NBSIR 78-1578

INVESTIGATION OF CONSTRUCTION

FAILURE OF REINFORCED CONCRETE

COOLING TOWER AT WILLOW ISLAND,

WEST VIRGINIA

by

H. S. Lew

S. G. Fattal

J. R. Shaver

T. A. Reinhold

B. J. Hunt

Center for Building Technology

National Engineering Laboratory

National Bureau of Standards

Washington, D.C. 20234

November 1979

Sponsored by

Occupational Safety and Health Administration

U.S. Department of Labor

Washington, D.C. 20001

U.S. DEPARTMENT OF COMMERCE

Luther H. Hodges, Jr., Under Secretary

Jordan J. Baruch, Assistant Secretary for Science and Technology

NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Director

ABSTRACT

The collapse of the natural-draft hyperbolic concrete cooling tower

unit no. 2 at the Pleasants Power Station at Willow Island, West Virginia

was investigated by the National Bureau of Standards. The investigation

included onsite inspections, laboratory tests of construction assembly

components and concrete specimens, and analytical studies.

Based on the results of these field, laboratory and analytical investigations, it was concluded that the most probable cause of the collapse

was due to the imposition of construction loads on the shell before the

concrete of lift 28 had gained adequate strength to support these loads.

The analysis of the shell indicated that the collapse initiated at the

part of the shell in lift 28 where cathead no. 4 was located. It further

showed that calculated stress resultants at several points in that part

equaled or exceeded the strength of the shell in compression, bending and

and shear. The failure of these points in that part of the shell would

have propagated to cause the collapse of the entire lift 28.

Key Words: Collapse; concrete; concrete strength; construction;

cooling tower; failure; hyperbolic shell; shell.

ii
PREFACE

An investigation of the construction failure of the natural-draft

concrete cooling tower at Willow Island, West Virginia on April 27, 1978

was carried out by the National Bureau of Standards (NBS) at the request

of the Occupational Safety and Health Administration (OSHA), the Department of Labor. This investigation was conducted pursuant to an interagency

agreement between OSHA and NBS.

The NBS field investigation team consisted of E. 0. Pfrang, J. 0.

Bryson, E. Anderson, S. G. Fattal, B. J. Hunt and H. S. Lew. Throughout

the course of this investigation, the team received full cooperation from

the OSHA regional and area offices. Assistance provided by Mr. David H.

Rhone, Regional Administrator and Mr. Stanley Elliot, Area Director, is

gratefully acknowledged by the NBS team.

l
TABLE OF CONTENTS

Page

PREFACE i

ABSTRACT ii

1. INTRODUCTION . 1

1.1 Background 1 1.2 Objective and Scope 2 1.3 Organization of the Report ...... 2

2. COOLING TOWER CONSTRUCTION . 4

2.1 Introduction 4

2.2 Physical Description of Cooling Tower Unit No. 2

at Willow Island 4

2.3 Construction Method 5 2.4 Hoisting System 10

3. SITE INVESTIGATION 12

3.1 Introduction 12

3.2 Investigation at the Site 12

3.3 Interview Statements of Workers 17

4. LABORATORY INVESTIGATIONS 20

4.1 Introduction 20

4.2 Concrete Tests 20

4.2.1 Examination of Constituents of the

Concrete 21

4.2.2 Determination of Strength and Stiffness

Concrete ..... 24

4.3 Anchor Bolt Tests ..... 28

4.4 Hoist Cable Tests 29

4.5 Chain Hoist Test 30

4.6 Grip-Hoist Test 30

5. EXAMINATION OF CONDITIONS OF CONSTRUCTION ASSEMBLIES AT THE

TIME OF THE COLLAPSE 32

5.1 Introduction 32

5.2 Examination of Hoisting System 32

5.3 Examination of Scaffolding System 37

5.4 Summary 37

iii
TABLE OF CONTENTS (Cont.)

Page

6. DEFINITION OF LOADS AND SHELL ANALYSIS 38

6.1 Introduction 38

6.2 Geometry 41

6.3 Mechanism of Load Transfer 44

6.4 Weight of Scaffolding - Load Case 2 47

6.5 Live Loads - Load Case 3 48

6.6 Weight of Cathead Gantry Assembly - Load Case 4 48

6.7 Hoisting Loads - Load Case 5 50

6.8 Dynamic Effects 56

6.9 Shell Analysis 58

7. INTERPRETATION OF SHELL ANALYSIS 63

7.1 Introduction 63

7.2 Strength of Shell Under Combined Axial Load and

Bending Moment 63

7.3 Strength of Shell Under Shear ..... 68

7.4 Probable Modes of Failure 70

8. SUMMARY AND CONCLUSIONS 71

9. ACKNOWLEDGEMENTS 73

10. REFERENCES 74

TABLES AND FIGURES 76

APPENDICES

A. Patent of Wall Construction System 152

B. Computer Programs for Shell Analysis .... 163

C. Results of Chemical Analysis of Cement 184

D. Results of Analysis of Concrete 187

iv
1. INTRODUCTION

1.1 Background

Two natural-draft hyperbolic concrete cooling towers were being

constructed at the Pleasants Power Station which is located on the Ohio

River at Willow Island, West Virginia. The shell of the first unit of

the two towers was completed in August, 1977. It has a base diameter

of 358 ft (109 m) and stands 430 ft (131 m) above the ground level.

Shortly after 10 a.m. on April 27, 1978, the top portion of the second

unit which had reached a height of 166 ft (51 m) collapsed during construction. Figure 1.1 is a photograph taken after the failure which

shows the completed tower unit No. 1 and partially completed tower unit

No. 2 in the foreground. A four-level scaffolding system which was

anchored to the collapsed portion of the shell fell with it killing all

51 workers who were on the scaffold.

The shell was constructed through the use of a patented lift form tech

nique. Except for the lower and upper portions of the tower, the construction procedure at Willow Island utilized a scheme to place a 5-ft (1.5-m)

lift per day. At the time of failure, 28 lifts had been completed with the

most recent one having been placed the previous day. The formwork which

supported the less than one day old concrete of lift 28 had been raised

into place for lift 29. According to eyewitness accounts by workers,

lift 28 began to collapse when the third bucket of concrete was being

hoisted up to the working platform. It was estimated that about 1.0 yd

O (0.8 m ) of concrete had been placed for lift 29 at that time. According

to an eyewitness the entire section of lift 28 collapsed into the tower

within a few minutes.

1
On the day of the collapse, an inspection team from the Occupational

Safety and Health Administration (OSHA), Department of Labor, arrived at

the site and began an investigation into the collapse. A team from the

National Bureau of Standards (NBS) joined the OSHA team two days later.

1 . 2 Objective and Scope of the Investigation

The National Bureau of Standards was requested to assist the field

investigation conducted by the Occupational Safety and Health Administration at the site of the collapse and to carry out a detailed study aiming

at the determination of the most probable cause of the collapse.

In response to this request, NBS carried out field, laboratory

and analytical studies. The NBS investigators also used data obtained

from onsite inspections, OSHA case records, the patent for the forming

system and drawings showing details of the tower construction.

1 . 3 Organization of the Report

The report is organized in ten chapters:

Chapter 2 presents dimensions of the tower and the material used for

construction. It also describes the construction method employed for

erection of the tower and the hoisting system used to transport construction materials from the ground level to the top of the tower.

Chapter 3 describes observations made and measurements taken by the

NBS investigation team at the Willow Island site. It also describes

typical daily construction activities and the chain of events which took

place prior to the collapse.

Chapter 4 presents the results of concrete tests which were used

to determine the strength and stiffness gaining characteristics of the

2
concrete. The test results of anchor bolts, hoist cable, chain hoist

and grip-hoist are also presented.

Chapter 5 examines possible component failures of the hoisting and

scaffolding systems with the aid of the test results presented in

chapter A as well as data obtained from the onsite inspection presented

in chapter 3.

Chapter 6 describes criteria and rationale for defining the loads

which were acting on the shell at the time of the collapse. It also

presents analysis of the shell by means of finite element computer

programs.

Chapter 7 compares the results of the shell analysis with various

strength parameters of the shell and discusses the most probable mode

of failure.

Chapter 8 summarizes the findings of the investigation and presents

conclusions drawn on the based of field, laboratory and analytical studie

Chapter 9 acknowledges those individuals who made contributions to

various phases of the investigation and to the preparation of this report

Chapter 10 lists the references cited in the text.

3
2. COOLING TOWER CONSTRUCTION

2.1 Introduction

This chapter describes dimensions of the tower and the materials used

for construction. It also describes the method of construction and the

hoisting system. The formwork and scaffolding systems used for construction of the tower were supported only by the previously completed portion

of the tower. They moved up as construction progressed with no support

other than from the partially completed structure. Because such complex

self-lifting formwork and scaffolding systems were used, they are described

in some detail.

2 . 2 Physical Description of Cooling Tower Unit No. 2 at Willow Island

Figure 2.1 shows the elevation and plan view of cooling tower unit

no. 2. It may be noted that both the diameter and the shell thickness

changed along the height of the tower. At the time of failure, the tower

elevation had reached 166 feet (51 m) as indicated by the dotted line

on the elevation. The shell portion of the tower was supported by 80

diagonal concrete columns of 34 inches (0.86 m) in diameter. The shell

tapered from the lintel at the base (diameter of 342 ft, 104.2 m) to the

throat (diameter of 214 ft, 85.2 m). It then flared out with increasing

height. The exterior of the shell was divided by 96 evenly spaced ribs

(see figures 1.1 and 3.1). The panel width between two adjacent ribs

varied along the height of the tower.

In the wall section, two layers of reinforcement were provided both

vertically and horizontally. The size and spacing of steel in both

directions varied with the height of the tower. A typical cross section

of the wall at lift 28 is shown along with reinforcing and splicing

4
details in figure 2.2. The figure shows that, within any one lift height,

all of the vertical bars were spliced in every third panel. The horizontal bars were spliced such that no two splices in any one lift occurred in

the same vertical plane for bars on the same face. The spacing of

splices along the horizontal bars were not specified in the drawings.

All reinforcing bars were specified to meet ASTM A 615 Grade 60 requirements.

Normal weight concrete was used for the tower. The mix proportions

per cubic yard of concrete were:^

470 lbs (214 kg) Type II (moderate heat of hydration)

Portland cement

1174 lbs (532 kg) Natural sand

1900 lbs (862 kg) Natural gravel

61 lbs (27 kg) Fly ash

14.1 oz (417 mL) Water reducing agent

5.0 oz (148 mL) Air entraining agent

Approximately 27.6 gal (104 L) of water were specified to produce 4.5 in

(114 mm) slump.

2 . 3 Construction Method

As mentioned in section 2.1, both the thickness and the diameter

of the shell changed along its height. To accommodate these changes

as well as to provide working surfaces for workers, a complex combination of formwork and scaffolding system was used. A schematic drawing

of the formwork-scaffolding system is shown in figure 2.3. It is seen

in this figure that the entire system was supported on previously

completed portions of the tower. As construction advanced, the system

traveled up the tower with no support other than from the partially

1 / See table 4.5(a) and (b).

5
completed structure of partially matured concrete. An interior view of

a similar tower under construction is shown in figure 2.4. The fourlevel scaffolding system was suspended from the upper part of the

tower. A detailed description of the system is illustrated in the U.S.

Patent of the system, which is reproduced in appendix A.

The lift form and scaffolding system consisted of four basic functional components. These were:

1) Jumpform beams

2) Jacking frames

3) Stiffback channels and formwork

4) Scaffolding

A cross section through the formwork-scaffolding system is shown in

figure 2.5 in which each of the above components is identified. In the

text that follows, the function of each of these components is described

and illustrated with the aid of photographs taken at Berwick, Pennsylvania

where similar hyperbolic cooling towers were under construction using

basically the same construction method used at Willow Island.

Figure 2.6 shows the manner in which jumpform beams were attached to the

exterior of the shell by hexhead bolts. Typical extruded aluminum jumpform

beams which were attached to the interior and exterior of the shell are

shown in figure 2.7. Two 10-ft (3-m) long beams were spliced together end

to end with clip plates forming a 20-foot (6-m) long section. Figure 2.8

shows the 20-ft long jumpform beams attached to the inside and outside of

the shell by means of a number of anchor bolts with internally threaded

cone-shaped inserts and the hexhead bolts (also see fig. 6, appendix A).

The assembly of the anchor bolt is shown in figure 3.8. It may be noted

6
in figure 2.8 that the relative positions of these jumpform beams change

in an alternating sequence for each lift of the shell.

The exterior jumpform beam differs from the interior jumpform beam

in that the flange of the exterior jumpform beam functions as the rib

mold of the shell. The exterior flange of both the interior and exterior

jumpform beams has welded lugs to receive the reciprocating and stationary

pawls for engagement of the jacking mechanism (see fig. 2.7 and figs. 5

and 8, appendix A).

The second major component of the system was the jacking frame

(fig. 2.8). It consisted of two structural channels which rode on both

faces of the outer flange of the jumpform beam (see fig. 5, appendix A).

A photograph of the jacking frame assembly is shown in figure 2.9. Movement of the jacking frame was restricted to one direction, i.e., parallel

to the jumpform beam, by means of a set of 12 steel wheels housed in the

jacking frame.

The jacking frame also contained a hydraulic ram with a connected

reciprocating pawl and a follower pawl (item 142 and 124 in figure 5,

appendix A). When jacking commenced, the reciprocating pawl was actuated

by the hydraulic pressure and engaged with the jumpform beam lugs raising

the jacking frame in increments equivalent to the lug spacing. The springloaded follower pawl ratcheted and held the jacking frame until another

hydraulic cycle could be repeated.

The third major component in the construction apparatus consisted of

the stiffback channels and formwork (fig. 2.8, 2.9 and 2.10). This component held the fresh concrete in place between each rib location. The

stiffback was a structural channel vertically supported by the jacking

7
frame's lower steel wheel axle (see figures 6, 7 and 8, appendix A).

Adjustable joists were bolted to the stiffbacks at the rib locations and

were capable of compensating for the shell diameter variation. The plywood

form was braced by the joist-stiff back framework. The stiffback and formwork

moved integrally with the jacking frame during the entire jacking process.

Jacking took place at both the inside and outside jumpform beams simultaneously at all 96 rib locations. Jacking was terminated once the stiffback formwork had cleared the previous lift and had been positioned for

placement of concrete for the next lift.

The fourth major component of the lift form-scaffolding system was a

four-level scaffold (fig. 2.5). Working platforms were suspended from

both the inside and the outside jacking frames at each of the 96 rib locations. Scaffold planking and guardrails spanned between the rib locations

forming the working platform. At the top level construction materials were

received via the hoisting system, steel reinforcing was distributed for

placement, and concrete was delivered around the tower by Georgia buggies.

Level 2 was used less in the construction process. From this level

the stiff back-formwork was accessible for adjustment. Once the formwork

was set for a new pour the use of this level became diminished.

The lower levels (levels 3 & 4) provided access to the bottom half of

the jumpform beams. The 10 ft (3 m) section of the jumpform beam was

separated from the upper half and unbolted from the shell and delivered

to the top level for new placement. Final shell surface preparation (i.e.,

patching, removal of threaded insert cones, grouting) was performed from

these levels. The entire scaffolding system moved with the jacking frame

to the new elevation.

8
The sequence of operations for the entire formwork-scaffolding system

from lift 27 to lift 29 is illustrated in figures 2.11 through 2.15.

Figure 2.11 depicts the position of the jumpform beams and the scaffolding

system prior to concrete placement for lift 27. It may be noted that the

inside jumpform beam was higher than the outside one. Two days after the

completion of lift 27, the formwork and scaffolding systems were raised

and simultaneously the lower half of the outside jumpform beam was unbolted

and lifted to a new higher position, shown in figure 2.12. Figure 2.13

shows the final position of the entire system prior to concrete placement

for lift 28. One day after the completion of lift 28, the formwork and

scaffolding systems were again raised and the lower half of the inside

jumpform beam was lifted to a new position, as shown in figure 2.14.

Figure 2.15 shows the final position of the formwork and scaffolding

systems with the inside jumpform beam extending above the outside one.

The daily preparation for concrete placement consisted of five consecutive procedures as follows:

1) Workers removed wedging and loosened stiffbacks and formwork from

previously placed concrete. Plywood forms were removed next and

cut to new desired size to accommodate changing diameter of the

tower. Forms were then placed again in the same place after

being cleaned and oiled.

2) Once all forms were replaced, jacking of the entire formwork

and scaffolding systems commenced until the next predetermined

elevation was reached.

3) The lowest trailing jumpform beam was unbolted and transferred

to its new top position on top of the upper half mate of the

9
jumpform beam. Therefore, each pair of jumpform beams (inside

and outside) at each rib location "leap-frogged" one another

throughout the construction process. Figures 2.11 through 2.15

illustrate this process.

4) The formwork was wedged into its new lift position, spacing

tolerances checked, reinforcing steel placed, and concrete

was then ready for placement.

2.4 Hoisting System

At the Willow Island site, up to and including lift 10, construction

materials were handled by moving cranes on the ground and concrete was

placed into the formwork by a pumping process. Above this level, the

materials and concrete were delivered to the working platform via six

cathead gantry cranes (hereafter referred to as catheads) powered by

twin-drum hoists.

The six catheads were spaced at equal intervals around the top

perimeter. Each cathead was supported from the recently completed shell

structure by four legs which were attached to the jacking frame at two

adjacent rib locations. The catheads moved up with the lift formscaffolding system as construction advanced.

A static line, which was attached to the slide plate at the interior

end of the cathead at one end (fig. 2.16) and secured to an anchor point

on the ground level at the other, guided all materials hoisted to the

top working level (fig. 2.17). The anchor point location changed from

a position near the wall of the tower to the center as construction

progressed upward. During hoisting operations, the tautness of the static

line was adjusted by means of a grip hoist attached to the ground anchor

10
point to keep the material being hoisted from hitting the scaffold at

the top of the tower.

Because of the changing shell curvature, the cathead had to be

periodically adjusted to maintain a level configuration. This was done

throughout the tower construction by adjusting the pinned telescoping

outer legs and by a chain hoist which was an integral part of the counter

static line (figure 2.16) which also counteracted the pull exerted by the

static line.

Three twin-drummed diesel hoists were spaced at equal intervals

around the tower base. Each twin-drum hoist had a single operator who

served two cathead gantries (fig. 2.18). A hoist line, wound on a drum,

passed through two ground-level sheaves and traveled up along the outside

face of the tower to the outside sheave of the cathead beam (fig. 2.17).

It continued to the inside sheave of the cathead beam and down to the

tower center ground location where pickups were made.

The hoist load was guided along the static line by a mechanical

pulley device. Figure 2.16 shows a bucket of concrete approaching its

maximum elevation at the inside working level. The interlink to the

static line through the pulley is also shown.

The hoist system was primarily used for lifting concrete and steel

reinforcing bars to the upper working platform on the inside of the tower

As was the case for the cathead, various adjustments to the hoisting

system were required from lift to lift due to the changing geometry of

the tower.

11
3 SITE INVESTIGATION

3.1 Introduction

During the course of the investigation, the NBS team visited the

Willow Island site on five occasions. The first visit was made on April 29,

1978, two days after the incident, while the last visit was on August 10 and

11, 1978. This chapter describes observations made, measurements taken of

static and hoist cables, and materials retrieved from the Willow Island site.

In addition, a summary of interviews conducted by OSHA personnel with

workers is presented. These data were used in the examination of possible

component failure of the hoisting and scaffolding systems in chapter 5 and

in the establishment of the loading conditions on the tower which existed

at the time of collapse in chapter 6.

3.2 Investigation at the Site

Figure 3.1 shows the exterior view of tower unit No. 2. Collapse

occurred as concrete was being placed for lift 29. The top 5 ft (1.5 m)

of the shell (lift 28) is missing in this photograph since the entire lift

had fallen into the center of the tower leaving a jagged edge along the top

of lift 27 (fig. 3.2). Evenly spaced ribs on the exterior surface of the

shell and 5-ft (1.5-m) lift lines may also be seen in figure 3.1. The

exterior scaffolding-type stairs shown in the left part of the photograph

were the only means of access to and from the top of the tower by the

workers (fig. 3.1),

Figures 3.3 and 3.4 show debris piled upon the floor of the tower

around the base perimeter. While the general positions of the debris

inside the tower have not been drastically changed, it is known that the

debris was lifted and moved to some extent during the rescue operations.

12
The debris mainly consists of the cathead gantries, scaffolding and formwork systems, safety nets, and equipment. The concrete of lift 28 apparently pulverized upon impacting the tower floor since no large pieces of

broken concrete were found.

As can be seen in figures 3.2 and 3.5, vertical bars at the top of

lift 27 were bent toward the inside of the tower. None of the vertical

bars examined at locations around the top perimeter of the tower showed

any signs of fracture.

Figure 3.6 shows the exterior view of a portion of the shell between

catheads no. 1 and no. 2. Dark patch marks seen in this figure were made

by the cement grout used to plug the anchor bolt holes. These marks clearly

indicate that two anchor bolts were used in lift 26 and others below. A

closeup view of a typical rib in lift 27 (figure 3.7) shows that while the

top anchor bolt (bolt C in fig. 2.15) was broken away, the bottom anchor

bolt (the thread inserts and the crimped rod—bolt D in fig. 2.15) was intact.

Careful examination of the jumpform beams on the ground revealed that in

many cases the top anchor bolts were still attached to the flange of jumpform

beams by hexhead bolts (see fig. 3.8 for the anchor bolt assembly). Further

examination also revealed that there were no tears in the flange of jumpform

beams at the bolt holes nor any fragments of the hexhead bolt in the bottom

anchor bolt in lift 27. It was concluded from these observations that at

the time of the collapse, while the top anchor bolts were attached to jumpform beams, the bottom hexhead bolts had been removed.

Localized crushing of concrete occurred surrounding the bottom anchor

bolt in lift 27 at those ribs where the legs of catheads had been supported

(figure 3.9). This indicates substantial overworking on these bolts during

13
the previous construction stages as a result of large movements in the

cathead legs.

