Journal of Loss Prevention in the Process Industries 23 (2010) 98–105
Contents lists available at ScienceDirect
Journal of Loss Prevention in the Process Industries
journal homepage: www.elsevier.com/locate/jlp
Failure case studies of SA213-T22 steel tubes of boiler through
computer simulations
J. Purbolaksono a, *, J. Ahmad b, A. Khinani a, A.A. Ali a, A.Z. Rashid a
a
b
Department of Mechanical Engineering, Universiti Tenaga Nasional, Km 7 Jalan Kajang-Puchong, Kajang 43009, Selangor, Malaysia
Kapar Energy Ventures Sdn Bhd, Jalan Tok Muda, Kapar 42200, Malaysia
a r t i c l e i n f o
a b s t r a c t
Article history:
Received 8 April 2009
Received in revised form
8 June 2009
Accepted 9 June 2009
Increased temperature and decreased hardness values of the tube metal and development of oxide scale
on the inner surface of boiler tubes over prolonged period of time are typical problems in power plants.
Appropriate life assessments or condition monitoring of boiler tube should be carried out from time
to time. Computer simulations may economically support the post-failure assessment method, i.e. visual
inspections, metallurgical examinations and mechanical strength measurements. However, estimations
obtained from the simulations may provide an advanced warning to take preventive actions prior to
failure. In this work two failure cases of the reheater and superheater tubes made of a typical material of
SA213-T22 steel are evaluated. As the oxide scales are increasingly developed on the inner surface, the
increasing of temperature and decreasing of hardness value in tube metal for both cases are determined.
The remnant life estimations are then made in the form of creep cumulative damages.
Ó 2009 Elsevier Ltd. All rights reserved.
Keywords:
SA213-T22 steel
Creep rupture
Remnant life assessment
Elevated temperature
Boiler tube
Finite element modeling
1. Introduction
The strength of low-alloy steel may change for prolonged period
of service. Estimations of change in temperature, hardness and oxide
scale thickness during service may be used to estimate the remnant
life of the component. In particular, estimations of the average
temperature in tube metal are important in heat recovery steam
generator (HRSG). It may provide an advanced warning of failure by
estimating temperature increase in water-tube boiler.
Port and Herro (1991) reported that almost 90% of failures
caused by long-term overheating occur in superheaters, reheaters
and wall tubes. Tubes that are especially subjected to overheating
often contain significant deposits. The deposits reduce coolant flow
and the tubes will experience excessive fire-side heat input. Scales
and other materials on external surfaces will slightly reduce metals
temperatures. The thermal resistance of the tube wall may cause
a very slight drop in temperature across the wall. When heat
transfer through the steam-side surface is considered, the effect of
deposits is reversed. Steam layers and scales insulate the metal
from the cooling effects of the steam. It results in reducing of heat
transfer into the steam and increasing of metal temperatures.
Starr, Castle, and Walker (2004) also described that oxidation on
the steam side of the tubing can induce premature failures due to
* Corresponding author. Tel.: þ60 3 89212213; fax: þ60 3 89212116.
E-mail address: judha@uniten.edu.my (J. Purbolaksono).
0950-4230/$ – see front matter Ó 2009 Elsevier Ltd. All rights reserved.
doi:10.1016/j.jlp.2009.06.005
the insulating effect of the oxide scales raising tube temperatures.
In addition, scale spallation could also increase tube temperatures,
as spallation debris may collect in the bottom of tubes, blocking
steam flow. Attention is drawn to a potential problem in which the
tube temperature and rate of oxidation increase with time as the
oxide builds up.
Chaudhuri (2006) described some aspects of metallurgical
assessment of boiler tubes. He discusses some failure problems in
carbon steel, reheater and superheater tubes. Ray et al. (2007)
reported remaining life assessment and creep analysis of superheater
and reheater tubes made of 2.25Cr–1Mo steel of a thermal power
plant. The tubes had operated for 17 years with average operating
temperature of 540 C and having design pressure of 40 MPa. The
remnant life is predicted through dimensional, hardness and tensile
measurements. Viswanathan, Foulds, and Roberts (1988) performed
estimation on the temperature of reheater and superheater tubes in
fossils boilers. They made correlation between hardness and Larsen–
Miller parameter for 1Cr–½Mo, 2¼Cr–1Mo and 9Cr–1Mo steels.
Finite element simulations have been used to investigate failures
in superheater tube (Othman, Purbolaksono, & Ahmad, 2009) and
reheater tubes (Purbolaksono, Hong, Nor, Othman, & Ahmad, 2009).
