
Concrete overcoring stress measurements provide data on the stress condition in the dam
at key points in time. In several cases, stress cells were installed in the overcored holes to
allow monitoring of the increase/decrease of concrete stress over time. The overcoring
stress measurements showed that relatively large longitudinal compressive stresses occur
at the crest of the spillway near the spillway end piers. The large longitudinal
compressive stress in this area is attributed to stress flow from the upper portion of the
non-overflow section down into the crest of the spillway with the spillway acting as a
wide notch in the dam. Several stress cells were located adjacent to and below the four
slots. The stress cells adjacent to the slots showed an increase in stress during slot cutting
followed by a complete unloading as the diamond wire saw proceeded deeper below the
stress cell. The stress cells under the slots showed an increase in stress due to
longitudinal flow of stress under the slot.
GLOBAL FINITE ELEMENT ANALYSIS
TVA supplied Acres with the original finite element (FE) model as used in the 1991
studies. The finite element model is shown in Figure 3. The finite element model is
relatively coarse but it was considered sufficiently detailed to allow an assessment of the
condition of the dam. The finite element program ANSYS was used to perform the time-
dependent nonlinear concrete growth analyses. ANSYS has a feature of allowing user-
specified material subroutines whereby the user can program their own material laws into
ANSYS. This feature was used to implement the time-dependent and stress-dependent
concrete growth law from Acres GROW3D program into ANSYS. The concrete growth
law is described in Curtis (1995). The solution procedure is shown in Figure 4.
Figure 3. Global Finite Element Model Mesh
Figure 4. Concrete Growth Law Solution Procedure in ANSYS
DAM AND FOUNDATION
6881 Elements 9301 Nodes
Establish Initial Conditions
(Gravity, Hydrostatic Loads etc)
Access Element Stresses
Compute Strain / Stress Increments in
Global Coordinates Based on Growth Law and Impose Them at Each Gauss Pt.
Carry out Equilibrium Iterations till Convergence
Save Results for Load Step
Next Load Step
The finite element model of the dam was calibrated using concrete growth rates
determined from data review. Temperature measurements at instruments within the dam
were used to assess the effect of temperature on the concrete growth rates. Similar to
previous work, it was found that the higher temperature on the downstream portion of the
dam leads to higher growth rates in this area.
From Figure 4, the initial conditions are established prior to running the nonlinear
concrete growth analysis. The model is stepped through time using one-year time
increments to determine the effects of continued concrete growth. The finite element
model was calibrated with vertical displacement survey data as shown in Figures 5 and 6.
Figure 7 shows the calibration to the Block 14 plumb line data. Figure 8 shows the
comparison of the measured and computed spillway closure from the EDM survey.
Figure 9 shows a comparison of measured and computed longitudinal stresses from
Block 16 adjacent to the spillway. A good match to the measured stresses was achieved.
The slot closure measurements at Block 16/17, shown in Figure 10, also compare well
with the closures predicted by the model.
The model results taken prior to slot cutting computed quite large horizontal shear
stresses acting longitudinally at the spillway end pier/spillway crest junction. The fact
that the model matches the measured longitudinal stresses and slot closures at this
location suggests that other stress components are probably reasonably well estimated by
the model. There was no evidence of Block 16 sliding longitudinally into the spillway
opening, therefore the construction joints at this location must possess quite high
cohesive strengths in excess of 400 psi!
Conclusions From Global FEA
The model has provided useful insight into the behavior of the dam and this information
was compared to other projects to assess rehabilitation requirements and life extension
planning. The calibrated model was used to investigate the various proposed remedial
measures, such as deepening of the slots, additional spillway slots, and the effect of
allowing the spillway slots to close.
Investigation of the mechanisms of crack formation revealed that the cracking in the
curved abutment blocks is due to the interaction of the longitudinal compression and the
curved geometry of the dam at these locations. The abutment slots prevent the
development of compressive force in the curved abutment blocks, which should prevent
future crack formation in these blocks. Therefore the abutment slots should be
maintained open.
The stability of the dam sections most affected by concrete growth and slot cutting was
examined. Forces acting on sections adjacent to the slots were determined from the FE
model. In some cases, it was found that blocks adjacent to slots may, in the future, move
slightly towards the slot to relieve longitudinal forces caused by continued concrete
growth. Stability in the upstream-downstream direction of the blocks in the vicinity of
Figure 5. Measured versus Computed Total Vertical Rise
Figure 6. Measured versus Computed Vertical Rise Since 1990
.