It was observed that the anchor bolts located at the top of lift 27

(bolt C in fig. 2.15) failed in two different modes. In many cases, the

anchor bolt together with the concrete surrounding it broke away from the

shell (see fig. 3.10). In other cases, the bolt itself fractured into two

pieces leaving a clean tapered hole in the shell (see fig. 3.11). It would

not be possible to leave such a clean hole in the shell if the cone-shaped

inserts had been pulled through the wall. This indicates some bolts fractured

At the time of the field investigations, it was reported by workers at

the site that the collapse initiated at the location where cathead no. 4 was

positioned. Because of this reason, cathead no. 4 was carefully examined.

The overall appearance of a typical cathead gantry is shown in figure 3.12.

Although all six cathead gantry assemblies deformed severely, the component

parts for each of the cathead gantries were still intact except for the

slide plates which were separated from catheads no. 4 and 5. Figure 3.13

shows the slide plate for cathead no. 4 which was found some distance away

from the main assembly. A large bolt to which the slide plate was attached

was bent severely. A similar deformation of the bolt was also noted at other

gantries (fig. 3.14). The diameter of the sheave which was attached to the

slide plate as well as others used for the hoist cable was 12 in (305 mm).

Figures 3.15 and 3.16 show the inside and outside legs of cathead no. 4.

All four legs remained straight and showed no sign of damage. As is seen in

figure 3.16 the cathead beam was severely bent. However, close examination

showed no sign of buckling of the beam. It was concluded from this observation that the beam was bent after the collapse had initiated.

14
The hoist cable used was 9/16 in (24 mm) 19 x 7 non-rotating wire

rope. Of the total of six hoist cables, one per cathead gantry, only the

cables for catheads no. 4 and no. 5 were broken. The cable for cathead

no. 4 was broken at a distance 408.5 ft (124.5 m) from the axis of the

drum hoist, while the cable for cathead no. 5 was broken at the hoist

end just above the dead-weight steel ball. The remaining four hoist lines

i were draped over the tower following the collapse. At both catheads no. 4

and 5 the hoist cable cut a deep groove in the shell. At cathead no. 4,

the cable cut a 4 in (100 mm) groove (fig. 3.17) and at cathead no. 5 a

5 in (125 mm) groove. The part of the cable from the bucket to the breakage point was still attached to the bucket. Close examination showed

that the concrete bucket for cathead no. 4 fell to the ground with concrete

in it. Figure 3.18 shows part of the concrete which remained inside the

bucket. On the other hand, the concrete bucket for cathead no. 5 was found

empty.

The drum hoist which served cathead no. 3 and no. 4 is shown in

figure 3.19. Subsequent to the collapse, the hoist cable for cathead

no. 4 was played out for visual observation of the condition of the cable.

It was noted that the drum hoist was operational. As may be seen from

figure 3.19, the operator of the hoist sat inside the shed which housed

the hoist. In addition to audio communication with a worker who was on

the top of the tower, the operator normally used paired visual markers on

the cable to judge the position of the material being hoisted. When the

painted markers on the cable passed over the wooden cross beams located

in front of the hoist, the operator reduced the speed of the drum so that

15
the final positioning of the material being hoisted could be handled

safely. On the hoist cable for cathead no. 4, no such markings were

noted

.

Based on the measurements and the cable profile, it was determined

that the static line was 1/2-in (13-mm) diameter 6 x 19 classification

wire rope. One end of the line was attached to the slide plate and the

other end was anchored at the ground level. As was mentioned in

section 2.4, the anchor point was moved as construction progressed. It

was moved to the center of the tower on April 17, 1978, on the day

lift 25 was placed.

The lengths of the static line and the hoist cable for catheads no. 4

and 5 were measured. The measurements were taken by a professional surveyor

for cathead no. 4 and by NBS personnel for cathead no. 5. They were:

Cathead gantry No. 4:

Static line length = 219.2 ft ( 66.8 m)

Hoist cable length = 470.4 ft (143.4 m)

Cathead gantry No. 5:

Static line length = 221.9 ft ( 67.6 m)

Hoist cable length = 408.6 ft (124.5 m)

The above lengths of the hoist cables were taken from the axis of the

drum to the top of the dead weight steel ball (fig. 2.16). For the static

lines, the lengths were taken from the clevis which was attached to a concrete hopper located near the center of the tower to pivot point G of the

clevis which was attached to the slide plate of the cathead gantry (see

fig. 6.7).

16
3.3 Interview Statements of Workers

During the course of the investigation of the collapse, many of the

workers at the Willow Island site were interviewed by OSHA personnel.

Based on the interviewed statements, the work schedule of a typical day

can be described as follows:

(1) At approximately 6:30 a.m. , the carpenters arrived at the construction site. Immediate removal of various wedges and loosening of formwork took place around the entire tower perimeter.

This operation was executed from the second working platform.

The plywood formwork was scraped, trimmed to the new desired

dimensions and oiled, and put back into place between the stiffback channels. Simultaneously, laborers removed bolts from the

lower jumpform beams. These bolts were readily accessible from

the lower working platforms 3 and 4 prior to upward jacking

of the scaffolding system.

(2) At approximately 7:30 a.m., the iron workers arrived at the

construction site and prepared bundles of reinforcing steel for

hoisting. Adjustments were made in the static line and the reinforcing steel was hoisted to the top. The iron workers also

distributed and placed the reinforcing steel around the tower

perimeter from the top working platform.

(3) The field engineer arrived at the site at approximately

7:30 a.m. and aligned jumpform beams at 16 control locations.

(4) Jacking of the entire scaffolding and formwork system commenced

at about 8:30 a.m. Upon completion of jacking the lower jumpform beams were removed from the shell around the entire tower

17
perimeter and raised manually to the top level. Formwork was

then repositioned and aligned for the new pour.

(5) At approximately 10:00 a.m., the first bucket of concrete was

hoisted to the top level. During the hoisting of this first

bucket of concrete, the static line tension was readjusted.

The concreting operation began at catheads no. 4 and 5 and progresses in two directions toward catheads no. 1 and 6 where

the operation terminated. While concrete was being placed,

tying of reinforcing steel, jumpform beam repositioning, surface

preparation at the cold joint and grouting, curing agent application, etc. , took place and terminated prior to completion of

concrete placement. The remainder of the day, until approximately

2:00 or 3:00 p.m., was used for concrete placement for the entire

five-foot lift. Following placement of the concrete all workmen

retired and the process was repeated on the following day.

Just prior to the collapse, the following chain of events were described

by those who saw the collapse from the center of the tower.

(1) Shortly before the collapse, the first bucket of concrete was

delivered to cathead no. 4 and emptied into Georgia buggies.

(2) Cathead no. 5 received its first bucket of concrete which was

also emptied into Georgia buggies.

(3) When the workers at the center of the tower heard a loud cracking

sound coming from the direction of cathead no. 4, they saw the

second bucket of concrete was about two-thirds of the way to the

top. The static line went slack. The hoist operator for cathead

18
gantry no. 4 also reported the loss of tension in the hoist cable

at about the same time. He then applied the brake on the hoist line.

(4) Cathead no. 4 slowly fell toward the inside of the tower together

with the scaffolding. The collapse of lift 28 commenced and it

appeared to progress toward cathead no. 5. Eventually, lift 28

peeled off with failure radiating circumferentially in two opposite directions. All scaffolding and forms fell toward the inside

of the tower.

(5) Those who were at the center of the tower took safety under the

concrete truck ramp. All workers who were on the elevated scaffolding were killed in the collapse.

19

/
4. LABORATORY INVESTIGATIONS

4 . 1 Introduction

The laboratory investigations included testing of concrete specimens

under simulated field conditions to establish the strength and stiffness

values of the concrete of lift 28 at the time of the collapse. The results

are also used in determining the elastic moduli of concrete of other lifts

which are needed for computer analyses of the shell (presented in chapter 6).

Because it was not possible to determine visually the operating condition of several components of the scaffolding and hoist systems they were

examined in the NBS laboratory. They were also tested to determine their

ultimate load carrying capacities. The components recovered from the

Willow Island site included Williams anchor bolts, two sections of hoist

cables, a chain hoist and a grip-hoist. The results of these component

tests are used in evaluating any component failures which might have triggered the collapse. Examination of possible component failures is presented in chapter 5. Laboratory investigations on concrete and testing

of each of the components are described below.

4 . 2 Concrete Tests

The purpose of the concrete tests was to establish the compressive,

tensile and bond strengths and the stiffness of the concrete of lift 28 at

the time of the collapse. Although the results of field-cured standard

cylinders, made of the same concrete used for lift 28 and tested at 24 and

25 hr, respectively, were known (see table 4.5 a and b) , supplementary tests

were necessary to establish the rate of stiffness development and rate of

gain in bond strength.

20
When the collapse occurred, about 18 hours had elapsed since the

completion of concrete placement for lift 28. However, because that part

of lift 28 where cathead no. 4 was located had been completed earlier,

it was estimated that the concrete in that vicinity had been cured about

20 hours prior to the collapse. Therefore, the 24-and 25-hr test results

could not be used directly in the analysis of the shell.

In order to duplicate the concrete used for lift 28 in the laboratory,

the constituent materials for the concrete were obtained from the concrete

supplier who provided the concrete for tower unit no. 2. The materials

included cement,- fine and coarse aggregate, fly ash, water-reducing admixture

and air-entraining admixture. These materials will be referred to in the

text as the NBS sample. To assure that these materials, which were obtained

six weeks after the collapse, were the same as those used for the lift 28

concrete, the individual materials were compared with small quantities of

the materials obtained by OSHA personnel from the concrete supplier on

May 2, 1978. These materials will be referred to as the OSHA sample.

Comparative analyses of the two samples, one obtained by NBS and the other

by OSHA, were carried out by NBS and other laboratories.

4.2.1 Examination of Constituents of the Concrete

Cement

Analysis of both the physical and chemical properties of the NBS and

OSHA samples were made to examine, first, whether these two samples conformed to the standard requirements given in ASTM C 150 [4.1] ' and second,

whether the two samples had the same properties. The ASTM specified

2 / Numbers in brackets refer to references listed in Chapter 10.

21
3 values, the test results of Law Engineering Testing Co. , and the test

results of the Cement and Concrete Reference Laboratory, NBS, are given

in table 4.1. Except for the 7-day compressive strength of cement mortar,

other physical properties conformed to the ASTM requirements. It is to

be noted that while both samples showed a lower 7-day compressive strength

than the ASTM specified value, the NBS sample showed 16 percent greater

strength than the 7-day value of the OSHA sample. The difference is even

greater for the 3-day strength wherein the NBS sample showed 27 percent

greater compressive strength than that of the OSHA sample.

The properties of two cement samples were determined by the chemical

analysis method given in ASTM C 114 [4.2], The analysis was carried out

by Law Engineering Testing Laboratory. The results of the two cement

sample analyses are given in table 4.2 together with the ASTM standard

requirements. It is seen that although the cement samples were essentially

of similar composition, both NBS and OSHA samples did not conform to the

ASTM standard requirements for Type II cement, in that the silicon dioxide

content of the samples was less than the ASTM minimum value. Therefore,

it may be concluded that, based on both physical and chemical properties,

the cement obtained from the concrete supplier would not have met the ASTM

standard requirements for Type II cement. In addition, while the chemical

properties of the two samples compare well, a substantially greater compressive strength of the NBS sample than that of the OSHA sample indicates

that the strength of concrete made of the NBS sample at early ages would

be greater than that of the OSHA sample cement.

3/ See Appendix C for the complete reports.

22
Fly Ash

Table 4.3 shows the elemental analyses of fly ash of the NBS and OSHA

samples. The values listed in the table were obtained by means of atomic

emission and absorption spectrometry which was carried out by NBS. The

tabulated results suggest that the two fly ash samples were essentially

the same. The total contents of alumina, iron oxide and silica in each

of the samples were calculated and compared in table 4.4 with the requirement given in ASTM C 618 [4.1], It is seen that both samples conformed

to the ASTM standard.

Fine and Coarse Aggregate

Fine and coarse aggregate consisted of natural river sand and gravel.

Those obtained from the concrete supplier were visually compared with the

OSHA samples. No difference between the two samples was noted.

Water-Reducing Admixture

A water-reducing admixture is a material used for the purpose of

reducing the quantity of water required to produce concrete of a given

consistency. These admixtures increase the slump of concrete for a given

water content. Some admixtures may also retard the setting time of concrete.

The NBS and OSHA samples were compared by means of infrared spectrophotometry which was carried out by NBS. The chemical composition of

the two samples as shown by the analysis indicates that the NBS and OSHA

samples were essentially the same.

An infrared analysis of concrete constituents from a piece of concrete of lift 28 was made to determine the amount of admixture used in

the concrete. This was carried out by the Portland Cement Association.

23
The results of the analysis are given in Appendix D. The analysis revealed

that the amount of the water-reducing admixture used in the concrete was

about the same as the amount specified in the concrete design mix specification, i.e. , 3 oz per 100 lb (89 mL per 45.4 kg) of cement.

Air-Entraining Admixtures

Air-entraining admixtures are generally used to improve the durability

of concrete exposed to cycles of freezing and thawing. The workability of

fresh concrete is also improved, and segregation and bleeding are reduced

greatly. The NBS sample of the air-entraining admixture obtained from the

concrete supplier was compared with the OSHA sample by means of infrared

spectrophotometry. This analysis was carried out by NBS. The results

showed no significant difference between the NBS and OSHA samples.

4.2.2 Determination of Concrete Strengths and Stiffness

After having determined that, except for possibly the cement, the

constituents of the concrete obtained from the concrete supplier by NBS

were essentially the same as the ones obtained by the OSHA personnel from

the same source shortly after the collapse, a laboratory test program was

initiated to determine various strengths and stiffness values of concrete.

Included were tests for compressive strength, pullout bond strength and

modulus of elasticity.

Test specimens were prepared and cured in an environmentally controlled

chamber. Temperature in the chamber was controlled to simulate the temperature conditions at the Willow Island site over the 24 hr period immediately

prior to the collapse. The chamber temperature was controlled using the

data obtained from the Parkersburg airport which is located about 5 miles

(8 km) from the Willow Island site. It should be noted that the airport

24
is situated at an elevation of about 170 ft above the Ohio River on which

the tower was situated. The temperature variation prior to the collapse

based on the airport data and the temperature variation used for curing of

concrete specimens are given in figure 4.1.

3 3 The concrete was prepared in several batches using 2 ft (0.057 m )

3 3 and 10 ft (0.28 m ) capacity mixers. The mix proportions were the same

as those reported for the lift 28 concrete (table 4.5). The mix proportions specified for one cubic yard of concrete were scaled down to match

the size of mixer. Prior to mixing, the components of the concrete were

prechilled for at least 18 hr in an environmental chamber at 45°F (7.2°C)

to simulate the temperature condition of the components for the lift 28

concrete at the batching plant. To duplicate the mixing condition, hot

water was added for mixing. Table 4.6 gives the data on fresh concrete

obtained at the time of specimen preparation.

The following tests were performed.

1. Compressive strength test of 6 x 12 in (150 x 300 mm) cylindrical

specimens

.

2. Bond strength tests using 8 x 8 in (200 x 200 mm) cylindrical

pullout specimens.

While several series of compressive tests were made to examine the

strength-gain characteristics, only one series was carried out for the

pullout bond tests. For the first 24-hr period after casting, all specimens were subjected to a simulated field temperature condition as described

above. Thereafter, the specimens were cured at 55°F (12.8°C). For 28-day

test, a separate set of three companion compression specimens were cured

at 73°F (22.8°C). The actual temperature of the concrete was recorded

25
periodically by means of a thermocouple inserted in a 6 x 12 in (150 x

300 mm) cylinder. The specimens cured in the chamber were tested at

0.5, 1, 2, 3, 5, 7, 14 and 28 days.

Compressive Strength

The compressive tests were performed according to the procedure

described in ASTM C 39 [4.1], Figure 4.2 shows the compressive test setup.

Deformation of each specimen was measured during the compressive test with

a compressometer such as is described in ASTM C 469 [4.1], The results of

deformation measurements together with strength data were used to determine

the modulus of elasticity.

Figure 4.3 shows the results of the compressive strength tests in

which the compressive strength is plotted against the maturity of the

concrete. The term "maturity" expressed in units of "degree-day" represents a simple function with which the combined effect of temperature and

time can be related to the gain in the concrete strength [4.5].

The NBS test data, plotted as squares, are shown in the figure along

with the results of compressive tests carried out by the Ohio Valley

Testing Laboratory (OVT) and the Pittsburgh Testing Laboratory (PTL),

plotted as triangles, and the field test data of 6 x 12 in (150 x 300 mm)

cylinders for the lift 28 concrete, plotted as circles. The test specimens

used by the Ohio Valley Testing Laboratory and the Pittsburgh Testing Laboratory were made on May 2, 1978, at the Willow Island site using the

concrete delivered by the concrete supplier. These specimens were field

cured for the first 24-hour period and thereafter, in 70°F (21.2°C) lime

water. The specimens prepared at the time lift 28 was cast were kept

26
at the base of the tower for the first 24-hour period and subsequently,

moved to a 70°F (21.1°C) fog room.

It is seen in figure 4.3 that the cylinder strengths of lift 28 agree

well with the OVT-PTL data. Comparison with the NBS data shows that at

early ages, the OVT-PTL data fall below the NBS data, while at later ages,

the OVT-PTL data lie above the NBS data. This indicates that the rate of

strength gain represented by the two sets of data are different. The

results of the cement analysis discussed in section 4.2.1 suggest that

this difference could be attributed to a slower early strength gain

exhibited by the cement of the OSHA sample, which was obtained on the

same day as the OVT-PTL specimens were prepared. Because the compressive

strength data of lift 28 matched closely with the strength gain characteristics of the OVT-PTL data, the OVT-PTL data were used in the analysis of

the shell. However, for the relationship between the compressive strength

and the modulus of elasticity and the pullout bond strength, the NBS test

results were used because they were the only data available.

Modulus of Elasticity

From stress-strain curves of the compressive test the values of the

secant modulus of elasticity are obtained at the stress level of 40 percent

of the maximum compressive stress (0.4 f c

). The stress-strain relationship

of concrete specimens tested at various ages together with the relationship

between the compressive strength and the modulus of elasticity is shown in

figure 4.4. In this figure, the secant moduli are plotted against the square

root of the compressive strength. A linear regression line is shown together

with a line obtained by using the ACI code equation [4.4]. In the analysis

27
of the shell, different moduli of elasticity determined by the regression

line, for each lift were used.

Pullout Bond Strength

The testing procedure for the pullout bond tests was essentially as

described in ASTM C 234 [4.1] except that only the slip of the bar at the

free end was measured. Figure 4.5 depicts the test setup. The test specimen consisted of a 48-in (1.2-m) long no. 4 (1/2 in, 12.7 mm) deformed

reinforcing bar (ASTM 615, Grade 60) cast in an 8 x 8 in (200 x 200 mm)

waxed cardboard cylinder mold. Figure 4.6 shows the pullout specimen and

mold. The specimens were made so that the bar was bonded to the concrete

for a length of 6 inches (152 mm).

Figure 4.7 shows the pullout test results in which the maximum test

load observed is plotted against the square root of the corresponding

compressive strength. The bond strength of the no. 4 bar obtained according

to the regression equation was used in the strength evaluation of the shell.

4 . 3 Anchor Bolt Tests

As described in chapter 3, special anchor bolts were used to hold the

inside and outside jumpform beams together, separated by a constant distance,

prior to casting of concrete. After the concrete had set, although friction

existed between the flanges of inside and outside jumpform beams and the

shell, the bolts served as the only positive means of transferring all construction loads to the shell. The components of the anchor assembly are

shown in figure 4.8. The assembly consisted of two 3/4-in (19-mm) hexhead

bolts (ASTM A 494), two rectangular washers, two tapered threaded couplings

and a centerpiece threaded crimped rod. By adjusting the amount of insertion

28
of the rod into the threaded couplings, the space between the two jumpform

beams was adjusted.

Two anchor assemblies after tension testing to failure are shown in

figure 4.9. In both cases the fracture took place at the middle where the

crimp was present. The average of two anchor assembly tests was 40,350

lb (179.5 kN). Because it was observed in the field that in many cases

the rod was fractured in the thread coupling through the net cross section

at the first thread, a specimen was tested in tension with the anchor

assembly embedded in a 12 x 12 x 12 in (305 x 305 x 305 mm) concrete cube.

The purpose of this test was to see whether the failure mode would change

if the anchor bolt were tested with concrete encasement. A photograph

showing the fractured part is shown in figure 4.10. As can be seen, the

male rod fractured through the net cross section at the first thread. The

maximum test load was 43,200 lb (192.2 kN). This latter test load indicated

that the strength of an unembedded anchor assembly would give a lower bound

capacity

.

4.4 Hoist Cable Tests

Two 100-ft (30.4-m) hoist cables were obtained from the Willow Island

site. One section was cut from the hoist end of the cable, and another

from the opposite end of the cable wound on the hoist drum. The hoist

cable had a diameter of 9/16 in (14.3 mm) and was made of non-rotating wire

rope having two 19 strand layers.

For laboratory tests, ten 6-ft (1.8-m) long specimens were prepared,

five from each of the two sections of cable (figure 4.11). Wire rope

sockets were attached to the ends of each specimen by means of molten zinc.

29
Figure 4.12 shows the test setup for the tension test. Of the ten

specimens, eight broke within 2 ft (0.6 m) of one of the socket ends,

while the other two broke close to the middle of the specimen. The

results of tension tests of the cables are given in table 4.7.

The tabulated values show that there is no significant difference

between the strength of the cable section which had been subjected to

continuous hoisting stress and that of the cable section which had been

worked on the hoist drum.

4 . 5 Chain Hoist Test

A 2-ton (17.8 kN) capacity chain hoist was used in cathead no. 4

as part of the counterstatic line. The function of the chain hoist was

to adjust the length of the counterstatic line so as to counterbalance

the force in the static cable and adjust the position of the cathead

gantry. Because any slippage in the chain hoist gear system due to overloading could result in a continuous increase in the chain length and

subsequent dynamic instability of the cathead gantry, it was desirable

that the maximum strength of the chain hoist be determined.

The test setup shown in figure 4.13 duplicates the field condition using

the cable slings and clevises of cathead no. 4. When the test load reached

15,310 lb (68.1 kN), the internal gear system fractured without the chain

slipping. The fracture of the gear system also shattered the cover plate

of the housing of the chain hoist (figure 4.14).