Othman et al. (2009) simulated the deformed superheater tube
using the finite element method. The simulation results have a good
conformity with the finding from the visual site inspection. Purbolaksono et al. (2009) reported evaluation on reheater tube failure.
The geometry and the scale thickness of the as-received failed tube
J. Purbolaksono et al. / Journal of Loss Prevention in the Process Industries 23 (2010) 98–105
Nomenclature
Cp
D
G
h
k
L
o
m
N
Nu
P
Pr
Re
S
T
DT
specific heat at constant pressure (J kg1 C1)
inner diameter of the tube (m)
Gas mass velocity
convection heat transfer coefficient (W m1 C1)
thermal conductivity (W m1 C1)
length of the tube (m)
mass flow rate (kg h1)
number of tube
Nusselt number
Larsen–Miller parameter
Prandt number
Reynold number
pitch
temperature ( C)
Increasing of metal temperature ( C)
were measured and used to generate the finite element models. The
scale thickness inside tube over service time is considered as
a linear scale growth. Results obtained from the simulations have
shown agreement with the result from the microscopic examination. Both results showed that the failed reheater tube had overheating for prolonged period of time.
Nowadays, a large percentage of power and chemical plants
worldwide have been in operation for long durations. There are strong
economic reasons and technical justifications for continued operation
of the power plants. In order to realize the continuing operations in
practice, however, appropriate techniques and methodologies are
needed to evaluate the current condition of the plant components and
to estimate their remaining useful lives. The techniques should also be
valuable with respect to relatively younger power plants in the
context of safety, availability and reliability, operation, maintenance
and inspection practices. An important ingredient in the continuing
operation of power plants is the remaining life assessment technology
(Viswanathan, 1989). The remaining life estimations could help in
setting up proper inspection schedules and operating procedures in
order to avoid premature retirement of the plants.
An accurate prediction of the temperature distribution in tube
metal of the superheater and reheater will aid the power plant
inspectors or engineers in evaluating the remaining life of the boiler
tubes. A continually increasing scale thickness may occur on the inner
surface of superheater and reheater tubes during the service.
Presence of oxide scales on the inner surface of boiler tubes may
significantly contribute to the increased tube metal temperature.
Consequently, in the prolonged exposure this phenomenon will
worsen situation that leads to potential tube rupture problems. When
a power plant is forced to shut down because of the single component
failure, the cost of the lost electric-power generation can run to
several hundred thousand dollars a day. It is essential to perform life
assessments through the operational condition-based monitoring of
the power plant regularly than allowing the equipments to fail. In this
work the life assessments are performed by using finite element
simulations and utilizing the empirical formulae through iterative
procedures. The first empirical formula is correlating scale thickness
with Larsen–Miller parameter (Rehn, Apblett, and Stringer,1981). The
second empirical formula is correlating the experimental hardness
data for 2¼Cr–1Mo steel with Larsen–Miller parameter (Viswanathan et al., 1988). The finite element analysis is carried out using
software package of ANSYS (ANSYS, 2008). Finite element models for
heat transfer analyses, that involve forced convections on the inner
surface due to the turbulent flow of steam and on the outer surface
t
W
X
99
time (h)
gas flow (kg h1)
scale thickness (m)
Greek symbols
viscosity (kg m1 s1)
density (kg m3)
m
r
Subscripts
aveh
average for considering hardness
aves
Average for considering scale
g
gas
i
inner
o
outer
s
steam
t
transverse
due to cross flow of the hot flue gas over bare tubes, are carried out in
order to determine temperature distribution in the tube. An iterative
procedure is performed to determine the average temperature and
hardness of the tube steel over period of time as oxide scale thickness
on the inner surface increases. Two failure cases in reheater (Case 1)
and superheater (Case 2) tubes are evaluated through the finite
element simulations and iterative procedures. The reheater and
superheater tubes are made of a typical material of SA213-T22 steel.
As the oxide scales are increasingly developed on the inner surface,
the increasing of temperature and decreasing of hardness in
tube metal for both cases are determined. The remnant life estimations are then made in the form of creep cumulative damages. Besides
computer simulations utilizing parameters of the operational
condition may economically support the post-failure assessment
method, i.e. visual inspections, metallurgical examinations and
mechanical strength measurements, the estimations obtained from
the simulations may provide an advanced warning to take preventive
actions prior to failure.