DISTANCE FROM SURVEY MARKER 3-1 (FT)
Settlment Surveys After 1990
-0.5
0.0
0.5
1.0
1.5
2.0
2.5
3.0
3.5
4.0
-200
-100 0
100
200
300
400
500
600
700
800
900
1000
1100
DISTANCE FROM SURVEY MARKER 3-1 (feet)
SE
TT
LE
ME
NT
(-)
---
Vert
ical
Mo
vem
en
t (i
nch
) --
- E
XP
AN
SIO
N (
+)
Special Survey
of 1991
Calibration
Result
Vertical Movement (1940 to 1991)
Right Abutment Spillway Left Abutment
-0.60
-0.50
-0.40
-0.30
-0.20
-0.10
0.00
0.10
0.20
0.30
0.40
0.50
0.60
0.70
0.80
0.90
1.00
1.10
1.20
-200
-100 0
100
200
300
400
500
600
700
800
900
1000
1100
SE
TT
LE
ME
NT
(-)
---
Mo
ve
men
t (in
ch
) --
- E
XP
AN
SIO
N (
+)
NOV 1990 - INITIAL SURVEY
Dec 2002
Nov 1995
Calibration
Results
Right Abutment Left AbutmentSpillway
1240
1260
1280
1300
1320
1340
1360
1380
1400
1420
1440
1460
1480
1500
1520
1540
-0.6
0
-0.4
0
-0.2
0
0.0
0
0.2
0
0.4
0
0.6
0
0.8
0
1.0
0
1.2
0
1.4
0
EL
EV
AT
ION
(f
eet)
EL 1316
EL 1277
EL 1423
EL 1503
1940
Refe
ren
ce
May 27, 1944
(observed)
Foundation, EL 1245 +/,
assumed to be zero.
1992 (Mean)
Downstream Upstream (inches)
Calibratio
n
Calib
ratio
n
Figure 7. Measured versus Computed Spillway Tilt due to Impounding (1944)
and Concrete Growth (1992)
Figure 8. Measured versus Computed Longitudinal Length Change
of Spillway from EDM
1540
Spillway Block 14 Plumbline Tilt
-1.10
-1.00
-0.90
-0.80
-0.70
-0.60
-0.50
-0.40
-0.30
-0.20
-0.10
0.00
0.10
1972
1973
1974
1975
1976
1977
1978
1979
1980
1981
1982
1983
1984
1985
1986
1987
1988
1989
1990
1991
1992
1993
1994
1995
1996
1997
1998
1999
2000
2001
2002
2003
2004
2005
2006
2007
2008
2009
2010
2011
2012
2013
2014
2015
2016
2017
SP
ILL
WA
Y C
LO
SU
RE
(In
ch
es
)
EDM Survey Data
Calibration
Spwy Slots cut in 1993.
2/
3
S
LO
T
H D - 1H D - 2H D - 4 H D - 3
7/8
SL
OT
16
/1
7
S
LO
T
24
/2
5
S
LO
T
BL
K
0
BL
K
8
BL
K
16
BL
K
2
7
2/
3
S
LO
T
H D - 1H D - 2H D - 4 H D - 3 H D - 1H D - 2H D - 4 H D - 3
7/8
SL
OT
16
/1
7
S
LO
T
24
/2
5
S
LO
T
BL
K
0
BL
K
8
BL
K
16
BL
K
2
7
Right Side Left SideSpillway
YEAR
Abutment Slots cut in 1994.
0.34 " (Immediate Recovery)
0.37 " (Additional
Recovery by 2003)
Longitudinal Stress Hole 3, Phase VI (1992)
Figure 9. Measured versus Computed Longitudinal Stresses in Block 16
Adjacent to Spillway
Closure of Spillway Slot 16/17
Figure 10. Measured versus Computed Slot Closure at Block 16/17 Slot
0
5
10
15
20
25
30
35
40
45
50
-500 0 500 1000 1500 2000
Tension (-) --- STRESS (psi) --- Compression (+)
Dep
th o
f Ov
erc
ore
Te
stin
g (fe
et)
Longitudinal Stress (RAW)
Calibration (AAR + Temp)
Longitudinal Stress (AVG)
Spillway Crest, EL
1503.5
Roadway, EL 1537.5
End of beam
touching
concrete
BAY 7
EN
D P
IER
2
-1.6
-1.4
-1.2
-1.0
-0.8
-0.6
-0.4
-0.2
0.0
0.2
1992 1993 1994 1995 1996 1997 1998 1999 2000 2001 2002 2003 2004
Year
CL
OS
E <
Mo
vem
en
t (i
nch
) >
OP
EN
JM 16/17 US-L
JM 16/17 DS-L
JM 16/17 DS-L (Gallery)
Calibration
Slo
t C
ut
Re
cu
t
Re
cu
t
Re
cu
t
the slots is adequate even if longitudinal movement into the slot has occurred. No
additional anchors to tie the upper portion of the dam to the dam concrete beneath the
slots are considered necessary at this stage.