4.6 Grip-Hoist Test

As shown in figure 4.15, the static line was anchored to the ground

by means of a grip-hoist. By jacking the hand lever on the grip-hoist,

the static cable length and tension could be adjusted. Because any slippage

30
of che static line through the grip-hoist could cause cathead no. 4 to

become unstable the grip-hoist capacity needed to be determined.

Figure 4.16 shows the tension test setup in which the grip-hoist was

placed between the head and platen of the testing machine with the static

cable attached to the lower end. At a maximum load of 19,100 lb (84.96 kN),

the hook on the lower end broke away from the housing of the grip-hoist

(fig. 4.17). Throughout the entire range of test load, no slippage of

the static line was observed.

31
5. EXAMINATION OF CONDITIONS OF CONSTRUCTION

ASSEMBLIES AT THE TIME OF FAILURE

5.1 Introduction

The construction assemblies which collapsed into the tower may be

divided into three groups: the hoisting system, the scaffolding system

O O and the formwork system. Because only 1.0 yd (0.8 m ) of concrete was

delivered to the top of the tower at the time of the collapse, it is

reasonable to conclude that no appreciable forces were being exerted on

the formwork system, and that any local failure of the formwork system

would not have been possible. Therefore, this chapter examines possible

failures of those components of the hoisting and scaffolding systems

which could have caused cathead no. 4 to become unstable under imbalance

of forces. Examination of each critical component is supported by the

data collected from site investigations and, where applicable, by test

results.

5.2 Examination of Hoisting System

Figure 5.1 identifies the main components of the hoisting system as

1. Hoist cable

2. Static line

3. Counterstatic line

4. Chain hoist

5. Cathead beam

6. Outside legs

7. Inside legs

8. Outside support brackets

9. Inside support brackets

32
10. Slide plate assembly

11. Static line anchor - Grip-hoist (not shown in fig. 5.1,

see fig. 4.15)

Failure of any of the above items might cause a sudden shift of forces

which would create an unstable condition for the cathead.

Hoist Cable

Two of the six hoist cables were broken. These were located at

cathead no. 4 and cathead no. 5. A comprehensive examination of the fractured area by the Federal Bureau of Investigation reported no kinking,

crushing or "bird caging” of the cable and other signs of damage which

could have caused distortion in the cable structure.

As mentioned in section 3.2, both of these cables cut a deep groove

in the upper edge of the remaining shell, lift 27, indicating that they

were not broken until after the debris had started plummeting towards

the ground. It is believed that a substantial amount of weight must have

been hanging on the cable to cut the deep groove and that the break of

the cable followed the initiation of the collapse. The laboratory results

indicated that the average breaking strength of the cable of ten specimens

is 27,260 lb (121 kN). Even under an impact factor of 2.0 applied to the

hoist load, a maximum operating load which would be produced by the concrete

bucket and the attachments would be about 4,932 lb (22 kN) (table 6.1).

Thus, it is reasonable to conclude that this small magnitude of the hoist

load, as compared with the breaking strength, could not have caused the

cables to break, and that the falling weight of the cathead together

with the scaffolding system must have fractured the cable.

33
Static Line

Examination of all six static cables during the field investigation

showed no sign of fracture along their entire length. All of the cables

were still attached to the sliding plate at one end and anchored to the

ground through a grip-hoist at the other end.

Counterstatic Line and Chain Hoist

The counterstatic line consisted of a set of two cable slings interconnected in series by a chain hoist. As presented in chapter 4, the laboratory test revealed that the weakest link in the counterstatic line was

the chain hoist which had a maximum test capacity of 15,310 lb (68.1 kN).

Because the test also showed the chain hoist would fail in fracture

rather than by slippage in the gear system, it can be concluded that no

slippage occurred due to the hoist load. This conclusion is also supported

by the closeness of the measured length of the counterstatic line in the

field (see sect. 3.2) with the computed value based on the reconstruction

of the cathead gantry configuration using the actual member sizes measured

in the field.

Cathead Beam

A close examination of the cathead beam of cathead no. 4 revealed

that although the two channels (C8 x 11 . 5)comprising the beam were bent

at the midspan (fig. 3.16), they did not exhibit any signs of buckling

between the two cathead legs. This also agrees with a simple elastic

analysis of the channel section as a compression member. The computed

buckling load exceeded by a considerable margin the force produced by

the hoisting load with an impact factor of 2.0; 23.4 kip capacity vs.

3.58 kip load (104.1 kN vs. 15.9 kN).

34
Inside and Outside Cathead Legs

Figures 3.15 and 3.16 show the inside and outside cathead legs,

respectively. As discussed in section 3.2, examination of all four

individual legs showed no signs of damage or buckling. Support brackets

for attachment of the legs to jacking frames were still intact. No welding failure of the square tubular legs to the supporting bracket was noted

Slide Plate Assembly

The slide plate assembly supported the static cable and the inside

sheave of the hoist line. It could slide along a track attached to the

inside of the two-channel cathead beam. By turning the large bolt above

the beam, the position of the slide plate could be adjusted. As can be

seen in figure 3.13, the slide plate became detached from the cathead beam

Judging from the bent shape of the bolt, the slide plate first separated

from the beam with the bolt still anchored to the clip angle which was

attached to the top flanges of the cathead beam. Subsequently, the downward force produced by the falling weight of the concrete bucket as well

as the debris weight on the hoist cable could have bent the bolt as well

as fractured the bolts which connected the clip angle to the cathead

beam. However, the separation of the sliding plate from the cathead

beam is not unique to cathead no. 4. A similar condition was also noted

at catheads no. 1, 5, and 6. Because they were buried under debris

it was not possible to ascertain the condition of the slide plate for

cathead no. 2 and 3. Based on this observation, no definite conclusion

could be reached as to when the slide plate for cathead no. 4 separated

from the beam.

35
The possibility of the slide plate being hit by the dead weight balls

(fig. 5.1), which were attached to the hoist end of the cable, was examined.

The measured length of the hoist cable (see sect. 3.2 and 6.2) places the

position of the bucket at about 60 ft (18.3 m) below the cathead beam.

This agrees well with eyewitness accounts that the bucket was two thirds

of its way to the top of the tower when the collapse initiated. Therefore,

it can be concluded that the bucket could not have hit the slide plate and

caused it to separate from the cathead beam.

Static Line Anchor

As described in section 3.2, the static line was anchored to the

ground through a grip-hoist. This allowed adjustment of tension in the

line. Laboratory examination of the grip-hoist which anchored the static

line for cathead no. 4 revealed that the two brass shear pins in the ratchet

mechanism were still in place. The shear pins limit the amount of torque

that could be applied to the grip-hoist to increase the tension in the

staticline. It was established from the manufacturer that these brass pins

shear off at a torque level that would induce a 3500-lb (15.6-kN) tension

in the static line. Conversely, if the tension in the static line is

greater than 3500 lb, additional tension could not be applied to the line

by means of torquing the ratchet mechanism. On the other hand, the gripping

mechanism in the grip-hoist had the gripping capacity in excess of 19000 lb

(85 kN). Since the maximum tension force produced by the concrete bucket

at the critical location is considerably greater than this force (4772 lb

vs. 3500 lb) (see sec. 6.7), it can be concluded that the tightening of the

static line through the ratchet mechanism could not have pulled the cathead

gantry inwards to initiate the collapse.

36
5 . 3 Examination of Scaffolding System

As described in chapter 2, the platform supports were attached to

jacking frames which in turn were supported by jumpform beams. The

jumpform beams were anchored to the shell by means of special anchor bolt

assemblies (see sec. 3.2). Because the flanges of the inside and outside

jumpform beams were interconnected by anchor bolts, the anchor bolts in

lift 28 could not be pulled out without either breaking the 3/4 in (19 mm)

hexhead bolt, breaking the crimp rod or tearing the flange of the jumpform

beams. Close examination of the jumpform beams to which the legs of

cathead no. 4 were attached revealed that there were no signs of excessive

deformation at the bolt holes. The minimum specified capacity of the hexhead bolt according to ASTM A 449 is 40,100 lb (178 kN) and the average

tensile strength of the crimped rod was 40,350 lb (179.5 kN) (see sec. 4.3).

Since the computed tension force produced in the anchor bolt assembly

by a combination of the loading conditions including the hoist load with

a dynamic amplification factor of 2.0 was only 16,220 lb (71.2 kN), it was

concluded that the bottom anchor bolts in lift 28 could not have failed

prior to the collapse.

5.4 Summary

In the foregoing sections of this chapter the critical components of

the hoisting and scaffolding systems were examined with the support of field

laboratory test data. It was shown that each of the components did not

fail prior to the initiation of the collapse. Therefore, it may be concluded that the collapse did not initiate due to any component failure of

the hoisting and scaffolding systems. In the following chapter, the shell

will be analyzed to see whether its capacity was adequate to support the

superimposed construction loads.

37
6 DEFINITION OF LOADS AND SHELL ANALYSIS

6.1 Introduction and Background Information

Criteria and rationale for the definition of loads acting at the time

of the collapse are developed in sections 6.1 through 6.8 for use in the

analytical investigations of the shell presented in section 6.9. A total

of five basic sources of loading (load cases) and the manner of their distribution on the shell are identified for this analysis as follows:

1) weight of the tower - continuous

2) weight of the scaffolding - 96 ribs

3) live loads on the scaffolds - 96 ribs

4) weight of cathead assemblies - 12 ribs

5) hoisting loads - 2 ribs

Loads attributed to the weight of the tower are generated internally

by the shell analysis programs used in this investigation from the tower

geometry and unit weight of concrete prescribed in the input. The procedure for evaluating loads from the other sources cited above is discussed

in sections 6.4 through 6.7. Assumptions made with regard to dynamic

loading in the derivation of loads are discussed in section 6.8. Geometric

considerations for the definition of loads and the mechanism for their

conversion into forces directly applied to the shell are discussed in

section 6.2 and 6.3, respectively.

The various steps used in the conversion of externally applied loads

into equivalent forces acting directly on the shell structure may not be

readily obvious to those having no prior familiarity with the construction scheme used. Therefore, some background information for the material

contained in subsequent sections is in order.

38
The scaled line drawing shown in figure 6.1 is a key sketch that

should assist the reader in understanding the construction scheme used

at the Willow Island site. It represents a typical sectional profile

of the scaffolding, the cathead assembly, the static and hoist lines,

and the concrete bucket. The position of the bucket shown at the top

is that assumed at the time concrete is unloaded into Georgia buggies

(not shown) located on the top deck of the inside scaffold. The sketch

of the bucket shown by the dotted lines represents an intermediate position during delivery.

The hoist line rides over two sheaves suspended from the cathead beam

as indicated. On the outside, it continues on down to the ground level

(fig. 2 .17) where it is wound around a hoist drum which controls its

movement. The hoisted loads (concrete bucket, bundled bars, construction

hardware, etc.) are attached to the inside terminal of the hoist line,

and, in addition, by means of a pulley mechanism (point K), ride on the

static line GKB supported at points G and B. The static line provides

stability against lateral movement during delivery of materials. The

hoisting loads exert a transverse pull and thereby induce a tensile force

in the static line. As hoisting of materials proceeds up from the ground

level, the tension in the static line increases to a maximum somewhere

close to but below the halfway mark, gradually diminishes thereafter,

and vanishes entirely when the hoist line becomes vertical. At that point

the static line is referred to as becoming slack (i.e., tensionless, or

stress-free) but without play.

To determine the hoisting loads acting at the time of the collapse,

a brief note of explanation is needed about the situation just prior to

39
the collapse. All reinforcing bars and various construction-related

equipment were delivered to the top. The first bucket of concrete

was delivered to cathead no. 4. Worker statements indicated that when

the first bucket was delivered to the top, adjustments in the length of

the static line were made at the grip-hoist below (point I in fig. 6.2 -

sec. 6.2) to remove excessive slack, if any, with the assistance of the

worker who unloads the bucket at the top. Therefore, if the first

delivery of concrete was used as a trial to adjust (tighten) the static

line for subsequent runs, the force in that line would be higher during

the delivery of the second bucket of concrete to cathead no. 4. The

collapse, in fact, did occur during the second delivery. In the meantime,

the first bucket of concrete was delivered to cathead no. 5 and unloaded

into Georgia buggies. This bucket had not commenced its descent at the

time of the collapse, a situation which was established on the basis of

the field data (fig. 6.3 - sec. 6.2). No other hoist lines were in

operation at the time of the collapse so that hoisting loads (load

case 5) occurred only at the two ribs flanking cathead no. 4.

In section 6.7, the hoisting loads at cathead no. 4 are calculated

based on the assumption that the static line becomes slack when the bucket

reaches its unloading position as shown in figure 6.1, so that its stressfree length is represented by the solid line GKB. This assumption is

further verified by direct measurements of the length of the static line

obtained at the site after the collapse, and used in the subsequent

investigation of hoisting loads.

Referring to figure 6.1, it is noted that most of the externally

applied loads (load cases 2 to 5) are transmitted to the pair of opposite

40
jacking frames present at each of the 96 rib locations (the exceptions

are discussed in sec. 6.3). The jacking frames transmit these loads

to the jumpform beams which in turn transmit them to the concrete shell

through a series of interconnecting anchor bolts. The cathead forces,

including the weight of the cathead assembly (load case 4) , and the forces

exerted by the static and hoist lines (load case 5), are transmitted to

the shell through the cathead leg support brackets attached to the jacking

frames, and through the counterstatic line which bears against a wide

flange beam and is attached to the lower diagonal as shown.

Before the analysis of construction loads could be carried out, it

was necessary to gather a substantial amount of information about the

construction scheme, including sizes, lengths and material compositions

of the various components comprising the scaffolding, hoisting and cathead

assemblies. All of this information was assembled through numerous site

investigations and laboratory testing of components recovered from the

wreckage, supplemented and corroborated by information from project

drawings and worker statements. The source and nature of this information is cited at the appropriate places in subsequent discussions.

6.2 Geometry

To evaluate the hoisting loads on the tower, it is first necessary

to define the geometric configuration of the hoisting cables used for the

delivery of construction materials to the top of the shell at the cathead

locations (fig. 2.18). The cables of particular interest are those serving

catheads no. 4 and no. 5 where concrete was being hoisted at the time of

the collapse. The required information was developed on the basis of data

obtained from the site following the collapse.

41
Part of the field survey data is presented in figures 6.1 through

6.3. Figure 6.1 is a sectional profile (through a vertical plane of

revolution) of the scaffolding and hoisting systems at a typical cathead

location. The scaffolding, jacking frames and jumpform beams occur at

each of the 96 ribs while the six catheads, including the static, counterstatic and hoist lines, the skid board and the wide flange beam, occur

at equal intervals (every 16 ribs) along the circumference of the shell.

Each cathead is located between two adjacent ribs and transmits its loads

to these ribs through the counterstatic line (fig. 6.1) and two pairs of

inclined legs at the interior and exterior of the shell, respectively.

In plan, the hoisting layout for catheads no. 4 and no. 5 is shown

in figure 6.2. In elevation, the layout for cathead no. 4 (others are

similar) is shown in figure 6.3. These figures are helpful in visualizing

the mechanism for the delivery of materials to the top of the partiallycompleted tower shell where the casting operations for lift 29 were being

carried out. The lifting of materials at cathead no. 4 is controlled by

the hoist drum operator stationed at U. Starting from the drum, the hoist

line proceeds toward and around an interior sheave attached to sheave

block T on the ground, toward and around exterior sheave Q, up and around

the two sheaves suspended from the cathead beam and, before hoisting

commences, all the way down to a concrete hopper at B near center A of

the tower. A stationary static line is attached to point G on the

cathead beam at the top. At the lower end, it passes through a clevis

secured to a concrete hopper at B, and terminates at grip hoist I which

in turn is secured to the ground.

42
The space coordinates of, and distances between, points given in

figure 6.2 were established by double triangulation for independent verification of results. Note that the cathead is radially aligned (along AQ)

but the horizontal projection BG (shown dotted) of the chord of the static

line is not. Even though this introduces a horizontal force component of

the static line acting on the cathead beam in the direction perpendicular

to the beam axis, its effect is not significant and has been ignored in

subsequent calculations of hoisting loads (sec. 6.7). In figure 6.3,

dimensions such as those along the cathead beam, the diameter and pivotal

distances of the sheaves, the lengths of the cathead legs, the location

of their bases relative to the shell, and the height of lift 28, were

obtained by direct measurement in the field and were verified against the

drawings. The lengths of the hoist lines for catheads no. 4 and no. 5

shown in figure 6.3 were likewise obtained by measurement in the field.

They represent the length from the hoist drum U to the point of attachment

of the concrete bucket (point K in the figure). The cathead leg dimensions

shown are the projections of the actual lengths of the cathead legs which

are inclined with respect to the plane of the figure. Likewise, line BG

in figure 6.3 is the projection of the chord length of the static line

which is inclined relative to the plane of the figure (i.e., point B lies

outside this radial plane).

The results shown in figure 6.3 were obtained from the foregoing

data assuming the cathead beam to be level (see sec. 6.6 for explanation).

They define the coordinates of the top support points for the static

lines for catheads no. 4 and no. 5 as well as their chord lengths and

corresponding horizontal and vertical projections. They also indicate

43
the positions of the concrete buckets relative to the respective catheads

at the time of the collapse, based on the hoist line lengths measured

after the collapse.

6 . 3 Mechanism of Load Transfer

It will be helpful at this stage to review the mechanism of load

transfer to the shell. Essentially, most items are connected to the

jacking frames at the ribs (fig. 6.4). The loads that are transmitted to

these frames are indicated in figure 6.4a and the points at which they act

are keyed to the cross-sectional layout shown in figure 6.4b by circled

numbers. Forces F-^ through F^ and moment M are reactions induced by the

weight of the scaffolds and live loads acting upon them. The connection

at point 5 is fixed and therefore capable of developing a a moment M.

As mentioned previously, these forces occur at all of the 96 ribs of the

shell. Forces A

p , A^, Az

, B

p , and B^ are the cathead leg reaction

components on the jacking frames (as defined in fig. 6.8 and sec. 6.7) and

forces P

c and Qc are forces induced by the counterstatic line bearing

against the wide flange beam (see also figs. 6.1 and 6.4b). These forces

as well as force F

c in the counterstatic line (fig. 6.8) occur only at the

six pairs of adjacent ribs where the six catheads are located (fig. 2.18).

Forces that are not directly transmitted to the jacking frames may be

identified by reference to figure 6.4b. The only force directly applied

to the shell is transmitted through the roller at the lower end of the

exterior scaffold. The roller at the lower end of the interior scaffold

will tend to sway away from the shell so that no bearing force can be

\

expected to develop at that point. On the exterior of the shell, the

upper diagonal with the adjustable ratchet transmits, through connection

44
B-^ at the top, a force which may be resolved into components normal and

parallel to the shell. The normal component is transmitted to the jumpform

beam while the jacking frame receives the parallel component through a

pinned extension. Similarly, the lower diagonal to which the counterstatic

line is attached exerts a force through its lower connection F^. The normal and parallel components of this force are transmitted respectively to

the jumpform beam and the jacking frame. In addition, the connection at

this point develops a horizontal pull normal to the plane of the figure

due to the inclination of the counterstatic line with respect to that plane.

The jumpform beams receive forces through two sets of end rollers and

a pawl attached to each jacking frame (see fig. 5, appendix A). The pawl

is mechanically engaged to one of the lugs on the outstanding flange of the

jumpform beam. The rollers are free to slide along the flange but are

constrained against movement in the normal direction. The forces on the

jumpform beam were calculated by treating the jacking frame as a two-span

continuous beam on two exterior roller supports and an interior pin support

(at the pawl) as indicated in figure 6.4a.

Figure 6.5 identifies forces and couples acting on the outstanding

flanges of opposite jumpform beams at a rib. Points A through F (also

appearing in fig. 6.4a) define the location of the jacking frame supports

relative to the jumpform beams. The two exterior diagonals noted above

are connected to points B-^ and F-^. Forces Fx , F

y

, F

z and circumferential

moment My are identified by reference to the local coordinate axes (x,

y, z) as shown. Check marks and zeros shown in tabular form indicate

respectively those actions (force component or moment) that can or cannot

develop at each point. Actions F and M occur in opposite pairs as

z y 45
noted in the figure, and develop as a result of the inclination of the

cathead legs and counterstatic line relative to the plane of the figure.

The final stage in the conversion of external loads (load cases 2

to 5) into equivalent actions applied directly to the shell leads to the

results .tabulated in figure 6.6. The points of application of these

actions, designated by the letters I, J, K and L (also appearing in

figure 6.5), are located on the middle surface of the shell. Actions

occurring at E and F (refer to fig. 6.5) are converted into their statically equivalent counterparts and placed at K. Actions at F-^ are converted in a similar manner and placed at L. The resultants of Fy forces

and My moments acting on the cantilever portion of the jumpform beams

(points A, B, C, D and B^) are distributed equally to the two anchor

bolts at I and J. The F x forces at I and J induced by the forces acting

on the cantilever portion of the jumpform beams are calculated by treating

the anchors at these junctions as simple supports.

The preceding steps involve certain assumptions regarding the distribution of forces to the anchor bolts. For instance, the equal distribution of Fy forces to anchor bolts at I and J implies that the jumpform

beams are axially rigid and no friction can develop at the shell interface. The other extreme situation would occur if the axial rigidity of

the jumpform beams were negligible relative to that of the shell, in

which case, nearly all of the Fy forces from the cantilever portion would

be transferred to the top anchor bolt. In reality, the axial stiffness

of the shell is many times that of the jumpform beams and therefore,

the actual distribution will be somewhere between the two extremes with

the greater portion of the force going to anchor bolt I. However, the

46
net effect of the two extreme assumptions on the results of the shell

analysis is not significant. In fact, after examining the combined effect

of all the assumptions made in the conversion of loads from the jumpform

beams to the shell, it was determined that additional refinements will

not alter the conclusions of this investigation.

6.4 Weight of Scaffolding - Load Case 2

The scaffolding assembly, including formwork and steel joist supports

(fig. 2.5) is axisymmetric so that it transmits equal forces to the 96

peripheral ribs of the shell (fig. 3, appendix A). The procedure for evaluating the forces at a rib due to the weight of a repetitive scaffolding

section (see fig. 6.7) is straightforward and involves no major assumptions.