2. Heat transfer parameters
Model of the tube section used is 25 mm in length. In modeling
of the steady state heat transfer for the problem using ANSYS
(ANSYS, 2008), the area of the model is divided into two regions, i.e.
scale region and tube region (see Fig. 1). The steam region is taken
into account in determining the convection coefficient of steam
film for fully developed turbulent flow in a circular tube.
The chemical composition of SA213-T22 steel is listed in Table 1.
Steam properties and the thermal conductivities for SA213-T22 and
oxide scale (magnetite) are shown in Table 2. The steam-side scale
is usually reported to be duplex (inner spinel (Fe–Cr–Mo)3O4 layer
and outer magnetite (Fe3O4) layer) or triplex (inner spinel layer,
middle magnetite layer and outer hematite (Fe2O3) layer). In this
study material of the scale is treated to be all magnetite.
Phenomenon of heat transfer inside the boiler tube is considered
as forced convection with turbulent flow. Correlation for fully developed turbulent flow in tube is expressed as (Incropera & DeWitt, 1996)
Nus ¼ 0:023ðRes Þ0:8 ðPrs Þ0:4
(1)
where Res is Reynolds number that may be expressed as
o
Res ¼ 4ms :
pDi ms
(2)
100
J. Purbolaksono et al. / Journal of Loss Prevention in the Process Industries 23 (2010) 98–105
Table 2
Properties of fluid and solid materials.
Inlet steam properties (Incropera & DeWitt, 1996)
Case 1
Temperature, 576 C
Thermal conductivity
0.0636 W/m C
Specific heat
2185 J/kg C
Dynamic viscosity
2.965 e-05 N s/m2
Mass flow rate
3600 kg/h
hot gas
oxide
scale
Case 2
Temperature, 540 C
0.0604 W/m C
2161 J/kg C
2.834 e-05 N s/m2
7200 kg/h
Water wall properties (French, 2000)
Tube material
SA213-T22
Thermal conductivity
34.606 W/m C
tube metal
Fe3O4 iron oxide (magnetite) (French, 2000)
Thermal conductivity
0.592 W/m C
in which Cpg and kg are specific heat and thermal conductivity of
the flue gas, respectively. The corresponding Reynolds number Reg
may be expressed as
steam
Reg ¼
remaining tube
metal thickness
hollow
radius
o
in which ms is mass flow rate of the steam; Di is the inner diameter
of the tube; ms is steam viscosity, and Prs is its Prandtl number that
is defined as
ms Cps
(3)
ks
in which Cps and ks are specific heat and thermal conductivity of the
steam, respectively.
Convection coefficient of steam film for fully developed turbulent flow in circular tube is expressed as (Incropera & DeWitt,
1996):
hs ¼ 0:023
ks
ðRes Þ0:8 ðPrs Þ0:4
Di
(4)
where ks is steam conductivity.
Heat transfer outside the boiler tube is considered as forced
convection due to cross flow of the hot flue gas over bare tubes.
A conservative estimated convection coefficient of flue gas hg on
outer surface of bare tube in inline and staggered arrangements
(see Fig. 2) is given by (Ganapathy, 2003)
0:6 0:33
12kg hg ¼ 0:33
Reg
Prg
D0
(5)
where kg is flue gas conductivity; D0 is outer diameter of the tube;
Prg is defined as
Prg ¼
(7)
where G is gas mass velocity and may be defined as
initial tube metal thickness
Fig. 1. Model of the superheater and reheater tubes with scale on the inner surface.
Prs ¼
GD0
12mg
mg Cpg
(6)
kg
Table 1
Chemical compositions of SA213-T22 steel.
Code
C
Si
Mn
P, max
S, max
Cr
Mo
SA213-T22
0.05–0.15
0.5
0.3–0.6
0.025
0.025
1.90–2.60
0.87–1.13
Wg
G ¼ 12
Nw LðSt D0 Þ
(8)
in which Wg is gas flow; Nw is number of tube wide; St is transverse
pitch (see Fig. 2), and L is the tube length.
However, the evaluation will be more accurate if the nonluminous coefficients for water vapour and carbon dioxide are
taken into account in determining the gas side convective heat
transfer coefficient. Hence, there will be an additional increase
on the overall heat transfer coefficient at flue gas temperature of
800–900 C by 5–6%, assuming that the external or direct radiation
from the furnace is absent.