A number of remedial measures such as deepening of existing slots and allowing the slots
to close were examined using the FE model with concrete growth. The need for
additional spillway slots was also reviewed. Results indicated that there is little
justification for deepening the slots or providing additional slots. Both instrumentation
data and model predictions suggest that the time interval between recutting of the
spillway slots can be increased considerably without impairing gate operation.
The strategy for slot cutting depends on the performance of the spillway gates. At
present, the gates operate satisfactorily without jamming. The survey data and the model
prediction indicate slight continued opening of the spillway width. However, given that
there is tilting of the end piers and the gate clearances are unknown at this time, the
spillway pier movements should be monitored and, if necessary, future measures such as
minor deepening of spillway slots could be undertaken if gate jamming becomes a
potential problem. Such measures would only be pursued after monitoring movements
over the next few years of operation.
LOCAL SPILLWAY PIER ANALYSIS
Regular inspections of the dam have identified cracking in the spillway piers near the
trunnion pins. The cracks are typically vertical and run above and below the concrete
around the trunnion pins. The spillway piers were not included in the global FE model.
In order to investigate the condition of the spillway pier, a local finite element model of
one spillway pier, including the trunnion anchorage steel, was developed. The model of a
typical pier is shown in Figure 11 and the model of trunnion pin and anchorage steel is
shown in Figure 12. The steel reinforcement in the pier was modeled with shell elements
using a smeared equivalent area of steel.
From the analysis, the anchorage steel restrains the concrete and the restrained concrete
expansion leads to higher horizontal concrete compressive stresses in the vicinity of the
anchorage steel. The concrete above and below the trunnion pin anchorage steel expands
horizontally at a greater rate than the concrete in the immediate vicinity of the trunnion
anchorage steel. The result of this differential horizontal expansion is the vertical crack
observed in the concrete downstream of the trunnion pin.
The leveling surveys show the total rise of the dam at the spillway section to be less than
the adjacent non-overflow section. The local finite element analysis also predicts lower
expansion rates in the spillway pier concrete due to the restraining compressive stress
imposed on the concrete from reinforcing steel. As the concrete expands, the reinforcing
steel resists the expansion, and this small restraint induces concrete compression. The
magnitude of this additional compressive stress is 100 to 200 psi.
Figure 11. Local Spillway Pier Model Mesh
Figure 12. Trunnion Pin and Anchorage Steel Plate Mesh (Concrete not shown)
A variety of modelling assumptions were made regarding anchorage connectivity to
concrete and concrete growth rates in the secondary concrete. Regardless of the
assumption used, the model predicts very high stresses in the anchorage steel. The
computed stresses are approaching the yield stress of the steel in both the bracket plate
and double plate anchorage steel. From the analysis, it was concluded that the anchorage
steel is probably yielded at the bolted connection upstream of the trunnion pin.
Subsequently, the trunnion pin extensometer data was re-examined. The measurements
are plotted as a function of time in Figure 13. The data was re-plotted in Figure 14 such
that trunnion pin movements versus reservoir elevation are shown. From the analysis and
measurements, the gate anchorage system is behaving inelastically at high reservoir
levels.
Figure 13. Measurements of Trunnion Pin Movement versus Time
Figure 14. Measurements of Trunnion Pin Movement versus Head Water Elevation
It is anticipated that inelastic behavior is occurring at the bolted connection of the
trunnion pin bracket plate to the double plate anchorage steel. As local yielding develops
at this location, the surrounding concrete takes on additional load which limits the
downstream movement of the trunnion pin. In time, as the steel stresses become higher,
the pin deflections for a given load will increase and the ability of the pier to sustain
additional load will be limited by the shear strength of concrete downstream of the
trunnion pin. The concrete downstream of the trunnion pin was not designed to carry this
Hiwassee Project - BAY 5 - Left & Right Trunnion Pins
0.000
0.005
0.010
0.015
0.020
0.025
0.030
0.035
0.040
2001 2002 2003 2004 2005 2006
US
(-)
---
Pin
Mo
ve
me
nt
(in
ch) -
-- D
S (
+)
1440
1450
1460
1470
1480
1490
1500
1510
1520
1530
1540
He
ad
wa
ter
Ele
vatio
n (
fee
t)
LEFT PIN (Corrected) RIGHT PIN (Corrected) HW
The meter is attached to the pin
and anchor ed in the concrete
ups tr eam.