The weights of the individual components, evaluated from data on sizes and

material composition obtained from the site and the drawings, were placed

at their respective centroids and are indicated by vertical arrows in

figure 6.7 (arrows with notation are for other load cases). These include

planks used for decking and supporting brackets, diagonals and straps,

formwork and joists, railings and posts, safety nets, hydraulic actuators,

lines and jacks, stiffbacks, jacking frames, jumpform beams and miscellaneous other hardware. In the analysis, the junctions of the lower deck

brackets and suspended outer straps (fig. 6.7) were assumed pinned so

that with the exception of the diagonally braced bracket at the exterior

of the shell (second deck from the top), the system was rendered statically determinate. This bracket, together with the two diagonals was

treated as a pinned truss with one degree of redundancy. The final forces

on the shell resulting from case 2 loading are tabulated in figure 6.6.

47
6.5 Live Loads - Load Case 3

Live loading consisted of the weight of 51 workers, reinforcing bars

Georgia buggies, electrical and welding equipment, gas cylinders, water

drums, tools, grouting materials, fresh concrete, miscellaneous hoses,

wires, cables and other hardware. The live loads were applied equally

to the top decks of the interior and exterior scaffolds (designated by

in fig. 6.7) and were assumed to be uniformly distributed around the

periphery of the shell. After examining all the evidence at hand (field

observations at the Willow Island, W. Va. and Berwick, Pa. sites, plus

worker statements), it was concluded that live load distributions other

than those assumed cannot be reasonably justified so that large concentration of such loading occurring at any one location, including in the

vicinity of catheads no. 4 and no. 5, were ruled out. Furthermore, it

is noted by reference to figure 6.6 that the contribution of live loads

to the total load on the shell is relatively small so that variations in

live load distribution will not significantly affect the shell analysis

results. The procedure for evaluating shell forces induced by live loads

(fig. 6.6) is the same as load case 2 discussed in the foregoing section.

6.6 Weight of Cathead Assembly - Load Case 4

The free-body diagram of a cathead in the radial plane is shown in

figure 6.8. Without both the static and counterstatic lines acting

(at G and C, respectively), the cathead assembly is rotationally unstable

If, for any reason, the static line were absent, the cathead could still

retain its stability provided the resultant overturning moment due to

gravity loads about fictitious point 0 (point of intersection of cathead

legs a and b) is counterclockwise so that the counterstatic line is in

48
tension. Using the values for the gravity load components and

W and their locations specified in figure 6.8, it can be readily shown

that the corresponding overturning moment is, in fact, counterclockwise

if the cathead beam is assumed to be in horizontal alignment. It can

also be shown that it does not take much backward tilting from this

level position (and this can be manipulated by means of the chain hoist -

fig. 6.7) to cause a rotational collapse of the cathead in the clockwise

direction (outward) in the absence of the static line.

Referring to figure 6.1, it is noted that, with the cathead beam

level, the proximity of the bucket to the top deck of the inner scaffold

is sufficient to permit unloading of the concrete into on-deck Georgia

buggies by means of a chute pivoted to the bottom of the bucket (see

figure 6.9 for chute detail). It should be noted that the position of

the bucket relative to the deck can be controlled, in addition to the

chain hoist, by the special bolt on the slide plate assembly to which

the static line and the inside sheave of the hoist line are attached

(fig. 6.7). The position of this plate was established by measurement

of dimensions of the component parts found among the wreckage at the site.

The assumptions that the cathead beam was level at the time of the

collapse is based on the foregoing arguments and is further corroborated

by the alignments observed at the Berwick, Pa. site.

The gravity loads which are transmitted to the six pairs of adjacent

ribs at the six cathead locations around the shell are designated in

figure 6.7 (shown in parentheses) and figure 6.8. They consist of the

weights of the cathead legs W-^, W

2

, W^, the weight of the cathead beam

assembly W, which includes the weights of the sheaves, the weight of the

49
hoist line but not the hoisting loads, the weight of the steel skidboard

Pgg, the weight of the wide flange beam Pyp, and tension T

Q in the static

line due to its own weight. It should be pointed out that tension T

Q

depends on the stress-free length of the static line assumed. The criterion used in determining this length as well as the corresponding cable

forces T, and F

Q due to hoisting loads (load case 5) are discussed

in section 6.7.

Having established the cathead position and component forces of the

assembly, the corresponding leg reaction components at supports A and B

(fig. 6.8) and the forces F

c

, P

c and Qc (figs. 6.4 and 6.8) transmitted

by the counterstatic line were determined by statics. The final forces

on the shell corresponding to load case 4 are tabulated in figure 6.6.

6 . 7 Hoisting Loads - Load Case 5

Once the geometry of the cable profiles are defined (sec. 6.2), cable

forces and displacements induced by the hoisting loads can be determined.

To minimize the human computational effort involved, a computer program

was developed for that purpose. The program is based on the representation

of the static line by the simple cable model depicted in figure 6.10.

The model assumes the static and hoist lines to be weightless (the

weight of these lines are considered elsewhere - see sec. 6.6). The hoist

line is very nearly parallel to segment KG of the static line when the

bucket is about midway to the top (location for maximum tension in the

static line) and the program assumes it to be parallel so that force F D

acts in the same direction as KG for all locations of point K. The input

parameters are as defined in figure 6.10. The fixed input parameters

(the values of which are indicated in the figure) are chord length L,

50
coordinates Xq and Yq of support point G relative to support point B,

and the cross-sectional area A and modulus of elasticity E of the static

line. The variable input parameters are, position K of the bucket along

the static line as defined by chord length segments Lp and L^, the normal

distance of that point from the chord as defined by parameter D, and the

hoisting load Wg. Two of the values used for Wg were 2900 lb (1314 kg)

and 5800 lb (2627 kg). They represent, respectively, the weight of the

bucket assembly and concrete (calculations shown in fig. 6.9), and twice

that weight to assess the effect of a 100 percent dynamic amplification

of hoisting loads on cable forces.

Figure 6.11 shows in notation form a partial listing of the equations

used for calculating the desired output parameters which are indicated by

asterisks. Not shown are the equations for calculating the cathead leg

reaction components Ap, Bp, A^, B^. and tension F

c in the counterstatic

line (fig. 6.8), and a refinement introduced in the program to account

for changes in the output parameters resulting from the displacement of

support point G due to the elastic deformation of the cathead assembly.

Note that s Q

, the stress-free length of the static line, is an

output value. The program was used in an iterative fashion by adjusting

the input parameter D to produce the value of sQ that matched the true

stress-free length of the line. The true stress-free length is discussed below.

The cathead leg reactions and the force in the counterstatic line

are determined using the free body diagram of the cathead (fig. 6.8)

under the action of the static line force T and hoist line forces Fg

and F

q

. The cathead sheaves are assumed to be frictionless so that

51
F q

= Fq. By reference to figure 6.3, note that cathead leg reactions

A.y and are normal to the plane of the figure and occur in opposite

pairs; reactions Ap and Bp act along projections a and b of the respective inside and outside cathead legs in the plane of the figure; and

reactions A

N and B^ are zero because no intermediate loads (such as,

Wp , W2 and Wp) are acting in this case (refer to sec. 6.6). Thus, the

resultant cathead leg reactions are axial. Figures 6.12 and 6.13 should

further assist in visualization of the spatial configuration of the cathead legs, the counterstatic line, and reaction components (also see

fig. 9, appendix A).

For a given location of point K (fig. 6.10), the program outputs

two sets of results. The first set corresponds to the assumption that

point G is fixed in space. The second set is obtained by an iterative

process in which successive trials are based on the adjusted position

of G corresponding to the elastic deformation of the cathead assembly

under cable forces from the preceding trial.

Figures 6.13 through 6.15 show the analytical formulations for

calculating the movement of point G attributed to the elastic deformation of the cathead assembly under the action of the static and hoist

line forces. The expression for the displacement A

c of the counterstatic

line under tension F

c is derived in figure 6.13. In figure 6.14, small

rotational displacement geometry is used to develop the expression for

(Xq, Yq)

,

which represent the movement of point G, due to the elongation

of the counterstatic line. The axial deformations of the cathead legs

are smaller than the elongation of the counterstatic line by about one

order of magnitude, and therefore, their contribution to the movement of

52
point G is negligible. Figure 6.15 shows the formulation for displacements (Xq, Yq) attributed to the flexural deformation of the cathead beam.

As in the case of the legs, the effect of the axial deformation of the

beam is neglected so that Xq is assumed to be zero. The total displacement (XG , Yq) then is obtained by adding the displacements due to the

elongation ac of the counterstatic line at the top and flexural deformation of the cathead beam.

In this case, superposition of tension T

q

in the static line due to

self weight and tension T caused by loads Wg gives results that are not

appreciably different from the tension in the static line (at support G

on the cathead beam) that would be obtained from a solution based on

both loads acting simultaneously. A more refined analysis was performed

for a few cases to verify this fact. The refined analysis considers

simultaneously the weight of the cable and the hoisting loads using an

approach similar to that advanced by Harrison [6.1], The reasons for the

small difference between the two analyses are obvious. The distributed

weight of the static line (100 lb or 45.3 kg) is only about 3 percent

of the weight of the bucket (2900 lb or 1314 kg). In addition, with

sufficient slack in the static line (this was the case to enable unloading

of concrete at the top of the tower), differences in the results (tension

in the static line as well as tension F D in the hoist line - fig. 6.8)

from the two alternate procedures tend to disappear.

A glance at the results tabulated in figure 6.6 will indicate that

by far the major portion of the forces acting on the shell is produced

by the hoisting loads (load case 5). In particular, tension T in the

static line is the major source of the bending moments in the shell (hoist

53
line forces FD and F

Q = F^ remain nearly constant as T varies and balance

out so that they contribute mostly to axial loads F^ but not to normal

forces Fx

- figs. 6.5 and 6.8). Tension T in turn is quite sensitive to

variations in the tautness (or, conversely, in the amount of slack) in the

static line. Thus, the criterion for estimating the actual stress-free

length of the static line at the time of the collapse would affect significantly the results of the analysis.

As mentioned in section 3.2, the length of the static line of cathead

no. 4 (the line was recovered from the wreckage) was measured on two

separate occasions in the field. The measured lengths were within 2 in

(51 mm) of each other. The average measured length, after deducting the

length of the portion from grip hoist I to clevis B (fig. 6.2), was

219.17 ft (66.85 m). With the bucket at the unloading position shown

in figure 6.1 and the static line assumed tensionless but without play

for that position, the stress-free length s Q

is represented by the solid

line GKB as shown. This length can be calculated from the space coordinates of points B and G (figs. 6.2 and 6.3) with the known position

of point K (the bucket is hanging vertically down). The computed length

is 219.19 ft (67.46 m), which is almost identical to the length measured

in the field.

If a certain amount of play in the static line were allowed when

the bucket is at the unloading position, it would permit workers on the

top deck to manipulate the bucket or pull it in, if need be, without

encountering resistance from the static line. However, even without

such resistance, a worker will have to exert a pull of about 400 lb

(1780 N) to bring the bucket in a distance of 1 ft (0.31 m). From the

54
layout shown in figure 6.1, with the bucket at the unloading position,

it is apparent that concrete can be discharged without difficulty into

Georgia buggies on the deck by means of the chute projecting inward

above the top rail. Thus, there is no reason to believe that such manipulation with the bucket was necessary. Furthermore, if there were any

play in the static line at the discharge level, the line would become

slack below that level and the possibility of the bucket swaying and

impacting the scaffold while being pulled up would increase accordingly.

Conversely, with no play at the discharge position, the static line will

develop tension at lower levels and will pull the bucket away from the

scaffold thus diminishing the likelihood of severe impact.

The foregoing arguments led to the conclusion that the field measured length was the actual length being used at the time of the collapse

and, therefore, was adopted as the basis for the stress-free length used

in the calculations of hoisting loads (load case 5) and the static line

tension due to self weight (load case 4).

The analysis of forces induced by hoisting loads leads to the results

tabulated in table 6.1a. The results obtained by considering the elastic

deformation of the cathead were used to develop the forces for load case 5

tablulated in figure 6.6. For purposes of comparison only, table 6.1a also

shows computer results obtained by assuming support G to be fixed in space.

It should be noted that the forces in this table correspond to the location

of the bucket which produces maximum tension in the static line. This is

about 123 ft (37.5 m) below the cathead beam which is at variance with

the 60-ft (18.3 m) position estimated from the field data (fig. 6.3).

From manufacturer's specification and worker statements the speed at which

55
the hoist line was being pulled in was estimated to be about 10 to 12

ft/sec (3.1 to 3.7 m/s). Hence, it appears reasonable to assume that

failure was initiated when the tension in the static line was approaching

or at, its maximum, and during the next several seconds while collapse

was in progress, it ended up at the 60-ft (18.3-m) mark below the cathead

beam.

Table 6.1b tabulates forces produced by twice the weight of the

bucket which is equivalent to 100 percent impact on the hoist load (or

100 percent dynamic amplification of the hoist loads). These values are

only used to demonstrate that distress conditions would not have occurred

in the mechanical system even if there had been reason to believe that

100 percent impact conditions were present (see chap. 5).

6.8 Other Effects

The posture adopted in the foregoing derivations was to use lower

bound values for loads in situations where they were not known precisely.

Hoisting loads, for instance, were treated as though they were stationary

because not enough information was available to assess the inherent dynamic

effects of the hoisting system used. The reasoning behind this approach

was that should such loads lead to structural distress conditions (which

turns out to be the case - see sec. 7), the distinct possibility of higher

loads having occurred at the time of the collapse will not change the

final conclusions of this investigation. The following paragraphs discuss

the assumptions made with regard to dynamic loading.

56
Dynamic amplification of loads occurs as a result of construction

activities on the scaffolds (load case 2) and during hoisting of materials (load case 5). Dynamic amplification of live loads (load case 2)

is negligible because, for the most part, they are stationary and welldispersed (see sec. 6.5).

Dynamic amplification of hoisting loads occurs principally as a result

of the initial acceleration of the hoist line to bring the bucket up to

speed from at-rest position on the ground, fluctuations in the speed with

which the bucket is lifted as the speed of the drum hoist varied, uneven

winding of the hoist line on the drum, or any other changes in speed made

by the hoist drum operator for whatever reasons during the bucket’s journey

to the top. An increase in the speed of the bucket will cause a rise in

tension in both the static and hoist lines while a drop in speed will

have the opposite effect.

The dynamic loading condition induced by start-up acceleration

depends on how fast the bucket is brought up to speed and the damping

characteristics of the hoisting system. This and the other factors

noted above could possibly produce substantial amplification of hoisting loads. However, not enough information is available to make a

quantitative assessment of these effects with sufficient accuracy to

justify their use as part of the basis for arriving at the conclusion

drawn.

It should be pointed out that the path of the bucket is curvilinear

and therefore, even when moving at a constant speed, the bucket will exert

a transverse dynamic force (outward normal to the path) on the static

line. The path will be elliptic with foci at supports B and G (fig. 6.3)

57
if the cable is inextensible and nearly so if it is elastic. Placing

the bucket at midspan, assuming a speed of 10 ft/sec or 3.1 m/sec (see

sec. 6.7) and using the equation of the ellipse yields an incremental

tension of about 60 lb (267 N) in the static line, which is negligible

compared with the 4772-lb (21235-N) tension (table 6.1) used in this

investigation.

6.9 Analysis of Shell

An analysis of the cooling tower shell subjected to the construction

loads described in the preceding sections was made with SHORE-III [6.2],

a finite element program developed for the static and dynamic analysis

of axisymmetric shells. The cooling tower shell was discretized by a

series of curved ring elements, starting at the top of lift 28. Modeling

was consistent with the shell structure described in Chapter 2 of this

report. Details of the SHORE-III analysis and a verification analysis

using SAP IV [6.3] are presented in Appendix B. Construction loads were

applied to the shell model and the resultant maximum forces in the shell

were determined along with their location.

The finite element model used in the cooling tower analysis is shown

along with an expanded view of the model for the top 3 lifts, in figure

6.16. The tower structure is discretized by a series of nineteen closed

ring elements and one open type element at the bottom which models the

column supports. In order to obtain better stress distributions and

properly load the tower, the top two lifts (28 and 27) were subdivided

into eleven elements. The other element divisions were generally chosen

to account for either changes in shell thickness or changes in the modulus

of elasticity of the shell material. Points for load application were

58
located at the elevations in lifts 28 and 27 where the form system was

bolted to the tower. Figure 6.16 shows the elevations, radii, thickness,

and elastic moduli used for each element in the model. While the elastic

modulus remains constant for an element, the thickness may vary linearly

from the top to the bottom of the element as dictated by the shell geometry

Material properties (elastic moduli) used in the analysis were obtained

from the test data presented in chapter 4 of this report using the

maturity of the concrete in the various lifts on the day of collapse.

The construction loads presented in figure 6.6 are applied to the

shell model at the load points indicated as distributed line loads by

means of a Fourier Series. This distributed line loading technique is

depicted in figure 6.17 for the normal, meridional, and tangential forces,

Fx , Fy and F

z

, respectively, and for the meridional moment My. The x,y,z

coordinate system shown applies to the construction loads as described in

section 6.3 while the corresponding u, v, w system is the element coordinate system used in SHORE-III.

The loading function used in the SHORE-III analysis is developed by

first distributing the concentrated loads about the centerline of the

jumpform beam over 0.358 degree circumferentially (10 in or 254 ram at

lift 28) for load cases 2 and 3 and 1.875 degrees circumferentially

(4.3 ft or 1.31 m at lift 28) for load cases 4 and 5. The 10 in (254 mm)

distribution width is the surface contact length between the shell and

jumpform beam. This distributed load is then expanded in a Fourier Series

which applies the load at the required points around the circumference of

the tower for the particular construction load case. The larger distribution angle used in load cases 4 and 5 was chosen because it reduced the

number of Fourier series harmonics required to adequately define thd loads

59
Ideally, a 0.358 degree distribution angle should have been used for all

cases. However, when the 0.358 degree distribution angle was chosen for

load cases 4 and 5, 150 harmonics were required to produce a load function

with an acceptable shape. Increasing the distribution angle to 0.859

degree made it possible to generate a load function with acceptable shape

using only 56 harmonics; thereby, reducing computer time and cost by

about two thirds. A comparison analysis using both the 0.358 and 0.859

degree distribution angles showed that the maximum stresses differed by

only a few percent and the stress distributions were essentially the same.

Consequently, the compromise between distribution angle and computer analysis time is justified. Figure 6.18 illustrates the loading function for a

normal force applied to the shell by the jumpform beams at a cathead.

This load would then be repeated at all six (6) cathead locations in load

case 4 to produce a symmetric loading condition.

The development of a convergent Fourier Series with only a few harmonics was found to be a difficult task for load case 5 which is applied

only at cathead no. 4. This is because as the number of application points

decreases (two points or ribs for case 5) the number of harmonics required

for a convergent series increases rapidly. In addition, the computer time

required for solution of a load case is related to the number of harmonics

in the series. An investigation of the stress distributions that occurred

in load case 4 where the six cathead gantry loads are applied to the shell

indicated that the internal shell forces of interest decayed rapidly to a

small value at approximately 20° from the cathead. The SAP IV verification

60
analysis also exhibited this stress distribution (see appendix B) . Thus,

it was concluded that since the catheads are 60° apart, the loads applied

at a cathead have little influence on the internal forces at the catheads

on either side. Consequently, the loads for load case 5 were applied at

all six cathead locations instead of just at cathead no. 4.

The loading functions were developed for each construction load

(cases 2 to 5) and a separate analysis was made for each case, including

case 1 which was internally generated by the program. Since SHORE-III is

a linear elastic finite element program, the principle of superposition

applies and the results for each load case may be combined algebraically

to find the resultant stresses for any combination of the construction

loads. The effect of using a line load with a distribution angle rather

than a concentrated load is to predict stress magnitudes which are lower

than those that occur in the shell.

Table 6.2 presents the stress resultants in lift 28 at cathead no. 4

for the sum of load cases 1, 2, 3, 4, and 5 at 0 = 0° and 2°, respectively.

Figure 6.19 shows the sign conventions for the respective stress resultants.

Values at 6 = 0° occur midway between the jumpform beams to which cathead

no. 4 is attached while the values at 9 = 1.875° occur at the meridian

along which the jumpform beams are attached to the shell.

Figures 6.20 through 6.23 exhibit the distribution of the meridian

stress resultant, N , the hoop stress resultant, Nq, and the moments, MA , M , with 0 for selected elevations in lift 28 (refer to fig. 6.17) for

b the sum of load cases 1, 2, 3, 4, and 5. Because they are symmetric

about the cathead, the distributions are shown for only one side of the

cathead. It should be noted that the values of the meridional stress at

0 = 1.875°, figure 6.20 and also table 6.2, increases significantly at

61
elevations of 165.721 and 162.125 ft (50.512 and 49.416 m respectively)

due to the loads being applied at points I and J (fig. 6.6). The shape

of the stress distribution for N. about the circumference of the shell,

<P figure 6.20, can be understood by considering the way the loads are applied

to the shell. The loads prescribed in cases 2 and 3 are applied at each

rib and the resulting stress distribution is symmetric about each rib and,

furthermore, is identical between all ribs. On the other hand, the loads

prescribed in cases 4 and 5 are applied only at the cathead locations.

The stress distributions illustrated in figure 6.20 reflected the superposition of these cases and the stresses due to load cases 4 and 5 control

the shape of the distributions shown. The hoop stress, N^, as seen from

figure 6.21 is highest at the top of lift 28 and decreases with decreasing

elevation. Finally, the maximum values for N, and N. occur near the top

(J) u bolt in lift 28 at the jumpform beams on either side of cathead 4. A

relative maximum for N, occurs near the bottom bolt in lift 28 at the

<P jumpform beams. The maximum moments M, and MQ also occur at the location

<p u of the jumpform beams (0 = 1.875°) with Mg a maximum near the top of lift

28 and M a maximum near the bottom of lift 28 as shown in figs. 6.22 and

$

6.23.

62
7. INTERPRETATION OF SHELL ANALYSIS

7 .1 Introduction

The overall safety and the reliability of the construction method

used at the Willow Island site depend on the integrity of the partially

completed shell structure to resist all construction loads. This

demands that sufficient strength must be developed in the previously

cast shell so that the structural resistance exceeds the construction

load effects by a reasonable margin of safety. Obviously, the shell did

not have adequate strength to resist the applied load. In order to determine to what extent the applied load effects exceeded the capacity of the

shell, the results of the analysis of the shell at a number of critical

locations are compared with the resistance values determined by applying

existing strength theories. It should be pointed out that no dynamic

amplification of the hoist load was included in the shell analysis.