3. Estimations of temperatures, scale thickness and hardness
The increasing scale thickness may occur in superheater and
reheater tubes during the service. Therefore, estimation must be
made of the average temperature in the oxide scale as a function of
time and scale thickness. In this work in order to perform a scale
growth prediction, steam-side scale formation for ferritic steel of
1–3% chromium correlated with the Larson–Miller parameter as
reported by Rehn et al. (1981) is utilized (see Fig. 3). The data of
Fig. 3 may be approximated as
log
X
¼ 0:00022P 7:25
0:0254
(9)
where X is scale thickness in mm.
In the Larson–Miller method, time and temperature are related
by the following equation:
P ¼
9
T þ 492 ðC þ Log tÞ
5
(10)
where P is the Larson–Miller parameter; T is the absolute temperature in degree Celsius; t is the service time in hours; C is a constant
equal to 20.
Correlation between hardness (HV) and the Larsen–Miller
parameter for 2¼Cr–1Mo steel in the as-quenched condition (see
Fig. 4) may be expressed as (Viswanathan et al., 1988)
Hardness ðHVÞ ¼ 961:713 0:020669P
(11)
J. Purbolaksono et al. / Journal of Loss Prevention in the Process Industries 23 (2010) 98–105
101
Fig. 2. Inline and staggered arrangements of the bare tubes.
The increasing of scale thickness DX in the reheater or superheater tubes (SA213-T22) may be obtained from Eq. (9) corresponding to the given running hours and average temperature in
the oxide scale. The steps of time as shown in Table 3 are used in the
iterative procedures. The iterative procedures used to determine
scale thickness and hardness of the tube as a function of time and
temperature are as follows:
3.1. Procedure 1
tube, the average temperature of Taveh_i is determined from average
of the temperatures on the inner and outer surfaces of the tube. Eqs.
(9) and (10) are used to calculate the scale thickness of Xia for the
service hours of 1 h and the scale thickness of Xib for the service
hours of 250 h (see Table 3) using Taves1. Next, by subtracting one
from the other, the scale increase of DXi (¼Xib Xia) is determined
and a new scale thickness of Xi (¼X0 þ DXi) is obtained. Eqs. (10) and
(11) with Taveh_i are used to determine the hardness of Hia for the
service hours of 1 h and the hardness of Hib for the service hours of
250 h. Hardness Hi is equal to Hia.
For initial step i ¼ 1, the design temperature for the steam is set
to Ts at the inlet of reheater or superheater tube. From the
numerical simulation in the absence of scale (X0), the average
temperature of Taves_i is the temperature on the inner surface of the
Fig. 3. Steam-side scale formation for ferritic steels of 1–3% chromium correlated with
the Larsen–Miller parameter (Rehn et al., 1981).
Fig. 4. Correlation between hardness and the Larsen–Miller parameter for 1Cr–½Mo,
2¼Cr–Mo and 9Cr–1Mo steels (Viswanathan et al., 1988).
102
J. Purbolaksono et al. / Journal of Loss Prevention in the Process Industries 23 (2010) 98–105
Table 3
Steps of time used in the iterative procedure.
Table 5
Parameters used to determine gas mass velocity G for validation of actual data.
Step
Hours
1
2
3
4
5
6
7
8
9
10
11
12
250
500
1000
2500
5000
10,000
20,000
40,000
60,000
80,000
100,000
120,000
3.2. Procedure 2
Set i ¼ i þ 1. The average temperatures of Taves_i and Taveh_i are
then determined from the numerical modeling with the new scale
thickness on the inner surface. The average temperature of Taves_i
obtained from the average of the temperatures on inner surface
temperature and at the scale/metal interface is then used to
calculate the incremental thickness of scale from 250 to 500 h using
Eqs. (9) and (10). For service hours of 500 h, P is calculated using Eq.
(10) and Xib is found from Eq. (9). For service hours of 250 h, P is
calculated using Eq. (10) and Xia is found from Eq. (9). Subtracting
one from the other (Xib Xia) produces the incremental scale
formation from 250 to 500 h, which is added to Xi to give a new
scale thickness of Xi. The average temperature of Taveh_i obtained
from the average of the temperatures at scale/metal interface and
on the outer surface is then used to calculate the hardness of the
tube metal for service hours of 250 h and 500 h using Eqs. (10) and
(11). For service hours of 500 h, P is calculated using Eq. (10) and Hib
is found from Eq. (11). For service hours of 250 h, P is calculated
using Eq. (10) and Hia is found from Eq. (11). Hardness for service
hours of 250 h (¼Hi) may be determined from the average of Hia and
Hib. Repeat Procedure 2 for further predictions up to the required
hours with the steps of time shown in Table 3.