Trunnion Pin Movement versus Head Water Elevation for Period of May 1
to May 31 , 2003
1512
1514
1516
1518
1520
1522
1524
1526
1528
0.005 0.010 0.015 0.020 0.025
Pin Movement (inch)
He
ad
Wa
ter
Ele
va
tion
(ft)
Left Pin Right Pin
additional load. Also, it is not accepted practice to rely on the downstream pier concrete
alone to carry the gate load. Therefore remedial action was recommended to maintain the
safety and reliability of the dam.
The geometry of the end piers differs from that of the intermediate piers and provides
significantly more concrete downstream of the trunnion pin. It was determined that the
concrete and reinforcement downstream of the trunnion pin for the end piers has surplus
capacity to resist the gate loads. Even though the finite element analysis estimated that
the reinforcement in the end piers has already yielded due to concrete growth, the
reinforcement still provides confinement and strength. The end piers therefore do not
require remedial repairs. Nevertheless, the end piers will be monitored along with the
intermediate piers to track their behavior in the future.
SPILLWAY PIER REHABILITATION
A remediation scheme was needed to allow repair of the intermediate spillway piers
during service. It was also imperative that the reservoir be maintained to a predetermined
elevation in order to provide a margin of safety against inelastic movement of the
trunnion pins. Therefore, the selected scheme needed to be reasonably simple and
require as short an installation time as possible. The physical dimensions and
accessibility of the existing piers also limited available options.
Six schemes were considered in the evaluation of repair options. The selected scheme
was comprised of a pair of post-tensioned anchors and a cast-in-place concrete bearing
pad at each of the six intermediate piers. The anchors were aligned parallel with the face
of the piers and inclined so that the upstream ends could be grouted into the mass
concrete of the spillway as shown in Figure 15.
Figure 15. Spillway Pier Anchor Repair
Design of Anchors
The existing anchorage was designed to restrain the full gate load without relying on the
strength of the concrete downstream of the trunnions. Seismic, flood, and normal
maximum headwater loading conditions were considered for design. Normal maximum
headwater was the governing load case, with a horizontal load of 295 kips at each end of
each gate. Based on an anchor inclination of 25 degrees from horizontal, the design post-
tension load in the new anchors was required to be 325 kips.
As the concrete in the piers will continue to grow after installation of the anchors, the
anchors will stretch and increase the post-tension force over time. The finite element
analysis estimated the annual strain rate due to AAR to be on the order of 14 microstrain
per year, corresponding to an increase in anchor force of about 2.1 kips per year. The
design accommodates this phenomenon by essentially over-sizing the anchors and
initially locking them off at a stress well below the normal allowable. Nevertheless,
anchors will require periodic relaxation over the life of the structure.
While the intermediate piers are normally subjected to nearly equal loading from the
gates on each side, a potential situation can occur where one gate may be fully open or
one gate may be dewatered for maintenance. In such a case, eccentric loading will be
applied to the pier. The anchor loads were estimated to increase by approximately
48 kips under such loading due to prying action at the trunnion pin. This allowance for
prying was accounted for in the design recommendations for future relaxation of the
anchors.
According to the guidelines established by the Post Tensioning Institute (PTI), the design
load in an anchor should not exceed 60% of its guaranteed ultimate tensile strength
(GUTS). The anchors used in the repairs each have a GUTS of 778 kips. In order to
verify timing for future relaxation of the anchors, provisions were made to install load
cells to monitor the anchor loads. The anchors will be monitored for increased loads due
to continued concrete growth and when they approach a load of 419 kips, they will be
relaxed to the initial lock-off load.
The vertical contraction joints within the spillway section of Hiwassee Dam are located
coplanar with one face of most piers. Nearly half the anchors are therefore located in
close proximity to those existing contraction joints. However, due to concrete growth,
the contraction joints are tightly closed and in compression, estimated to be on the order
of 600 psi at the spillway crest. Free-edge effects at the contraction joints were therefore
not considered in the design. In fact, future longitudinal stress increase in the spillway is
expected to improve available bond zone capacity over time.