Dynamic amplification is known to exist in hoisting systems such as the

one used at the Willow Island site, and inclusion of the dynamic effects

would have increased the forces in the shell over those determined in

this report.

7 . 2 Strength of the Shell Under Combined Axial Load and Bending Moment

For a given cross section and reinforcement, an interaction diagram

such as the one shown in figure 7.1 can be constructed in terms of ultimate axial forces as ordinates and ultimate bending moments as abscissa.

For various combinations of axial forces and bending moments, the interaction curve defines a failure envelope in that all points lying on and

outside the interaction curve constitute failure. On the other hand,

combinations which fall within the area bounded by the interaction curve

63
represent a safe condition. In this section a number of cross sections

in the critical region of the shell are examined using the interaction

diagrams. It was shown previously in the analysis of the shell that

maximum forces in lift 28 would occur in the vicinity of the two ribs

where the legs of cathead no. 4 were positioned. This critical region

includes the section between the two ribs as well as the immediate

vicinity along the ribs.

In developing the interaction diagrams for the critical cross

sections, a shell section of unit width (1 ft) was treated as a rectangular column. In an actual shell element the forces and moments occur

in adjacent cross sections parallel and perpendicular to the meridian.

By treating a shell element as a column-like element, the forces and

moments in each direction are dealt with separately. Any effect on the

compressive strength of concrete due to biaxial compression is small [7.2].

Furthermore, the forces acting normal to the reinforcement have little,

if any, effect on the stress that can be developed in the reinforcement.

Thus, the treatment of a shell element as an uniaxial element is considered

a reasonable and expedient approach.

The cross sectional dimensions of the shell sections used for determining interaction diagrams are shown in figure 7.2. The vertical section

has two no. 4 bars and the horizontal section has two bars with an equivalent steel area for a 12-inch (305 mm) wide cross section. The actual

shell cross section in lift 28 had # 4 vertical bars spaced at 8.7 in

on center (220 mm) on each face and #4 horizontal bars spaced at 12 in

(305 mm) on center on each face (see fig. 2.2).

64
Because the strength-gain characteristics of the lift 28 concrete as

shown by the results of cylinder tests were significantly different from

that of the NBS specimens (fig. 4.3), it was decided to use the lift 28

test data to estimate the concrete strength. The strength of concrete at

the time of the collapse, approximately 20 hr after placement of concrete

(sec. 4.2), is estimated to be 220 psi (1.52 MPa) based on the results of

24 and 25 hr tests, 283 psi (1.95 MPa) and 299 psi (2.06 MPa), respectively

(Table 4.5 a and b). The procedure followed to establish the concrete

strength is shown in figure 7.3. It is seen that a linear interpolation is

made between the time of set (6 hr) and the time when the field cured

cylinders were tested. The time of set was established by the experimental

results with cement mortar, with consideration given to the inclusion of

flyash and water reducing admixture in the lift 28 concrete. Approximately

the same value of the compressive strength can be obtained from the

strength-maturity plot shown in figure 4.3 with a value of 43 degree Fdays being the maturity at the time of collapse. It should be pointed

out that the actual concrete strength at the top of lift 28 could have

been lower than 220 psi (1.52 MPa) due to the temperature condition of

lift 28 was different than that of the test cylinders which were cured

on the ground level and due to the rise of free water to the top of the

wall both of which tend to lower the concrete strength. If these factors

were taken into consideration, the concrete strength could have been as

low as 200 psi (1.38 MPa) rather than 220 psi (1.52 MPa). However, in

the evaluation of cross-sectional strengths, 220 psi was used.

The reinforcing steel must be embedded adequately in the concrete to

develop its yield strength. The length of embedment required to develop

65
yield is proportional to the square root of the strength of concrete.

According to ACI 318-77 [4.4], at a strength level of 220 psi (1.52 MPa)

the embedment length for a no. 4 bar in compression is 40.5 in (1029 mm).

Because the ACI code requirement incorporates a 25 percent increase in

embedment length over that required to develop yield [7.1], the value

obtained from the ACI code should be reduced by 25 percent to reflect

the actual needed embedment length. Thus, the corrected compression

embedment length for a no. 4 bar becomes 32.4 in (823 mm).

The maximum amount of stress that can be developed in the steel was

computed based on available length of embedment. The bars in the circumferential direction, which are in compression, could develop full yield

strength of 60 ksi (414 MPa) except at lap splices. At the locations

where lap splices occur the maximum stress that can be developed in the

/ bar is limited by the length of splice. For the splice length of 22 in

(559 mm), which was provided for no. 4 bars (fig. 2.2), the maximum compression stress in the steel is limited to 60 ksi x 22/32.4 = 40.7 ksi

(281 MPa). For the bars in the meridian direction the available length

of embedment varies depending on the distance from the top edge of lift 28

to cross sections being considered. For a cross section located 32.4 in

(823 mm) below the top of lift 28, the steel can develop to yield (60 ksi,

414 MPa). Under combined bending moments and axial forces, some bars

were in tension. For these bars embedment length of 28.8 in (732 mm) was

computed using the pullout test results. However, at all locations where

the tension embedment length would be a governing factor, the magnitude

of these moments and axial forces are relatively low and their combined

effects fall well within the interaction diagram.

66
For the shell cross sections of unit width shown in figure 7.2 and

using the limiting concrete and steel stresses described above, interaction

diagrams for axial load and bending moment are obtained for critical locations in lift 28 (fig. 7.4). In developing the interaction diagram, it

was assumed that the crushing strain of concrete is 0.003 in per in^ and

that the reinforcing steel has an elastic-perfectly plastic stress-strain

relationship with a nominal yield strength of 60 ksi (414 MPa).

Figure 7.5 shows an interaction diagram for a cross section resisting

meridional forces (hereafter referred to as the horizontal cross section)

for all cross sections 32.4 in (823 mm) below the top of lift 28. Below

this point a sufficient embedment length is available for the vertical

bars to develop yield. In lift 28 the meridional force and bending moment

are maximum at cross sections along a line through points BAB as shown in

figure 7.4. Maximum force and moment combinations for cross sections along

this line are plotted in figure 7.5 in which the point corresponding to

location A is shown as a square and the point corresponding to location B

as a circle. It is to be noted in this figure that at both locations the

cross section is controlled predominantly by bending moment. Because the

cathead gantry loads are supported at the rib locations, the bending moment

is substantially greater at this location than at the center of the panel,

exceeding the capacity of the shell cross section. This indicates that a

compression failure of concrete would initiate at the inside of the shell

^In rectangular beam tests, strains 0.003 to 0.004 have been measured

near maximum load carrying capacity. Many tests of beams and columns

have shown that a satisfactorily accurate prediction of ultimate

strength can be made using the crushing strain of 0.003 [4.4], Limited

test data are available for concrete strength of 220 psi (1.52 MPa).

However, a more accurate determination of the concrete strain at

maximum resistance was not made in this investigation because the

effect on the interaction diagram is not large.

67
at the rib locations and propagate along the circumference at a distance

1 ft (0.3 m) above the bottom of the panel.

The locations of lap splices of the horizontal bars were not known

for the cross section resisting hoop forces (hereafter referred to as the

vertical cross section). Hence, two interaction diagrams were prepared

depending on the maximum stress that can be developed in the steel. If

lap splices of the bars are a sufficient distance away from a section

being considered, the steel can develop yield (60 ksi, 414 MPa). For

cross sections where lap splices occurred, the maximum stress that can be

developed is limited by the length of the splice (40.7 ksi, 281 MPa).

The interaction diagrams obtained using these two steel stresses are shown

in figure 7.6. These diagrams give the combinations of the hoop force

and bending moment for failure of the cross section. In lift 28, the hoop

force and associated bending moment are maximum at locations C and D. They

are plotted on the interaction diagram for these sections in figure 7.6

in which the point corresponding to location C are shown in square and the

corresponding to location D in circle. It is seen in this figure that with

point D lying considerably outside of the interaction curve a compression

failure of concrete at the top of the shell would take place at the rib locations, points D, due to combined effect of high moment and axial force with

or without the presence of lap splices. At the center of the panel, the

shell section is subjected predominately to axial force.

7.3 Strength of Shell Under Shear

In the presence of axial compression the shear capacity of a reinforced concrete section increases. This complex interaction between shear

and axial compression is not fully understood. In the absence of a suitable means of determining the shear strength of a shell section, the

68
expressions 'in the ACI code [4.4] are used for insight into the strength

of the shell in shear. Those expressions are:

Vc - 2 1 +

2000 A

g

and not greater than

vc “ 3. 5wl + 4 500 A,g

where

V

Q = nominal shear strength provided by concrete

N

u

- axial normal load to cross section

Ag = gross area of section

f'

c = compressive strength of concrete in psi

b = width of section

d = distance from extreme compression fiber to centroid of

tension reinforcement •

Radial shear values in the vertical and horizonal cross sections of

elements at the top of lift 28 are given in table 7.1 together with corresponding shear capacities computed in accordance with the above formulas.

It may be seen in the table that the radial shears in the vertical and

horizontal sections (Qq and Q.) are very high at the top of lift 28 along

u (p two ribs, while the radial shears at the center of the panel are either

zero or very small. The radial shear in both the vertical and horizontal

sections in the region near the top of lift 28 along the ribs, points D in

figure 7.4, exceed those values computed by the formulas. Although the

calculated shear values may be influenced by the distribution of concentrated loads along a line, as discussed in section 6.5, this clearly indicates

that relatively high radial shear forces were present in this region of the

shell.

69
7.4 Probable Mode of Failure

It was shown in the previous sections that maximum stresses occur in the

region of the panel bounded by the two ribs where cathead gantry no. 4 was located.

The analysis showed that a compression failure of concrete would have initiated

in lift 28 along the circumference of the shell at a distance 1.0 ft (0.3 m) above

the bottom of lift 28, the line through points BAB in figure 7,4. In addition,

because both the large hoop forces and moments along the ribs, a band of the

compression failure would also have spread along the rib of the panel. The

presence of high radial shear in the panel could have further weakened the shell

where combination of high axial loads and moments occurred and could have initiated the failure before the capacity in axial forces and moments was reached.

Therefore, it may be concluded that the failure of lift 28 resulted

from inadequate strength of the shell section where cathead no. 4 was

located to resist the applied construction loads. The failure was brought

about by compressive crushing of the concrete due to combinations of

axial forces and moments, and/or cracking due to the high radial shear.
8 SUMMARY AND CONCLUSIONS

As a result of the cooling tower construction collapse at Willow

Island, West Virginia, the National Bureau of Standards (NBS) was requested

by the Occupational Safety and Health Administration (OSHA) to carry out

a detailed study aiming at the determination of the most probable cause

of the collapse. In response to this request, NBS has undertaken a comprehensive field investigation, laboratory tests of construction assembly

components and concrete specimens, and chemical analyses of concrete.

In addition, mathematical models of the tower were prepared and analyzed

by computer using shell analysis program. The findings presented in this

report are based on the results of these field, laboratory and analytical

studies with other information such as structural drawings of the tower

and OSHA case records.

1. At the time of failure the concrete bucket was in transit

from the base of the tower to cathead no. 4. The measured

length of the hoist cable indicated that the bucket was about

60 ft (16.4 m) below the cathead beam. This also agrees with

eyewitness accounts. Therefore, it is believed that the concrete bucket did not hit the cathead to cause it to fail.

2. Although the hoist cables for catheads no. 4 and 5 were broken,

field observations and the laboratory test indicate that the

breakage of the cables occurred after the onset of collapse

of lift 28. Thus, the breakage of the cables did not trigger

the failure.

71
3. Field and laboratory investigations revealed that the major

components of the hoisting, scaffolding and formwork systems

did not fail prior to the collapse. Thus, the collapse did

not initiate due to any component failure of these systems.

4. Based on the results of field cured cylinder tests, the compressive strength of concrete of that part of lift 28 where

cathead no. 4 was located was estimated to be about 220 psi

(1.52 MPa) at the time of the collapse.

5. The analysis of the shell as well as eyewitness accounts

indicated that the collapse initiated in lift 28 at the area

where cathead no. 4 was located. The analysis showed that

calculated stress resultants at several points in this area

equaled or exceeded the strength of the shell in compression,

bending and shear. Failure at any of these points would have

propagated causing the collapse of lift 28.

6. These results of the analysis indicate that the most probable

cause of the collapse was due to the imposition of construction

loads on the shell before the concrete of lift 28 had gained

adequate strength to support these loads.

72
9 . ACKNOWLEDGMENTS

The authors acknowledge the contributions made by the following

persons

:

Dr. E. 0. Pfrang provided technical guidance and constructive criticism.

Dr. E. V. LeyendeCker, Mr. J. R. Harris and Dr. B. Ellingwood critically

reviewed the report.

Mr. R, Williams coordinated the laboratory tests and Messrs. F. Rankin,

J. Owens and T. Ruschell performed the tests.

Mr. C. Mullen coordinated graphic work and prepared figures.

Mrs. A. Salazar typed the several drafts of this report manuscript.

Mr. Ray Mele and Mr. John Day carried out the graphic work.

73
10. REFERENCES

[4.1] 1978 Annual Book of ASTM Standards, Part 14, American Society

of Testing and Materials, Philadelphia, PA 19103.

[4.2] 1978 Annual Book of ASTM Standards, Part 13, American Society

of Testing and Materials, Philadelphia, PA 19103.

[4.3] Lew, H. S. and Reichard, T. M., "Mechanical Properties of

Concrete at Early Ages," Journal of the American Concrete

Institute, Vol 75, American Concrete Institute, Detroit,

Michigan, October 1978.

[4.4] ACI 318-77, "Building Code Requirements for Reinforced Concrete,"

American Concrete Institute, Detroit, Michigan, 1977.

[6.1] Harrison, H. B., "Iterative Nonlinear Structural Analysis,"

Journal of Structural Division, American Society of Civil

Engineers, New York, NY, July 1976.

[6.2] Basu, P. K. and Gould, P. L., "SHORE-III - Shell of Revolution

Finite Element Program Users' Manual," Research Report No. 49,

Structural Division, Washington University, St. Louis, September

1977.

[6.3] Bathe, K. J., Wilson, E. L. , and Peterson, F. E., "SAP IV -

Structural Analysis Program for Static and Dynamic Response

of Linear Systems," Earthquake Engineering Research Center

Report No. EERC 73-11, College of Engineering, University of

California, Berkeley, April 1974.

*> [6.4] Winter, George, Urquhart, L. C., O'Rourke, C. E., and Nilson,

Arthur H., "Design of Concrete Structures", McGraw-Hill Company,

New York, Y.Y.

,

1973.

74
[7.1] Orangun, C. 0., Jirsa, J. 0. and Breen, J. E., "A Reevaluation

of Test Data on Development Length and Splices," Journal of

the American Concrete Institute, Vol. 74, American Concrete

Institute, Detroit, Michigan , March 1977.

[7.2] Kupfer, H., Hilsdorf, H. K. , and Rusch, H., "Behavior of

Concrete Under Biaxial Stress," Journal of the American

Concrete Institute, Vol. 66, American Concrete Institute,

Detroit, Michigan, August 1969.

[8.1] Department of Labor, Bureau of Labor Standards, "Safety and

Health Regulations for Construction," (29 CFR part 1926),

Washington, D.C.

75
TABLES AND FIGURES

76
Table

4.1

Physical

Properties

of

Type

II

Cement

Samples

OSHA

Sample

3750

CO

PQ

$3 0) rHa i 00 i i i i S i CO i i i i 03 i i i i i CO i co i i i i CO

/~N /—S

d P

<u VO rH •H

r1 • • s a

60 Cl rH CO p S rH rH

O o

cd CM o •w/ v«/ rH

m

u CO

a CM

co o o u

<u <5 CO 00 o Si SI H cc VD Oh

CO rH CO <r

60 o Cu

a> P c

a) /-x CN •H P CD 00 CM B e

*H rH • 8 60 a in Oh

m

dg rH <3- <*

w cd

O

CO • rH o o U

u 3 r>. o Si

cti CO CO rH CM

rJ PQ CM CM CSS SZ

u

U sio

rH rH P c «d

(8 o /*> hC JS m CO CM 4J 4-1 rH ® • i o CO 01 cj CM rH rH CO u

rH S-/ 0) o X rH a H * o o o

CO

X o o o 4-> 4J < eg 00 m in O o X CM rH CM X a

4= Ph

4J S K 00^ *

a>d «K

a Q) -H • cd Vh CO P 14-1 4-J CL

d ^ /-N •H u CO •H JJ

a; 0 p * 4-1 CO

4-» 0 CO (U HC 4-1 0) 4-J d •K > « <u h Q)

CO

V rH « o •H In CO CO 4-)

4->

u o CO •H co cd a) 0)

CO

p > CO MH ^ CO 4-J 4-1H rH CO

Q o a; *H 00 0)M CO CO o o cd

a

d o \ V4

O to So 0 •rH rH

rH 4Q 0) axM a. S cd cd a) rH 4-1 cd

cdu c Cu 0 0 TO TO 0 *H •H

d

a •H •H (0 CJ

O CM •H o d •H

•H pH 'w' CJ o CO

H x-

HH PH

CO 40 c • • •

P-4 rH CM CO

77

ASTM C

204

[4.2]

ASTM C

109

[4.2]

ASTM C

266

[4.2]
Table 4,2 Chemical Properties of Type II Cement

Chemical

Elements

ASTM

C 150

%

Wet Chemical Analysis

NBS Sample

%

OSHA Sample

%

Silicon Dioxide Min. 21.0 20.6 20.0

Aluminum Oxide Max . 6.0 5.3 5.2

Ferric Oxide Max. 6.0 3.0 3.9

Calcium Oxide — 64.1 63.5

Magnesium Oxide Max. 6.0 2.4 2.1

Sulfur Trioxide Max. 3.0 2.1 2.0

Total Alkalies — 0.38 0.44

Loss on Ignition Max . 3.0 1.2 1.6

Insoluble Residue Max. 0.75 0.22 0.25

Tricalcium Silicate — 60 60

Dicalcium Silicate — 14 12

Tricalcium Aluminate Max. 8.0 7.1 7.1

Tetracalcium Alumiof errite 9 12

78
Table 4.3 Element Analysis of Flyash

Element

NBS Sample

(in %)

OSHA Sample

(in %)

Aluminum 12.8 11.3

Boron 0.05 0.05

Barium 0.15 0.20

Carbon 0.50 1.30

Calcium 2.00 3.00

Chromium 0.01 0.01

Copper 0.01 0.01

Iron 9.8 10.5

Potassium 2.00 2.00

Lithium 0.15 0.12

Magnesium 2.00 2.00

Manganese 0.05 0.05

Sodium 0.30 0.30

Nickel 0.01 0.01

Rubidium 0.01 0.01

Silicon 20.9 20.7

Strontium 0.07 0.10

Titanium 0.50 0.50

Table 4.4 Chemical Composition of Fly Ash

NBS Sample OSHA Sample ASTM C 618

Aluminum Oxide 24.1 21.3 , Iron Oxide 14.0 14.9 —

Silicon Dioxide 44.6 44.2 —

Sum 82.8 80.5 70.0

79
Field

Data

On

Concrete

Compression

Test

Specimens

Field

Data

On

Concrete

Compression

Test

Specimens

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OUT

AS

FAR

AS

POSSIBLE

FILL

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FAR

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LABORATORY,

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-,

i? 1 | ! \

•| j VJ! s ,Cs ct v t • - 2 C i c 3 if >Oo £ 2 <V» i - r*

*- * 0 - t*l s : Vi 1

lA) r• X

0| l X

0)

IS 1 J: s

V) P ^ N <3? $ ® <

i| 3

£3 1 i> j_rt C_3

CD 3 ua-o o

t*- i 4-j u tre2

fn

tss F\* jir • & c i u

SfJ

5. £ ; 0s 000 K t Ry n s i2* 5M

V*»

0,

*.> aa

Os

f?

ty

*\

k

ts 5)S' s L; ~

A'!

M£. *1 • •

'c.< S, 1 V ! 1 & *v • ••S' *N l Ev A - SV V < : >.-e i?, £ ;

eT.I S) < M i T- iV (1 • sS t Lr . v.

ji. tN $ N <

SY) ij 55

E» jii

c. w

u

t'O

c. ra

vS

u fO TO-a A'' t

'* < .’

xfeUx S-. .

80
Table 4.6 Data on Fresh Concrete

Range Average

1. Unit Weight of Concrete 140.5 - 144.4 143.1

lb/ft3 (kg/m 3

) (2251 - 2313) (2292)

2. Slump 6-9 7.4

in (mm) (152 - 229) (188)

3. Air Content

%

4.5 - 5.9 5.1

4. Temperature of Concrete 66.3 - 66.9 66.4

at the Time of Pour

°F (°C)

(17.2 - 20.6) (19.1)

Table 4.7 Results of Tension Tests of Cables

Hoist End Section Drum End Section

1 28.53 kip (126.91 kN) 26.55 kip (118.10 kN)

2 26.82 (119.30) 27.00 (120.01)

3 27.20 (120.99) 26.20 (116.54)

4 27.75 (123.44) 27.45 (122.10)

5 27.40 (121.88) 27.78 (123.57)

Average 27.54 kip (122.50 kN) 27.00 kip (120.01 kN)

c

81
Table 6.1 fathead Forces and

(for load case 5)

Reactions Induced by Hoisting Loads -

(a) (b)

Impact Factor 0% 100%

ib\g)

2900

(1315)

5800

(2630)

ft ?m)

219.17

(66.80)

219.17

(66.80)

Condition of

Support G Fixed Elastic Fixed Elastic

F

0

lb (N)

2460.14

(10943.24)

2451.69

(10905.66)

4928.35

(21922.31)

4931.73

(21937.42)

T

lb (N)

5378.64

(23925.37)

4772.26

(21228.06)

8860.59

(39413.85)

7670.59

(34120.47)

F

c lb (N)

11311.67

(50316.80)

10131.61

(45067.63)

18760.19

(83449.45)

16339.25

(72680.58)

A

p

lb (N)

18630.78

(82873.81)

16913.27

(75233.95)

31858.23

(141712.41)

28461.90

(126604.79

B

P lb (N)

296.91

(1320.72)

496.14

(2206.94)

1213.33

(5397.15)

1602.59

(7128.67)

A

Z lb (N)

3453.00

(15359.70)

3134.69

(13943.79)

5904.57

(26264.83)

5275.10

(23464.81)

B Z lb (N)

57.71

(256.71)

96.44

(428.99)

235.84

(1049.07)

311.51

(1385.67)

L

1 ft (m)

95.0

(29.0)

100

(30.5)

95.0

(29.0)

100

(30.5)

L

2 ft (m)

123.56

(37.66)

118.56

(40.41)

123.56

(37.66)

118.56

(40.41)

82
Table 6.2a Stress Resultants in Lift 28 for Sum of Load Cases

at Angles of 0 and 1.875 Degrees

Sum of Load Cases; 1 2, 3, 4, 5 6=0°

Point of

Load

Application Location

Top

of

Element Node

N

* k/ft

N

0 k/ft

N

e* k/ft

M

* in-k/f

t

M

e in-k/f

t

%

k/ft

Q

6 k/ft

TL 28 1 1 0.000 -27.11 0.0 0.000 -2.449 0.0000 0.0

I 2 2 0.034 -25.81 0.0 1.912 -1.255 0.0295 0.0

K 3 3 -0.047 -22.95 0.0 8.804 -3.632 0.7103 0.0

4 4 -0.114 -18.64 0.0 23.532 -7.699 0.9382 0.0

5 5 -0.127 -14.94 0.0 36.840 -10.286 0.9345 0.0

L 6 6 -0.099 -11.72 0.0 47.892 -11.707 0.8683 0.0

J 7 7 0.023 - 9.46 0.0 54.312 -11.344 0.9559 0.0 TL 27 8 8 0.102 -13.18 0.0 60.768 -24.696 0.4627 0.0

Sum of Load Cases; 1, 2, 3, 4, 5 e = 1.875°

Point of

Load

Top

of

N

e N

e* M

* M e Q* Q

e Application Location Element Node k/ft k/ft k/ft in-k/f

t

in-k/ f t k/ft k/ft

TL 28 1 1 -0.001 -31.91 0.000 0.000 -76.416 0.000 4.3880

I 2 2 -5.341 -28.28 -1.122 -8.732 -81.792 8.594 1.4400

K 3 3 -3.644 -22.59 -1.276 18.492 -62.628 4.707 -0.5434

4 4 -1.869 -17.64 -1.362 35.400 -34.080 3.234 0.2383

5 5 -0.814 -13.97 -1.671 49.908 -13.788 3.074 0.1788

L 6 6 0.374 -10.94 -1.343 70.644 3.307 2.322 -1.5720

J 7 7 -3.811 - 9.06 -0.792 89.520 13.704 -3.506 -0.5855

TL 27 8 8 -2.012 -12.28 -1.387 62.556 5.300 -1.526 1.6810

83
Table 6.2b Stress Resultants In Lift 28 for Sum of Load Cases

at Angles of 0 and 1.875 Degrees

Sum of Load Cases: 1, 2 , 3, 4, 5 6 = 0

°

Point of Load

Application Location

Top

of

Element Node

N

* KN/m

Ne KN/m

N

e* KN/m

M

* KN-m/rn

M

e KN-m/m

Q*

KN/m

%

KN/m

TL 28 1 1 0.002 -395.7 0.00 0.000 -0.908 0.000 0.00

I 2 2 0.493 -376.7 0.00 0.708 -0.465 0.431 0.00

K 3 3 -0.679 -335.0 0.00 3.263 -1.346 10.368 0.00

4 4 -1.667 -272.1 0.00 8.721 -2.853 13.694 0.00

5 5 -1.858 -218.1 0.00 13.653 -3.811 13.640 0.00

L 6 6 -1.438 -171.1 0.00 17.749 -4.339 12.674 0.00

J 7 7 0.342 -138.1 0.00 20.128 -4.204 13.952 0.00 TL 27 8 8 1.486 -192.4 0.00 22.521 -9.152 6.754 0.00

i

Sum of Load Cases: 1, 2, 3, 4, 5 9 - 1.875°

Point of Top N

0 N

e* M

* M

e Q* Q

e Load of

Application Location Element Node KN/m KN/m KN/m KN-m/m KN-m/m KN/m KN/m

TL 28 1 1 -0.01 -465.8 0.00 0.00 -32.03 -0.00 64.05

I 2 2 -77.95 -412.8 -16.38 -3.24 -30.31 125.40 21.02 K 3 3 -53.19 -329.7 -18.62 6.85 -23.21 68.70 -7.93

4 4 -27.28 -257.5 -19.88 13.12 -12.63 47.20 3.48

•\ 5 5 -11.88 -203.9 -24.39 18.50 - 5.11 44.67 2.61

L 6 6 5.47 -159.7 -19.60 26.18 1.23 33.89 -22.95

J 7 7 -55.63 -132.3 -11.56 33.18 5.08 -51.17 -8.55 TL 27 8 8 -29.37 -179.2 -20.25 23.18 1.96 -22.27 24.54

84
Table 7.1 Shear Forces in Lift 28

Loading Cond.

1,2, 3, 4,

5

Q 0 Shear

Strength

(in pounds)

Shear

Strength

(in pounds)

At Center 0 2539 30 2225

At Rib 4388 2594* 8594 2286*

* Capacity of the section is less than shear force.

Qq = Radial shear in the vertical cross section

= Radial shear in the horizontal cross section

1.0 lbf = 4.44822 N

85
Figure 1.1 A General View Showing Completed Tower Unit No. 1 and Unit No. 2

86
40 PEOESTAL LOCATIONS

80 SUPPORTING COLUMNS

1 ft = 0.305 m

1 in = 25.4mm

Figure 2.1 Elevation and Plan View of Cooling Tower Unit No. 2

87
4 - 1

" MIN LAP FOR #8

2’ 9” MIN LAP FOR #6

2 -3" MIN LAP FOR #5

(a) Bar Sizes and Spacing

( b ) Splice Detail for Horizontal Bars for Lift 28

H

I

PANEL "C" PANEL "A" PANEL "B" PANEL "C" PANEL "A"

1 ft = 0.305m

(c) Splice Detail for Vertical Bars 1 in = 25.4mm

Figure 2.2 Detail of Wall Reinforcement

88
Jumpform beams

Cathead gantry

>

Interior

> scaffold

Figure 2. A Interior View of a Tower Under Construction
Cathead gantry

Figure 2.5 A Cross Section through the Formwork and Scaffolding Systems

90
Figure 2.6 Attachment of Jumpform Beam to the Exterior of Shell

91
Figure 2.7 Jump form Beam Detail

92
INTERNALLY THREADED CONNECTOR GOING INSIDE AND OUTSIDE JUMPFORM BEAM PAIRS

Figure 2.8 Attachment Detail of Jumpform Beams and Jacking Frames

93
Figure 2.9 Jacking Frame Assembly

94
Figure 2.10 Formwork Support System

(Note five levels of adjustable joists are used per lift.)

95
Inside jumpform beam

Elevation 156.283*

i Diameter 27L142*^

Figure 2.11 Position of Jumpform Beams prior to Concrete Placement for Lift 27

96
Figure 2.12 Raising of Formwork and Relocation of Jumpform Beam after Casting

of Lift 27

97
Elevation 161.125' Diameter 268.645'

ty

Elevation 156.289*

Diameter 271.142' ty

Figure 2.13 Position of Jumpform Beams Prior to Concrete Placement for Lift 28

98
Figure 2.14 Raising of Formwork and Relocation of Jumpform Beam after Casting

of Lift 28

99
Elevation 165.969' i Diameter 266.167' '

Elevation 161.125'

Diameter 268.645'

&

Elevation 156.289'

Diameter 271.142'

Figure 2.15 Position of Jumpform Beams Prior to Concrete Placement

for Lift 29

100
Slide plate assembly

Figure 2.16 Cathead Gantry Assembly

101
Cathead

crane

o

102

Figure

2.17

A

Schematic

View

of

the

Hoisting

System
Twin drum hoist

Figure 2.18 Location of Drum Hoist Relative to Cathead Gantries

103
Figure 3.1 Exterior View of Tower Unit No. 2

'

Hoist cable for cathead no 1

if

104
Figure 3.3 Debris Piled up Inside the Tower

Figure 3.4 Distribution of Debris Around the Base of the Tower

105
Figure 3.5 Top of Lift 27 at Cathead Gantry No. 1

location of cathead

gantry no 2

Part of concrete broken

away with upper

anchor bolt |

Lower anchor

bolt

Anchor bolts

Anchor bolts

Figure 3.6 Exterior View of Top Portion of the Shell Between Cathead

Gantry No. 1 and No. 2

106
Figure 3.7 Typical Rib in Lift 27

Mexhead bolt

Threaded

coupler

Figure 3.8 Anchor Bolt Assembly

107
%

Localized crushing

Figure 3.9 A Close-up View of the Bottom Bolt in Lift 27 Which

Supported a Cathead Gantry

Figure 3.10 Close-up View of Top of Lift 27 at a Rib Location

108
Figure 3.11 Close-up View of Top of Lift 27 Showing an Anchor Bolt

Hole in Concrete

Figure 3.12 An Overall View of Cathead Gantry

109
Figure 3.13 Slide Plate for Cathead Gantry No. 4

Figure 3.14 Displaced Slide Plate of Cathead Gantry No. 6

110
Figure 3.16 Outside Legs of Cathead Gantry No. 4

111
Figure 3.17 A Groove in the Shell Cut by the Hoist Line at

Cathead Gantry No. 4

Figure 3.18 Concrete Basket Used for Cathead Gantry No. 4 Which

Was Recovered from Wreckage

112
Figure 3.19 Drum Hoist for Cathead Gantry No. 3 and 4

Time (t), days

Figure 4.1 Temperature Variation Used to Cure Test Specimens

113
Figure 4.2 Compressive Test of Cylinder

114
°C

-

day

27.8

55.6

277.8

555.6

1111.2

2777.8

;sd '(aj) qiSuaJiS aAissajduiOQ

>

115

Figure

4.3

Compressive

Strength

vs.

Maturity
Secant

Modulus

(E) @

0.4

f'c,

106

psi

V"kPa

0 20 40 60 80 100 120 140 160 180

Figure 4.4 Compressive Strength vs. Modulus of -.lasticity

116

MPa
117
-f

i*H> es*

o

LA

CO

CM

ULUS

Square

root

of

compressive

strength

in

psi

[\]Vc)
Figure 4.8 Williams Anchor Bolt Assembly

Figure 4.9 Appearance of Williams Anchor Bolts After Tension Test

119
120
Figure 4.11 Typical Wire Rope Test Specimen

Figure 4.12 Tension Test Setup for Wire Rope

121
Figure 4.14 Fractured Cover Plate for Gear Box

Figure 4.13 Test Setup for Chain Hoist

122
Figure 4.15 Anchorage of Static Line to Ground Through Grip-Hoist

123
Figure 4.16 Test Setup for Grip-Hoist

Figure 4.17 A Close-up View of Grip-Hoist After Tension Test.

The Hook Pulled Away from the Housing.

124
125
SLIDE PLATE ASSEMBLY

Figure 6.1 General Sectional Configuration of Construction Scheme

126
TOP OF LIFT 27

FIELD SURVEY DATA

C: cathead no. 5 static line clevis at concrete hopper Elevation of C - 105.37 ft (32.13 m)

BD // DE II EF X AD

Horizontal Distances, ft (m)

BD DE AD EF 8.4583 6.5417 1.0625 2.0 (2.5798) (1.9952) (0.3241) (0.61) CE CF CT AF 2.7708 1.5417 92.5625 8.6075

(0.8451) (0.4702) (28.2316) (2.6253)

Coordinates, ft (m)

Location and Description of Point

Space Coordinates

North South Elevation

A: Center Marker 919.44

(280.43) 1023.39(312.13)

97.81

(29.83)

B: Cathead 4 Static 927.94 1022.66 104.92

Line Clevis at Hopper (283.02) (311.91) (32.00)

T: Sheave Block for Hoist 1000.00 1000,00 98.45 | Lines of Catheads 3 and 4 ! (305.00) (305.00) (30.03) 1 N: Cold Joint Top of 1009.23 923.50 264.69

Lift 27 at Cathead 4 (307.82) (281.67) (80.73)

U: Hoist Drum for 1079.80 976.84 99.53

Catheads 3 and 4 (329.34) (297.94) (30.36)

Q: Ground Sheave for Holst * 1040.92 687.84 98.61 Line of Cathead 4 (317.48) (270.79) (30.08)

V: Sheave Block for Hol6t 852.22 973.94 98.49

Lines of Catheads 5 and 6 i (259.93) (297.05) (30.04)

W: Cold Joint Top of 869.48 898.75 265.54

Lift 27 at Cathead 5 (264.58) (274.12) (80.99)

R: Ground Sheave for Hol6t 851.94 854.35 98.90

Line of Cathead 5 (259.84) (260.58) (30.16)

S: Holst Drum for 784.79 924.36 99.36

Catheads 5 and 6 (239.36) (281.93) (30.30)

Figure 6.2 Field Survey Data for Location of Hoisting Cables

127
FIELD DATA

COMPUTED RESULTS

N IOOI.OI N 874.06

AT CATHEAD 5

I ft = 0 306 m

Figure 6.3 Sectional Elevation at Cathead 4 and Computed Results

of Cable Layout for Catheads 4 and 5

128
9) O

o

CO

129

(a)

Jacking

frames.

(b)

Scaffolding

and

cathead

profile
* NEAR SIDE JUMPFORM BEAM

(FAR SIDE IS IN OPPOSITE DIRECTION)

I ft. = 0.305m

figure 6.5 Forces on the Jumpforra Beams

130
Load Case(2) - Weight of Scaffolding Load Case (3) - Live Loads

Loc

i J K L Force^\

Fx -674 366 91 -226

F

y -2426 -2426 0 0

F

z 0 0 0 0

M

y

0 0 0 0

Load Case (4) -Weight of Cathead

Loc

I J K L Forces\

Fx -1293 784 -55 232

F

y -928 -928 0 0

* F

z 213 -123 71 -66

*M

y 67 66 101 91

\ Loc

Force s^\

i J K L

Fx

337 -212 32 0

F

y -970 -970 0 0

*F

z 0 0 0 0

*M 0 0 0 0

y

Load Case (5) - Hoisting Loads

v

'v Loc

Forces^

I J K L

Fx -19025 12296 426 2964

Fy -1996 -1996 0 0

* F

z 2151 -2151 707 -1544

*M

y 798 798 1243 2115

Near side jumpform beam (force on far side beam equal and opposite)

Units: Moment in ft — lb (1.00 ft lb = 1.36 N — m)

Force in lb (1.00 lb = 4.45 N)

Figure 6.6 Summary of Loads Applied to the Shell

131
Cathead beam

Figure 6.7 Definition of Gravity Loads

132
A, B = supports of inside and outside cathead legs, respectively.

A

N’

= cat head leg reaction components in plane of figure and normal to AA

^ and BBp, respectively.

Ap, Bp = cathead leg reaction components in plane of figure and parallel to AAp and BBp, respectively.

A

Z’ = horizontal cathead leg reaction components of A and B, respectively, normal to plane of figure.

W = resultant weight of cathead beam assembly

T, T

q = tension in static line due to hoisting loads and self weight

Fp, Fq

= tension in hoist line (F^ = F

Q

)

F = tension in counterstatic line

c

a = AAp = 18.65 ft

b = BBp = 17.64 ft

W = 1071 lb

Wp = 500 lb

W

2 = 449 lb

W

3 = 239 lb

6 A

= 77.740°

e B = 79.434°

e c

= 53.031°

a

o = 12 -*°

1 ft = 0.305 m

1 lb = 0.453 kg

Figure 6.8 Free Body Diagram of Cathead under Gravity and Hoisting Loads

8

. 65

'

17

. 35

'
. r

WELDED PIPE

CHUTE DETAIL

VOLUMES:

Bucket

Chute

HOISTING LOADS

= 695 in 3 (11.40 x 10 6 mm3

)

= 130 in 3 (2.13 x 10 6 mm 3

) 1/2 yd. Concrete = 13.5 ft 3 (0.378 m

3

)

WEIGHTS:

Bucket

Chute

Concrete

Steel Balls (two)

Mi sc. Hardware

= (.283 x 695) = 197 lb (89.2 kg)

= (.283 x 130) = 37 lb (16.8 kg)

= (150 x 13.5) = 2,025 lb (917.3 kg)

= 580 lb (263 kg)

65 lb (29 kg)

= 2,904 lb (1316 kg)

B = 2,900 lb (1314 kg)

TOTAL

Use W

Figure 6.9 Concrete Bucket Dimensions and Weights

134
CONSTANT INPUT PARAMETERS

L = BG = 218.56 ft (66.66 m)

A = 0.098 in 2 (63.21 mm

2

)

E = 13 x 106 psi (89.7 x 106 kN/m 2

)

X

G

= 115.94 ft (35.36 m)

Yg

= 185.27 ft (56.51 m)

VARIABLE INPUT PARAMETERS

L

x = BQ

L

2 = GQ

D = KQ

Wg = Bucket Weight

Figure 6.10 Simple Model for Analysis of Cable

Forces Due to Hoisting Loads

135
*1

*2

8 3

e 4

6

5 =

e i =

8 o =

tan

-1 (D/Lj) "

tan

'1 (D/L 2

)

IT2 ~ Y 1

IT2 Y

2

=. refer to fig. 6.10

(63 + B 4 )/2

6 - y

2

B

5 - 83

* F

* F

D

* t

D,

Do =

* 6 =

s„ =

*

* F

c

* A

P

* B

p

* A„

* Bo

Wg sin Bp/sin (83 + 83) = force on static line

Wg sin 82/sin ( 63 + B 2

) = tension of hoist line

F /(2 cos 65) = tension in static line

L^/sin B3 = length BK, fig. 6.10

L

2 /sin B 4 = length GK, fig. 6.10

Df + D

2 = stressed length of static line under tension T

Ts/ (AE + T) = static line elongation

s - 6 = stress-free length of static line

x coordinate of G after elastic deformation of cathead

y coordinate of G after elastic deformation of cathead

force in counterstatic line due to hoisting loads

Cathead leg reaction components in directions

h shown in figure 6.12 due to hoisting loads

* Designates results printed out by computer

Figure 6.11 Procedure for Cable Analysis Program

136
(b) View of cathead legs

in tangential plane.

CATHEAD LEGS

(c) Partial shell profile.

(d) Configuration of cathead leg

reactions relative to shell.

1 ft = 0.305m

Figure 6.12 Configuration of Cathead Legs and Reaction

Components Relative to Shell

137

18.22’

I.L.

17.35’

O.L.
PROFILE

FORMULATION OF LOAD-DISPLACEMENT EQUATIONS

Cable: A = 0.154 in 2 (99.33 mm 2

)

E = 13000 ksi (89.7 x 106 kN/m 2

)

(AE)

1 - (AE)

3

= 2 x 106 lb (8.9 x 106 N)

(AE)j - 3.2 x 10 6 lb (14.2 x 106 N)

R - 0.522 F„

«1 “ F

CLX

/ 2 (AE)

1

6 2 - F

c L

2/2 (AE)

2

" RE

3 6ec

Awf = F

c (3L4 - 4a )/12EI

«4** f ( 4wf . Lj, L

6

, 9 c>

b

6

)

A

c - «1 + «

2 + « 3 +

A

c - KF (K - const.)

1 ft = 0.305 m

* { 4 “ displacement of point C due to lateral deflection Ayy of WF beam

Figure 6.13 Geometric, Material and Load-displacement

Characteristics of Counterstatic Line

138

SO
SMALL DISPLACEMENT GEOMETRY;

B

1 B2 = bab

*JL BB

X

, BjB-j = A

1A 3 = GG1 = CC2 = BjB 2 cos(-i - 0 B

)

B

2 B3 = Bl B2 sin (-~ — 0g) * A

2A3 ~ A^A3 tan — 0^)

Oq = tsn [ (

A

2A3

"4*B 2B

3

) / A^B^ ] 9 A

3

G-^ = A^G

G^G

2 A3G^ tan oiq A

2A3

j Xq B^B 3

^ ^1^2

C

X C2 = A2A3 — A^C tan cbq* C

2

CC-^ = tan (C^C 2 /CC2

)

ecl = -1 - 9 C - c^cq, c

x c = (cc 2 + qc|) 1/2

6 C = C^C sin 0 C1 , A c = KF

C (from figure 6.13)

XG

= < A

c

/fic>x

G> YG

= (A

c /«

c )yG**

* = rotation of BB^ about point B

** Xq, Yq = displacement of G due to F

c

Figure 6.14 Displacement of Point G due to Axial

Elongation A

c of Counterstatic Line

139
Known

:

T, F d

, F q = F

d

, 6 1# a

Q = 12.4°

E = 29 x 10 6 psi (200 x 10 6 kN/m 2

), I = 65.2 in4 (27.1 x 10 6 mm4

)

Lj = 5.3 ft (1.6 m), L

2 = 3.7 ft (1.1 m)

L

3 = 1.8 ft (0.55 m), L4 = 10.5 ft (3.20 m)

Computed

:

&1 = P

1 L^/3EI, &2 = (3L1 - L2

) P

2 L

2 /6E'I

Mf = P^L-p ^

2 = P2 L2

, M

3 = P

3 L

3

e = (2M

X + 2M2 + M

3 )L/6EI, A3 = L-J^e

YG

= A

1 + A 2 + A3

X” = 0

*

X

Q

= X£ + X£ = X£ (see fig. 6.14)

Y g

= Y£ + Yg ** (see fig. 6.14)

* (XG> Y

g

) = displacement of G due to cathead beam deflection

** (Xq,Yq) - displacement of G due to elastic deformation of cathead

Figure 6.15 Displacement of Point G due to Flexural

Deformation of Cathead Beam

140
EXPANOED VIEW OF TOP 3 LIFTS. F.E. MODEL

Points

ol bail

application Elamairt no. Elevation (ft.) Radius (ft)

16& 916 133 427

165.721 133 466

146 176 133 664

164.337 133 626 143666 134031

142776 134237

142.126 (34 366

1III2I 134 Ml

MO 876 134707

166916 01170

167 264 136.848

166.283 136611

E (ksf) i tgS

* Bolt hole location

Element no. Elevation (ft.)

i669B9

161126

160 283

161444

146 606

141771

122 446

13.499

69776

30599

00

I

77777777777

Radius (ft) Thickness (ft.) E (ksf) x 105

133427 134867

135916

137173 139437

139.709

144 991

152743

162.131

170300

178 960

i *

Figure 6.16 Cooling Tower Finite Element Model

141
SHORE III LINE LOADS

Figure 6.17 Shore-Ill Line Loads

142
Insure 6.18 Normal Force (Fx

) at Cathead for Load Case 4

143
Figure 6.19 Sign Conventions

144
Meridional

stress

resultant

- N$,

kips/ft

Angle -0, deg

Figure 6.20 Meridional Stress Resultant at Cathead No. 4

145

kN/m
Hoop

stress

resultant

-

Me,

kips/ft

ELEVATION

Angle deg

Figure 6.21 Hoop Stress Resultant at Cathead No. 4

146
Angle deg

Figure 6.22 Circumferential Moment at Cathead No. 4

147
Angle -6, deg

Figure 6.23 Meridional Moment at Cathead No. 4

148
20 30 40 Moment, kip-in

Figure 7.1 An Interaction Diagram under Combined Bending

Moments and Axial Forces

Figure 7.2 Physical Dimensions of Vertical and Horizontal Cross

Sections of a Shell Element Used for Analysis

- 120

- 100

- 80 *

- 60

- 40

0

SECTION

149
Figure 7.3 Procedure Used to Estimate Compressive Strength of

Concrete at the Time of Failure

CATHEAD GANTRY

No. 4

Figure 7.4 Locations of Maximum Hoop and Meridional Forces

and Bending Moments in Lift 28

150

MPa
N*-

UP

0123456189 kN-m

Figure 7.5 Combined Effect of Axial Force and Bending Moment

at Locations A and B 012 345kftl-m 67 ®

figure 7.6 Combined Effect of Axial Force and Bending Moment

at Locations C and D

151
APPENDIX A

Patent of Wall Construction System

152
Appendix A

Patent of Wall Construction System

United States Patent ti9j Scheller

Hi] 4,040,774

[45] Aug. 9, 1977

[54] APPARATUS FOR CONSTRUCTING

CONCRETE WALLS

[75] Inventor: Herman Scheller, Eaton, Colo.