Installation Access to the work area was accomplished from either a man basket or drill platform
suspended by a crane from the bridge deck over the spillway. The drill operation is
shown in Figure 16.
Figure 16. Anchor Hole Drilling
The existing semi-circular downstream face of the piers did not provide a flat surface
necessary for applying anchors loads. Also, as a result of the AAR, the outer 4 inches of
the pier were heavily cracked and deteriorated. The existing concrete was removed to an
approximate depth of 16 inches to allow construction of a reinforced concrete bearing
pad, or saddle, and to provide a substrate of sound concrete. The new saddle was
constructed to transfer the anchor load into the pier over a sufficient area and to provide a
bearing surface perpendicular to the axis of the anchors. Lateral ties were also installed
through the sides of the pier to resist bursting forces near the anchor heads.
After construction of the concrete saddles, two 6-in diameter holes for the anchors were
drilled into each pier. The holes were approximately 38 feet long, with the bond zone
extending approximately 16 feet into the spillway concrete mass. The bond zone at the
bottom of the hole was water pressure tested and verified to be suitably watertight. The
upper portion, along the free-stressing length, was water tested using only static head in
the hole. Two of the piers exhibited significant leakage from cracks through the side of
the pier near the nose. In order to provide an adequate substrate for the anchor loads and
ensure that the encapsulation grout would not be lost, the cracked concrete was removed
and replaced with new concrete. The new concrete was tied to the existing pier with
grouted steel reinforcement dowels.
The anchors were 2 1/2-inch diameter by 40-ft long ASTM A722 (150 ksi) epoxy-coated
bars with Class I corrosion protection. They were fabricated in the factory with pre-
grouted corrugated poly tube along the bond zone. The anchors are re-stressable, with
the free-stressing length comprised of a grease-filled smooth sleeve factory-sealed to the
top of the corrugated tubing. The entire assembly was encapsulated in non-shrink cement
grout by partially filling the hole and inserting the anchor such that grout was displaced
out the top of the hole. Centralizers were used to ensure uniform encapsulation. The
bearing plate trumpet was sealed to the top of the smooth sleeve with an o-ring. The
remaining annulus outside the trumpet was injected with grout after the bearing plates
were installed. Following anchor installation and tensioning, corrosion protection was
completed with a pre-fabricated grease-filled PVC cap installed over the exposed end of
the anchor bar and sealed to the top of the bearing plate. All anchors were either proof or
performance tested in accordance with PTI guidelines and then locked off at a load
nominally 15% higher than the minimum design load of 325 kips. During the testing of
the first two anchors, the piers and adjacent spillway blocks were monitored with
instrumentation for any adverse movements. No adverse affects were noted and
monitoring was discontinued. The anchored piers are shown in Figure 17.
CONCLUSIONS
This paper has shown the importance of both instrumentation and analysis in the
assessment of concrete dams affected by AAR. Both the instrumentation data and the FE
analysis identified a structural problem with the spillway gate trunnion pin anchorage.
The method of analysing the instrumentation data also proved key in recognizing the
inelastic behavior of the trunnion pin anchorage. In the global analysis of the dam, the
instrumentation data was extremely useful in model calibration and understanding the
present and future behavior of the dam
Figure 17. Completed Spillway Pier Repair
ACKNOWLEDGEMENTS
The authors wish to thank the Tennessee Valley Authority for permission to publish this
paper. Numerous TVA staff were involved in this project including M. Cones, R. James
and J. Peyton. The bulk of the instrumentation data presented in this paper was collected
and prepared by D. Tanner of TVA. Also, F. Feng of Acres participated in the anchor
design and analysis.
REFERENCES
Acres International for Tennessee Valley Authority. “Hiwassee Dam Spillway Pier
Repair – Design Basis”. Document No. 14520.29.09. May 20, 2004
Acres International for Tennessee Valley Authority. “Hiwassee Dam – Spillway Pier
Repair Design Calculations”. Document No. 14520.37.03.04. August 23, 2004.
Acres International for Tennessee Valley Authority. “Stress Analysis of Hiwassee
Spillway Pier”. Document No. P14520.29.07. June 2004.
Curtis, D.D. “Modeling of AAR Affected Structures Using the GROW3D FEA
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Tanner, D. T. “The Use of Monitoring and Finite Element Analysis in Evaluating
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