[73] Assignee: Research-Cottrel, Inc., Bound Brook,

N.J.

[21] Appl. No.: 681,723

[22] Filed: Apr. 29, 1976

[51] Int. Cl.’ E04G 11/04; E04G 1 1/28

[52] U.S. 0 425/65; 249/20; 254/139

[58] Field of Search 425/63-65;

249/20-22; 264/33-34; 254/139, 142

[56] References Cited

U.S. PATENT DOCUMENTS

661,318 11/1900 Nelson 254/139

957,521 5/1910 Talbot 249/20

3,472.477 10/1969 249/20

3,521,336 7/1970 Rohlf 425/63

3,761,551 9/1973 Ogata et al 264/33

3,779,678 12/1973 Scheller 425/65 Primary Examiner—Francis S. Husar

Assistant Examiner—John McQuade

Attorney. Agent, or Firm—Harold L. Stowell

[57] ABSTRACT

In an apparatus for constructing high-rising, poured

concrete walls, pairs of spaced-apart, upright supports

are preliminarily mounted on a foundation and then detachably attached to both sides of at least a partially hardened level of concrete wall and at intervals along

the length of the wall for repeated, upward, step-wise

use as the wall is being formed; a plurality of carriages are mounted on adjacent supports along both sides of

the wall for continuous upward movement as the wall is being cast, and adjustable concrete shaping assemblies

are mounted on the carriages. Each assembly opposing

a similar assembly to define a continuous mold into which new concrete is poured on top of previously poured concrete to form the wall.

3 Claims, 10 Drawing Figures

153
U.S. Patent Aug. 9, 1977 Sheet 1 of 6 4,040,774

FIG i

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U.S. Patent Aug. 9, 1977 Sheet 2 of 6 4,040,774

FIG. 3.

155
U.So Patent Aug. 9, 1977 sheet 3 of 6 4,040,774

FIG 5.

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U.S. Patent Aug. 9, 1977 Sheet 4 of 6 4,040,774

157
U.S. Patent Aug. 9, 1977 Sheet 5 of 6 4,040,774

o

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Sheet 6 of 6

4,040,774

U.S. Patent Aug- 1977

159
1 4, 040,774

APPARATUS FOR CONSTRUCTING CONCRETE

WALLS

CROSS-REFERENCE TO RELATED SUBJECT

MATTER

Related subject matter is disclosed and claimed in my

U.S. Pat. No. 3,779,678 granted Dec. 18, 1973. BACKGROUND OF THE INVENTION

Modern concrete wall casting techniques frequently

utilize pairs of spaced-apart, shaping forms held in position by various types of movable supports, the concrete

being poured between te forms and on top of the section of wall poured earlier and partially set. After the last poured concrete has at least partially set, the forms are removed and relocated above the former position and

then the procedure is repeated until the wall is completed. SUMMARY OF THE INVENTION

This invention provides a new and improved apparatus for efficient casting of shaped concrete walls and, in the exemplary embodiment, includes a plurality of pairs of spaced-apart, upright supports extending from the foundation initially and then attached to opposing sides of the wall being formed along the entire length of the

wall, sections of each support being detachable at the bottom of the support and from the wall at intervals and

re-attached at the upper end of the support to form a continuously advancing support as the wall is formed. BRIEF DESCRIPTION OF THE DRAWINGS

FIG. 1 is a perspective, stylized view of the base

portion of a cooling tower veil being formed by the

casting apparatus of this invention; FIG. 2 is a vertical section of the apparatus showing

the preferred form of supports and carriage; FIG. 3 is a top view of the apparatus shown in FIG.

2; FIG. 4 is an enlarged, fragmentary side elevation of a form or shaping assembly; FIG. 5 is an enlarged, vertical section showing the carnage elevating and retaining mechanism; FIG. 6 is an enlarged, vertical section showing the intermediate form in relation to the carriage; FIG. 7 is a rear elevation of the sheet portion of the form assembly; FIG. 8 is a fragmentary horizontal sectional view of

the form assembly and related apparatus; FIG. 9 is a perspective view of the hoisting apparatus; andFIG. 10 is a side elevation of the hoisting apparatus. DESCRIPTION OF THE PREFERRED

EMBODIMENT

The preferred embodiment of the apparatus is set forth herein, for illustrative purposes, in connection

with the construction of a concrete veil or outside wall of a nature draft cooling tower, which frequently rises 450 feet above the ground and has a diameter of 350 feet

at ground level. In such construction, the round concrete wall decreases in diameter as it rises until narrow

section or throat is reached, and then increases in diameter toward the top to form the hyperbolic shape. As

illustrated in FIG. 1, the tower veil or wall 10 has an

outer surface 12 and an inner surface 14, and a series of

carriages 18 located entirely around the top of the wall

2

During the initial construction of a cooling tower and

after the ground foundation has been laid, pairs of opposed, spaced-apart, generally vertical supports or Hbeams, designated 20 for the inside 14 and 20' for the outside 12, are anchored on the foundation. Each sup- port 20 and 20', not shown in FIG. 1 burl located at each carriage 18, is 20-30 feet long and is composed of

several sections 22 of extruded aluminum, generally H-shaped beams, jointed at their ends by bolted plates 24 as shown in FIGS. 2 and 5. The beams of each pair are separeated from each other by a distance while will generally be the thickness of the wall 10 being formed, which at the base of a tower is about 30 inches. The

beams are spaced apart, as desired, as best shown in FIG. 5, by a plurality of horizontal spacers or internally threaded inserts 26 attached to the beam flanges 20A by

bolts 26a before concrete is poured between the beams Each pair of beams are separated by about 10 feet from

similar pairs around the wall so that the beams are lo- cated at spaced intervals completely along both the

inner and outer wall surfaces 12, 14. As the concrete is poured between the beams and the wall increases in height, a lower section of each beam is detached from

the side of the wall 10 by removing bolts 26a from the spacers 26, leaving the spacers in the wall. Then the section is attached to the top of the same beam. As the concrete cures, the lower sections of each support held

tightly against the wall surfaces easily support the car- riages 18 and other structures mounted on them. The

beam sections 22 may be lifted into new position by a simple block and tackle or other lifting means mounted

on the carriages 18. The beams 20-20' are also provided

with spaced apart lugs 30 (FIG. 2) permanently affixed

to the outer flange 20B of each beam for a purpose later described. A plurality of carriages 18, shown in FIGS. 1, 2, 3, 4, 5 and 6, are separately mounted for vertical movement

on beams 20-20' and adjacent carriages support there between concrete form assemblies 16 which shape the surfaces of the wall being formed. Additionally, as shown on FIG. 2, the carriages provide a working plat- form for men and equipment which extends entirely around the wall being formed. Each carriage 18, reference being made to FIGS. 2, 4, 5 and 6, includes a pair of similar frames 19 composed of

several common structural beams welded and bolted together to form a strong, unitary structure. In detail and as shown in FIG. 2, each frame includes a vertical channel 32, which is welded to a horizontal tubing 46, which in turn is bolted with bolt 132, to channel 34, which is welded to channel 44. Two adjacent frames 19, one being constructed opposite hand, are jointed to- gether to form each carriage 18 mounted on a beam 20 by bolted connector plates 33 (FIG. 5), angle 38, roller frame 148, and machine bolts 108, 126, 128, 130 and 132, which also serve as axles to other components. A pipe 42 and a thread rod assembly 40 are bolted between

adjacent frames 19, and serve as an adjustable support

to channel 34. Wheels 110 are attached to channel 32 and tube 46 of each frame by axles 102, 104, 106 and 108, and ride against opposing sides of the two outer flanges 20B of the beam holding the carriage against the beam. A hydraulic jack 118 (FIG. S) is attached to plate 33 and roller frame 148 in each carriage by bolts and mounts and exerts upward thrust against the carriage when hydraulic pressure is supplied through flexible hose 120. A reciprocating bar 140 is attached to the ram

of the hydraulic jack 118, which in turn has a pivoted

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pawl 142 which is spring loaded (spring not shown) to insure positive engagement between pawl 142 and lug 30 on beam 20-20'. Further, wheels 144 are mounted on roller frame 140 which extends between adjacent beams 32, and are in rolling contact with reciprocating bar 140 to resist lat- eral forces on bar 140 and prevent disengagement be- tween pawl 142 and lugs 30. In addition, a follower pawl 124 welded to a short pipe 125, which in turn is mounted for rotation on a shaft 126 extending between beams 32, engages lugs 30 and is held in engagement by

spring 146. Normally, the carriages progress in an upward direction by extension of pawl 142 until tooth 124 engages a higher lug, 30 followed by withdrawal of pawl 142 for relocation against a higher lug 30, and repeating. However, if necessary, a carriage can be lowered by extending hydraulic jack 118 so that the load carried by fol- lower pawl 124 is now transferred to extendable pawl

142 and engaged cleat 30. When the load is thus trans- ferred, pawl 124 can be pivoted so that by slowly re- tracting the jack will allow the carriage to settle and

follower pawl 124 to engage in the next lower cleat 30. A working scaffold is formed by planks 36 resting on beams 34 of adjacent carriages and an outer rail 44 between which ropes or planks may be placed. Concrete shaping form assemblies 16, generally shown in FIGS. 4, 6 and 7 span the distance between

adjacent carriages 18 on both sides of the wall being formed and form and sides of the mold into which fresh concrete is poured. Each form assembly 16 includes a forming sheet 64, two end stiffback channels 56, to which lateral, telescoping steel trusses 70 are bolted, with incorporated wedges 74, that strengthen the

trusses 70 when in the proper position. Forming sheet 64 may be 3/4j plywood for example, or any ridgid material that can be easily cut, or added to, to provide

a smooth casting surface, telescoping trusses 70 provide

lateral support for sheet 64, throughout their effective length, once the incorporated wedges 74 are wedged

firmly. The stiftback channels 56 are structural channels that are vertically supported at the bottom by axle 104 which extend beyond beam 32, thru slotted holes 63 (FIGS. 6 and 8) to a retainer plate 62. Further, axle 108

also extends beyond beam 32 to the retainer plate 62. The form panel 64 is held firmly against the extruded

flange 20A of beams 20—20'. This is accomplished by wedging the opposite side of the stiffback channel 56 (FIG. 6) with a steel wedge 58 against the extended axle

108.Similarly the top of channel 56 forces form panel 64

tightly against the extruded flange 20A of beams 20-20', with a hardwood wedge 60, driven between its opposing side and angle 38. With reference the FIGS. 6 and 7, the forming sheet 64 is held vertically by a bracket 66 which merely serves to keep the sheet 64 from falling out when the entire form assembly is unwedged as re- quired while "jacking". As in the case of cooling tower construction, as the tower moves upward in height the diameter decreases

to the throat or neck of the tower, requiring the car- riages 18 to become closer together. With reference made to FIG. 8 it can be seen that as the carriages 18 move upward and wedges 74 of the steel trusses 70

being loosened, they are forced together, or can open

freely as required. FIG 8 also illustrates that while the form assembly 16 has its wedges loosened, the form

4

panel 64 is only supported from falling by bracket 66 and can easily be pulled out be sliding it upward from between the support beams 20-20' and the steel trusses 70 for cutting. The nature of the extruded aluminum beams 22 that make up the support beams 20-20' permits that they can

be repeatedly bent or flexed approximately 3 inches over each length without permanent damage

Using this characteristic, (FIG. 2) a frame 76, comprising two opposite hand weldments are attached to

either half 19 of the carriage 18 on one side of the wall

10. The frame 76 encases the outer flange 20B of a portion of beams 20-20' which is adjacent to firm concrete. Similarly a frartie 82 is mounted on the top of

either half 19 of the carriage 18 which encases the outer

flange 20B of beams 20-20' above the form enclosure. An arm 78 is bolted to the lower frame 76, and in turn

a stream-bolt ratchet 80 connects to the top frame 82. By extending or contracting the steam-bolt ratchet 80,

lateral pressure is applied sufficient to bend the support beam 20 and properly align the wall 10 being cast. The vertical strength afforded by such a forming

system is also used to best advantage by eliminating excessively high mobile cranes or tower cranes required

in high rise construction, as used in previous forming

systems. Gantries are erected at selected points around

the top of the tower as illustrated in FIGS. 9 and 10. Each gantry is comprised of a beam 88, which is bolted

to four pipe legs 86, which in turn are attached to four separate carriages 18. On the upper side of beam 88, two

sheaves 90 are attached. A commercial hoisting engine 93 on the ground is used with these gantries so that the load line 92, originating from the hoisting engine on one

side of wall 10, is supported over the wall on sheaves 90

so that materials can be hoisted on the opposite side of

wall 10 as the arrows indicated in FIG. 10. Also, a stationary static line 94, which extends from beam 88 to a strategic point or concrete hopper on the ground, is used as a guide for the load line 92, which is attached to a small block 96 so that its sheave wheel rolls on static line 94. The static line 94 serves to make the ground

loading point (not shown) central, which expedites concrete handling from a concrete truck or hopper, as well as keeping the workmen from working beneath the scaffolding along the perimeter of the tower. The two variations in the form surface, or flanges 20A

of jack beams 20 and 20', are shown as in cooling tower

construction, the outer wall surface 12 is generally ridged and the inner wall surface 14 is generally smooth. The plurality of vertical ribs thus formed on

the outside surface 12 of the completed tower serve to induce air turbulance, thus enhancing heat transfer. In casting the initial courses of the wall, an ordinary

general crane is used to raise the plastic concrete from

the ground to the working scaffold where it is distrib- uted by wheeled carts to the forms around the periph- ery of the structure. It will be appreciated that several cranes may be used simultaneously so that such a large structure can be cast at a reasonable rate. When the structure height exceeds reach of the cranes, a number

of gantries are attached to selected carnages spaced around the wall. These gantries now provide the means

to transport plastic concrete and other materials to the work area. The concrete is raised in buckets (approxi- mately J cu. yd capacity) by cables running over the

gantries and returning to hoisting engines on the ground. Thus, the necessity for using a large tower

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crane is avoided and the work can progress much more

quickly and safely than with such a crane.

1 claim:

1.

Apparatus for forming a concrete wall comprising

a. a plurality of generally vertical supports spaced

apart along both sides of the wall being formed, each support being located opposite to another

support on the other side of the wall and detachably connected to said opposing support by a plurality of

spacers some of which are contained in the wall previously formed, each support consisting of separate segments detachably joined together at their ends, and the lower portion of each support being

held contiguous the surface of the wall previously formed by said spacers while the upper portion extends above the portion of the wall previously formed.

b. a carriage mounted on each support with means for advancing the carriage upwardly along the supports as the wall is formed,

c. a plurality of concrete casting assemblies located along and on both sides of the wall being formed

adjacent its top, each assembly supported at its ends by two carriages on adjacent supports and each

assembly having a generally vertical casting surface extending between the adjacent supports so that the assemblies and the supports together define a continuous mold extending on top of the wall into which concrete is poured, whereby as the concrete hardens the carriages and assemblies are moved

upwardly on the supports to form a new level of

6

wall and, at intervals, support segments are de- tached from the bottom of each support and at- tached to the top to provide a continuous track for the carriages,

d. each casting surface comprises the inner surfaces of

generally horizontal, plywood sheets which combine to form a continuous casting surface,

e. each casting assembly comprising a plurality of

generally horizontal braces positioned contiguous

the outer surface of the plywood sheets to support

the sheets while the concrete is setting against the inner surface of said sheets, and

f. each casting assembly comprising at least a pair of

slotted beams movably earned by adjacent carriages, said plurality of horizontal braces mounted

between said slotted beams which are adapted to be wedged against said adjacent carriages, so that by wedging the slotted beams against the carriages the inner surface of the sheets are drawn into contact

against the adjacent supports thereby providing

said continuous mold.

2. The apparatus defined in claim 1 further including

hoisting means attached to said carriages, a ground

supported hoisting engine, and a hoisting cable connecting said hoisting engine and said hoisting means.

3. The apparatus defined in claim 2 further including a stationary cable connecting the hoisting means to a

central loading point below on which the hoisting cable

is guided.

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APPENDIX B

Computer Programs for Shell Analysis

163
Appendix B

Computer Programs for Shell Analysis

B.l Introduction

Failure of the cooling tower shell under the imposed construction

loads at the location of cathead gantry no. 4 was believed to be a

plausible cause for the initiation of the total collapse of lift 28.

Thus, an analysis of the shell structure as it existed at the time of

the collapse was made using the construction loads. The complexity of

the loading conditions as well as the variation in shell thickness and

material properties required that the finite element method be employed

as the means of analysis. A survey of finite element programs available

for the analysis of this type of shell structure with the constraints

mentioned previously indicated that SHORE-III[6.2] satisfied these

requirements and has been used previously [B.l, B.2] to analyze this

type of structure.

As a means of verifying the results obtained in the SHORE-III

analysis, a second finite element program, SAP IV [6.3], was selected.

SAP IV is a general finite element program which provides an alternate

method for load input and an alternate shell model from those used in

the SHORE-III analysis. "

This appendix will present a discussion of the pertinent features

of the two finite element programs used in the analysis of the shell

and a comparison of the results obtained from each analysis, for

selected loading conditions.

164
B.2 Discussion of SHORE-III and SAP IV

SHORE-III (SHORE) is a finite element program for the linear

elastic static and dynamic analysis of arbitrarily loaded axisymmetric

plates and shells. SAP IV (SAP) on the other hand is a general finite

element structural analysis program for the linear elastic analysis of

three dimensional structural systems. Although both programs are capable

of performing either a static or dynamic analysis, this discussion will

be limited to the static analysis since all construction loads used in

the analysis were treated as being static. Furthermore, since SAP has

a rather large element library (truss, beam, plane stress or strain,

three-dimensional solid, pipe, etc.) only the plate/shell element which

was used in the SAP analysis will be discussed.

The capacity of SAP is primarily dependent upon the total number

of nodal points needed to model the shell while SHORE is restricted to

a model comprised of no more than fifty (50) elements. This element

restriction for SHORE is not serious for this analysis since the stress

distributions at the lower elevations of the shell do not have a significant effect on the results obtained for lift 28. Thus, the number

of elements used to model the lower portion of the shell can be reduced.

The shell model is developed for SHORE by discretizing the meridian

curve of the shell with a series of curved rotational ring elements.

SAP requires a discretization along the meridian and around the circumference of the shell in order to develop its model. Thus, the SHORE

model is composed of a series of continuous ring elements along the meridian of the shell while the SAP model is a three-dimensional assemblage

of flat plates.

165
Only a brief description of the elements used in the SHORE and SAP

analysis will be presented. References which present the details on the

formulation of the individual elements are given in the description. The

curved rotational ring element [B.3] used in SHORE has an element stiffness matrix which is derived from displacement fields that may vary from

linear to sixth order and includes the exact geometry of the shell as

well as the effect of transverse shear deformation. The extra coefficients in the higher order displacement fields are eliminated by

kinematic condensation at the element level. Sixth order displacement

fields were used throughout in this analysis. The element used in the

SAP analysis is a quadrilateral of arbitrary geometry formed from four

compatible triangles. A constant strain triangle [B.4] and a linear

curvature compatible triangle with nine (9) degrees of freedom [ B . 5

]

are used to represent the membrane and bending behavior, respectively,

of the SAP element. A central node is located at the average of the

coordinates of the four corner nodes and has six degrees of freedom

which are also eliminated by condensation at the element level.

Each element in both SHORE and SAP have constant material properties (moduli of elasticity) although the properties may vary as

required from element to element in the respective models. The thickness of a particular element in SAP must be constant, but may change as

required throughout the SAP shell model. The thickness of the SHORE

element can vary linearly along the meridian as dictated by the shell

geomet ry

.

Various loads including thermal effects may be used as input for

both the SHORE and SAP analysis with distributed loads (gravity,

166
pressure, etc.) and thermal loads being treated as consistent equivalent

nodal loads in both analyses. However, a major difference exists in

the manner in which other loads are input for analysis. SAP requires

that all external loads including moments applied to the shell structure

be input as concentrated loads at nodal points while for the SHORE

analysis all loads are expanded in Fourier harmonics with respect to

an element nodal point and the final result is obtained by superimposing

the results of each harmonic.

Finally, both finite element programs solve the resulting set of

linear simultaneous equations for the structural model by a modification

of the Gaussian elimination scheme which takes advantage of the symmetric

narrow banded nature of the global matrices used.

B .3 Comparison of Results

Several of the most significant differences between the two finite

element models are considered and comparisons of results are made. The

differences include the effect of element discretization, boundary conditions, the method of applying the loads, the precision of calculations

and the type of elements used.

One of the biggest differences between the two analyses lies in the

initial scheme for discretizing the shell. While SHORE uses continuous

ring elements as shown in figure B.l, SAP requires that the rings be

broken into numerous elements. The radial grid adopted is shown in

figure B.2. The changes in radial increment were chosen to reduce the

number of nodes required for the model in hopes of making the program

manageable on the computer while preserving resolution at cathead gantry

no. 4 and providing loading points at the other catheads. As it was,

167
the SAP analysis required more than eight hours of computer time. The

effect of having varying radial increments was to introduce some variations in the solution which could be attributed directly to the variations

in the size of the elements. This can be best illustrated by considering

the gravity loads which should be identical at any angle. Figures B.3

and B.4 illustrate the radial variation of hoop stress, N0 , and the

meridional stress, , respectively, 1.5 ft (.457 m) below the top of

lift 28. The radial variations in the stresses as predicted by SAP are

due to the changes in size of adjacent elements and to the poor aspect

ratio of the larger elements. The smallest elements at cathead no. 4

(elements 126 and 127) are 2.18 ft (0.66 m) long by approximately 0.6 ft

(0.18 m) high for an aspect ratio of 3.63 while the largest elements near

the top (elements 112 and 120) have an aspect ratio of approximately 58.

As can be seen in figure B.3 and B.4 element size change has more of an

effect on (fig. B.3) than the aspect ratio, while the aspect ratio has

more of an influence on N

0 (fig. B.4). However, figure B.3 and B.4 do

illustrate that the stress distribution becomes more uniform near cathead

no. 4 where the mesh is finer and the aspect ratio of the elements is more

favorable to obtaining a good solution. In the SHORE analysis, the

radial stress distribution is dependent only upon the equations used to

develop the ring elements. The SHORE and SAP analyses for the gravity

load showed good agreement between and N

0 along the meridian at

cathead no. 4. The largest difference was less than 5 percent.

Different boundary conditions are used in the two models. ^he

SHORE model uses an open type element developed specifically to represent the columns at the base of the shell. The effect on the model is

168
to smear the stiffness of the columns into a ring. In the SAP model

the shell at the top of the ring beam tying the columns together was

fixed by prescribing zero displacements. This was done to reduce the

number of nodes in the SAP model. Consequently, the stress distributions

within the bottom ring of the shell are quite different for the two

models. However, in element 17 of the SHORE model (see fig. B.l) and

the corresponding elements in the SAP model, the stress distributions

are similar. The same results are presented in reference B.l where the

effect of different boundary conditions on a hyperboloidal shell were

studied. Consequently, the effects of the boundary conditions at the

base seem to be far enough away from the area of interest, the top two

lifts, so that none of the differences in results are due to the differences between boundary conditions used in the two models.

The method of applying the loads seems to have a great impact upon

the results. SHORE requires that the loads be applied as line loads

acting over some finite length. The length is chosen to allow a reasonably rapid convergence of the fourier series used to generate the load.

SAP requires that the loads be point loads applied at nodes. For this

problem where the loads are essentially applied to the shell by bolts,

the point load approach is more realistic. The effect of using the

distributed line loads is to cause the stress distribution to be more

uniform near the point of application and the maximum stress predicted

should be lower than the real stress experienced by the shell.

The loading function used in the SHORE analysis is developed by

first distributing the concentrated loads about the centerline of the

jumpform beam. The loads are distributed over 0.358 degrees (10 in

169
or 254 mm at lift 28) for the scaffold loads (load cases 2 and 3) and

0.859 degrees (24 in or 610 mm at lift 28) for the cathead gantry and

hoist loads (load cases 4 and 5) . The 10 in (254 mm) distribution width

is the surface contact length between the shell and jumpform beam. This

distributed load is then expanded in a Fourier Series which applies the

load at the required points around the circumference of the tower for

that particular construction load. The larger distribution angle used

in load cases 4 and 5 was chosen because it reduces the number of Fourier

series harmonics required to adequately define the loads. Ideally, a 0.358

degree distribution angle should have been used for all cases. However,

when the distribution angle was reduced from 0.859 to 0.358 degrees, the

number of harmonics required to produce a load function with an acceptable

shape increased from 56 to 150 and computer time and costs almost tripled.

A sample analysis was conducted using both the 0.358 and 0.859 degree distribution angles and the maximum stresses differed by only a few percent.

The stress distributions were also essentially the same. Consequently,

it would appear that the compromise between distribution angle and computer

analysis time is justified. Figure B.5 illustrates this distributed line

loading for the normal, meridion and tangential forces F , F , and F , x y z respectively and for the meridion moment M^, applied to the shell by the

jumpform beams. The bold vectors represent the points loads used in the

SAP analysis.

The development of a convergent Fourier series with only a few harmonics was found to be a difficult task for load cases 4 and 5 which areapplied only at cathead 4. This is because as the number of application

170
points decreases (two points or ribs for cases 4 and 5), the number of

harmonics required for a convergent series increases rapidly. In addition,

the required computer time to analyze a load case using SHORE is related

to the number of harmonics. An investigation of the stress distributions

that occurred around the circumference of the shell in load case 4 where

the six cathead gantry loads are applied, for both the SHORE and SAP

analysis, indicated that the internal shell forces decayed rapidly to a

small value at approximately 20° from the cathead as shown in figures B.6

and B.7. Thus, it was concluded that since the catheads are 60° apart, the

loads applied at a cathead have little influence on the internal forces at

the catheads on either side. Consequently, in order to reduce the number

of harmonics necessary to obtain convergence, the loads for load cases 4

and 5 were applied, in the SHORE analysis, at all six cathead locations

instead of just at cathead no. 4. Figure B.8 illustrates the loading

function for a normal force applied to the shell by the jumpform beams

at a cathead. This load would then be repeated at all six cathead locations in load case 4 to produce a symmetric loading condition.

The loading functions were developed for each construction load

(cases 2 through 5) and a separate analysis was made for each case,

including the gravity load which was internally generated by both programs. Since both SHORE and SAP are linear elastic finite element

programs, the principle of superposition applies and the results for

each load case may be combined algebraically to find the resultant

stresses for any combination of the construction loads.

171
Comparison between values of membrane stress, Ng , and bending moment

Mg , obtained for solutions to load case 4 using SHORE and SAP are illustrated in figures B.6 and B.7. The comparisons are illustrated for selected

elevations in lift 28 and are typical of results from other load cases.

It is interesting to note that although the loads are distributed over an

are length in the SHORE analysis, the stress magnitudes predicted compare

well with those obtained using SAP. However, because loads are applied at

points in the SAP analysis, the stresses would be expected to be larger

than those obtained using the SHORE analysis. Several features of the SAP

model and solution process may contribute to the apparent inconsistency.

A major feature is that SAP calculates stresses at the center of the element

which essentially represent the average stress in the whole element. Since

even the smallest elements are over two feet wide and the stress distribution is sharp the peak stress may be missed by a significant amount. A

second feature involves the precision of the calculations. SHORE carries

out all calculations in double precision on a 32-bit word machine while

the version of SAP used carries out single precision calculations on a

36-bit word machine. Consequently, roundoff errors may have occurred in

the SAP solution process, especially since there were over 6400 equations

to be solved and the band width of the stiffness matrix was 612. Another

reason the SAP program calculates smaller stresses may be the poor aspect

ratio of the elements and the relative size of adjacent elements. Finally, there are differences in the types of elements used in the two

models. SAP uses a plate/shell element which provides only for membrane

stresses and bending moments. SHORE on the other hand uses a shell

element which accounts for transverse shear and thus provides for a better

estimate of bending moments since the elements are relatively thick (0.667

172
ft, 0.2 m) in lift 28. This appears to be corroborated by the fact that

the comparison between membrane stresses in figure B.6 is better than the

comparison between bending moments in figure B.7.

B.4 Conclusion

Despite the differences in the values of the stresses there are

several encouraging points which arise from the comparison of the two

solutions. Both models gave stress distributions of similar shape for

corresponding stress components. Also, the fact that such different

models could lead to the prediction of stresses and moments for which

the peak values agreed within a few percent for membrane stresses is

encouraging, especially since insufficient time was available to refine

the SAP model. Consequently, the SAP model is considered to be a first

cut at verification of the SHORE results, while the SHORE program is

designed specifically for the solution of problems involving shells of

revolution.

Based primarily on the difference between the SHORE closed ring element, which included transverse shear deformation, and the SAP plate/shell

approximation which does not, it is believed that the stress resultant and

moments obtained in the SHORE analysis are a better approximation of the

stress levels experienced by the shell in lift 28 for the specified construction loads. The actual stress levels may be higher since the distributed loads used by SHORE to represent the loads applied to the shell at

the jumpform beam bolts tend to smear the loads over a larger surface area

of the shell than actually occurs and the shell model is not sufficiently

able to model the stress distributions that occur at the bolt locations.

Both the SAP and SHORE models give stress distribution that agree

reasonably well for the type of loads applied in each analysis. Thus, the

areas of high stresses in lift 28 could be predicted from the results of

either model.
B . 5 REFERENCES

[B.l] Sen, S.K. and Gould, P.L., "Hyperboloidal Shells on Discrete

Supports," Journal of Structural Division, American Society of

Civil Engineers, New York, NY, March 1973.

[B.2] Sen, S.K. and Gould, P.L., "Free Vibration of Shells of Revolution Using F.E.M.," Journal of Engineering Mechanics Division,

American Society of Civil Engineers, New York, NY, April 1974.

[B.3] Brombolich, L.J. and Gould, P.L., "A High-precision Curved Shell

Finite Element," AIAA/ASME 12th Structures, Structural Dynamics

and Materials Conference, American Institute of Aeronautics and

and Astronautics, American Society of Mechanical Engineers,

Anaheim, CA, April 1971.

[B.4] Felippa, C.A., "Refined Finite Element Analysis of Linear and

Non-linear Two-dimensional Structures," SESM Report 66-2,

Department of Civil Engr. , Univ. of Calif., Berkeley, 1966.

[B.5] Clough, R. W. and Felippa, C. A., "A Refined Quadrilateral Element

for Analysis of Plate Bending," Proceedings 2nd Conference on Matrix

Methods in Structural Mechanics, Wright-Patterson Air Force Base,

Ohio, 1968.

174
EXPANDED

VIEW

OF

TOP 3

LIFTS,

F.E.

MODEL

175

Figure

B.l

SHORE-III

Finite

Element

Model

for

Lifts

26,27

and

28
Element

no.

Elevation

(ft.)

Radius

(ft)

Thickness

(ft)

E

(ksf) x

IfflS

176
POINT

OF

LOAD

THICKNESS

(ft)

LIFT

NO.

ELEVATION

(ft)

RADIUS

(ft)

E

(Kef)

x

10 5

APPLICATION

r

*d

CT> CM

to to a a .296 .296ino>

Csl o

CM

in o o o

CM CM CM

in in in H r-4 <H CM CM CM CM

5 5

$ $ $ $

in in in in

co co cn co

I I II I I I

to 5 <d <3 3 3

CD CD CO 00 CO GO f' iID ID in M CO CM r-i O at o id U) OH O O (D in O rt CM CM CM CM CM CM CM CM CM CM CM CM CM CM CM CM CM CM CM r-i P r-i H V H r-t >

I'' C' f' r* r- C' r- t" 8

ID ID <D <D <D ID ID <D ID ID ID ID ID <D ID ID ID <D ID §

Figure

B.2

SAP-IV

Finite

Element

Model

for

Tower

Unit

No.
GRAVITY LOADS SAP Bf

Figure B. 3 Comparison of from Gravity Load, Load Case 1

178
GRAVITY LOADS SAP IV

NUMBERS

Figure B.4 Comparison of Nq from Gravity Load

179
Figure B.5 Load Functions for SHORE-III and SAP-IV Programs

180
Hoop

stress

resultant

- N©,

kips/ft

Figure B.(S Comparison of Hoop Stress Resultant for Load Case 4

181

kN/m
Circumferential

moment

- M©,

kip.ft/ft

- -1.75

- - 1.50

- - 1.25

15 20

Angle -6, deg

- - 1.00

- -.75

-.50

-.25

0.0

.25

Figure li.7 Comparison of Circumferential Momemt for Load Case 4

182

kNm/m
Figure B.8 Loading Function for Normal Force, Load Case 4

183
APPENDIX C

Results of Chemical Analysis of Cemet

184
LAW ENGINEERING TESTING COMPANY

MATERIALS TESIING ENGINEERS

SOIL AND EOUNDAIION INVESTIGATIONS

412 Av».. N. E. • P. O. Bo* 0815. SU K ATLANTA. GEORGIA 30324

REPORT CEMENT ANALYSIS

QWat: LETC0 MC LEAN Office: Atlanta, Georgia

NBS SAMPLE #45 Date: August 9, 1978

LETC0 JOB NO. M- 110 Lab. No. 21 531

B

BRAND OF CEMENT ? PLANT ? SAMPLE RECEIVED 7-15-78

PHYSICAL DATA

SETTING TIME (Gillmore)

Initial Set:

Final Set:

HOURS MINUTES

3 10 ~A 50

SOUNDNESS (Expansion)

FINENESS (Surface Area, Blaine)

COMPRESSIVE STRENGTH, PS I 1 Day Break

3 Days Break

7 Days Break

28 Days Break

AIR CONTENT (Percent by Volume)

0.00 %

3720 Sq. Cm./Gm.

1680

1900

7.2

CHEMICAL ANALYSIS

Silicon Dioxide (Si

O

2

) Aluminim Oxide (A1 ? 0^)

Ferric Oxide (Fe-Otr

Calcium Oxide (CaO)

Magnesium Oxide (MgO)

Sulphur Trioxide (S03 ) Alkalies (Na~0 & 0.658

Loss on Ignition

Insoluble Residue

Tricalcium Silicate (3CaO.Si Op)

Di calcium Silicate (2CaO.SiOo)

Tricalcium Aluminate (3Ca0.AT 2 0

3

] Tetracalcium Aluminoferrite

(4Ca0.Al 2 0

3 .Fe2 Q

3

)

k

2

°)

20.0

5.2

3.S

63.5

2.1 2.0

0.44

o~ —OS'

60

T2 ~i:r

12

NOTE: The Type of this cement is not known,

cc: H.S. Lew Washington, D.C. Retpectfully tu bmilted,

LAW ENGINEERING TESTING CO.

185
LAW ENGINEERING TESTING COMPANY

MATERIALS I ES I INC. ENCINJll KS SOIL AND EOUNDA I ION IN VEST 1CA I IONS

412 Pl.st.r t Ay... N E • P O. Box 13815 Sta K ATLANTA GEORGIA 30324

jIE!*©II7 CEMENT ANALYSIS

€2tes*t: LETCO MC LEAN

?rs>fcrt:NBS SAMPLE #29

LETCO JOB NO. M- HO

BRAND OF CEMENT_?

PHYSICAL DATA

Offic«: Atlanta, Georgia

/ Date: August 9, 1978

Lab. No. 21 531 -A

PLANT ? ; SAMPLE RECE I VED 7-1 5-78

SETTING TIME (Gillmore)

Initial Set: HOURS

2 MINUTES49 Final Set: 4 45

SOUNDNESS (Expansion) 0.00 %

FINENESS (Surface Area, Blaine) 3710 Sq

COMPRESSIVE STRENGTH, PSI

1 Day Break

3 Days Break 2140

7 Days Break 2200

28 Days Break —

AIR CONTENT (Percent by Volume) 7.4

CHEMICAL ANALYSIS

Silicon Dioxide (Si

O

2

) 20.6

Aluminim Oxide (A1 ? 0 3 ) 5.3

Ferric Oxide (Fe

? 03

) J 3.0

Calcium Oxide (CaO) 64.1 Magnesium Oxide (MgO) 2.4

Sulphur Trioxide (S03

) 2.1 Alkalies (Na~0 & 0.658 K„0) Q.38

Loss on Ignition 1 .2 Insoluble Residue Q.22

Tricalcium Silicate (3Ca0.Si0p) 60

Dicalcium Silicate (2CaO.Sio

2

7 14 Tricalcium Aluminate (3Ca0.Al20

3

) 7.1 Tetracal cium Alumi noferrite

( 4 Ca 0 .Al 2 03 .Fe2 0

3 ) _9

NOTE: The Type of this cement is not known.

cc: H. S. Lew

\

Retpedfully tubmitled,

LAW ENGINEERIN

DAN WELCH

186
APPENDIX D

Results of Analysis of Concrete

187
Technical Service Report

Project: Slow Setting Concrete Wall Section

Project No.: CT-0477 Date: November 7, 1978 Customer: National Bureau of Standards (Washington, D.C.)

Objective

Determine if Pozzolith 200-N, a water-reducing admixture, was present in concrete placed in a wall section of a thin wall structure at a concentration sufficient to account for an unusually slow setting behavior.

Conclusion

The concrete in question contained a water-reducing admixture,

similar in composition to Pozzolith 200-N, at a level not ex- ceeding that normally recommended by the admixture manufacturer

(3 to 5 fl. oz./lOO lbs. of cement).

Sample Identification and Background Information

A piece of hardened concrete, weighing about 11 lbs., and three

small bottles of various liquid substances, identified as Pozzolith 200-N, Starch Hydrozylate, and Amine Derivatives Mixture,

respectively, were received from Dr. H. S. Lew, Structures and

Materials Division, Center for Building Technology, on 9/5/78.

Methods of Test

The concrete sample was subjected to chemical analyses to determine the presence and addition level of a water-reducing admixture (ASTM C-494 , Type A). The liquid sample identified as Pozzolith 200-N was characterized to obtain its chemical

composition and certain physical properties. The two remaining

liquid samples, identified as Starch Hydrozylate and Amine

Derivatives Mixture, were subjected to infrared analysis to determine the principal ingredients present.

Re s

u

Its and Discussion

The liquid sample identified as Pozzolith 200-N consisted principally of a mixture of corn syrup (Starch Hydrozylate) and

triethanolamine (Amine Derivatives Mixture). The latter com- prised 11% of the admixture formulation as received. Chloride

ion (Cl

-

) in the amount of 0.24% also was detected. A comparison

of this sample with one analyzed previously by us in October

1975 suggested they were quite similar.

Chemical analysis revealed the concrete sample most likely con- tained no more than a normal dose of a water-reducing admixture

similar in composition to Pozzolith 200-N (ASTM C-494, Type A).

In this case, a normal dose is an addition rate of 3 to 5 fluid ounces per 100 lbs. of cement, as recommended by the admixture

manufacturer (Master Builders). A more precise determination of admixture concentration is not possible at this time, unless a calibration curve was to be

prepared employing actual job materials at the levels specified

in the mix design.

L. M. Meyer, ManagerTechnical Services Section Chemical Analyses by:

ig

CT-0477

J. R. Polky

Research Chi mist

Copy to -

J. J. Shideler

D. L. Glochowsky

Assistant Research Chemist

A. A. Alonzo

Associate Research Chemist

188
NBS-1 14A (REV. 7-73)

U.S. DEPT. OF COMM.

BIBLIOGRAPHIC DATA

SHEET

1. PUBLICATION OR REPORT NO. 2. Gov't Accession.

No.

3. Recipient's Accession No.

4. TITLE AND SUBTITLE

Investigation of Construction Failure of

Reinforced Concrete Cooling Tower

at Willow Island, West Virginia

5. Publication Date

6. Performing Organization Code

7. AUTHOR(S)

H.S. Lew, S.G. Fattal, J.R. Shaver, T.A. Reinhold, B.J. Hunt

8. Performing Organ. Report No.

9. PERFORMING ORGANIZATION NAME AND ADDRESS

NATIONAL BUREAU OF STANDARDS

DEPARTMENT OF COMMERCE

WASHINGTON, D.C. 20234

10. Project/Task/Work Unit No.

11. Contract/Grant No.

12. Sponsoring Organization Name and Complete Address (Street, City, State, ZIP)

Occupational Safety and Health Administration

U.S. Department of Labor

Washington, D.C. 20001

13. Type of Report & Period Covered

Fianl

14. Sponsoring Agency Code

15. SUPPLEMENTARY NOTES

16. ABSTRACT (A 200-word or less factual summary of most significant information. If document includes a significant

bibliography or literature survey, mention it here.)

The coolapse of the natural-draft hyperbolic concrete cooling tower unit no. 2 at

the Pleasants Power Station at Willow Island, West Virginia has been investigated.

This investigation included onsite inspections, laboratory tests of construction

assembly components and concrete specimens, and analytical studies.

Based on the results of these field, laboratory and analytical investigations, it was concluded that the most probable cause of the collapse was due to the imposition

of construction loads on the shell before the concrete of lift 28 had gained adequate

strength to support these loads. The analysis of the shell indicates that the collapse initiated at the part of the shell in lift 28 where cathead no. 4 was located.

It further showed that calculated stress resultants at several points in that part

equaled or exceeded the strength of the shell in compression, bending and shear.

The failure of these points in that part of the shell would have propagated to cause

the collpaoe of the entire lift 28.

s»7. KEY WORDS (six to twelve entries; alphabetical order; capitalize only the first letter of the first key word unless a proper

name; separated by semicolons)

Collapse; concrete; concrete strength; construction; cooling tower; failure;

hyperbolic shell; shell.

8. AVAILABILITY Unlimited

s: For Official Distribution. Do Not Release to NTIS

I

I Order From Sup. of Doc., U.S. Government Printing Office Washington, D.C. 20402, SD Cat, No, 03

[ ! Order From N ational Technical Information Service (NTIS)

Springfield, Virginiji 22151

19. SECURITY CLASS

(THIS REPORT)

UNCLASSIFIED

20. SECURITY CLASS

(THIS PAGE)

/ UNCLASSIFIED

21. NO. OF PAGES

22. Price

U SCOMM- DC 29042-P74
'