Since the initial increment of time determines the further estimation results, it is proposed to use the steps of time as shown in
Table 3. A smaller increment of time might provide a better estimation, whereas a bigger increment of time for initial iteration may
be resulting in inaccuracy estimation or less conservative prediction. However, for estimations starting from the service hours of
20,000 h, an increment of time may be proposed to be taken at
every 20,000 h.
4. Finite element modeling
The geometry of the as-received tubes and heat transfer
parameters governing the problem are used to generate the finite
element models. It is important to note that all the geometrical
units used for modeling are in m. Hence, the meshing size control of
0.0001 is used to generate the 2D solid triangular elements in order
to allow the model having appropriate size of elements. The
properties of the elements are then defined as 2D-axisymmetric
solid elements.
Table 4
Geometry of the as-received tubes, service time and the inner scale thickness.
Case
1
2
Inner
radius, m
Tube
thickness, mm
Service
time, h
Scale
thickness, mm
Date of
failure
0.0219
0.0163
3.5
4.0
117,522
115,494
0.58
0.2
10/10/2003
04/08/2003
Gas flow, kg/h
Number of tube wide
Transverse pitch, m
Tube length, m
Case 1
Case 2
500,000
50
0.1016
8
800,000
50
0.1016
8
As described in Procedure 1, the finite element model is produced
in the absence of oxide scale at the initial step (i ¼ 1). The bulk
temperature and convection coefficient hg of the flue gas are applied
on the right edge of the model (Fig. 1). Next, the bulk temperature
and convection coefficient hs of the steam are applied on the left
edge of the model. In the presence of the oxide scale (i>1) the model
will have two domain areas, i.e. scale and tube metal. In order
to make connectivity of the domain areas at scale/metal interface,
a merge-size control of 0.00001 is used. The merge-size control
should considerably be smaller than the meshing size control. The
bulk temperature and convection coefficient hg of the flue gas are
applied on the right edge of the tube metal region and the bulk
temperature and convection coefficient hs of the steam are applied
on the left edge of the oxide scale region. The convection coefficients
of hs and hg are obtained by using Eqs. (4) and (5) respectively.
5. Case studies
The most common approach for calculating the cumulative
creep damage is computing the amount of life expended by using
time fraction as measures of damage. When the fractional damages
add up to unity, then the failure is postulated to occur. The most
prominent rule is given as (Robinson, 1938)
Xt
si
tri
¼ 1
(12)
where tsi is the service time and tri is the time to rupture.
The actual data of the available reports on the failed reheater
(Ahmad, 2004) and superheater (Ahmad, 2003) tubes at Kapar
Power Station Malaysia are used for the case studies. Detailed
descriptions of the failed tubes are as follows:
- Operating steam temperature of the reheater and superheater
tubes are 576 C and 540 C respectively. The average operating
steam pressures in Case 1 and Case 2 are 40 bar (4.0 MPa) and
100 bar (10 MPa) respectively.
- The average flue gas temperatures were reported at around
800 C for the reheater tube and 850 C for the superheater tube.
The detailed data taken from the reports are shown in Table 4.
Parameters used to determine gas mass velocity G and the estimated convection coefficients hs and hg for the internal and
external surfaces are shown in Tables 5 and 6 respectively. Fig. 5
shows the isometric view of a half expansion model from 2D
axisymmetric model and the temperature distribution of Case 2 for
the service hours of 115,494 h. It can be seen that the temperature
on the external surface may be estimated at around 574 C.
Diagram of Larsen–Miller parameter with stress variation to
rupture of annealed material 2.25Cr–1Mo steel (ASTM) (Smith,
Table 6
The estimated convection coefficients hs and hg for internal and external surfaces
respectively.
Case
hs, W/m2 C
hg, W/m2 C
1
2
2053.65
6084.32
126.01
129.99
J. Purbolaksono et al. / Journal of Loss Prevention in the Process Industries 23 (2010) 98–105
103
Fig. 5. Temperature ( C) distribution of Case 2 for the service hours of 115,494 h.
1971) as shown in Fig. 6 is utilized to determine the rupture time. In
order to obtain conservative estimation, the minimum curve
correlating stress variation and Larsen–Miller parameter is used.
The Larsen–Miller parameters for Case 1 and Case 2 may be
obtained from Fig. 6. It is shown that the parameters for Case 1 and
Case 2 are 39,900 and 37,800, respectively. The estimated life for
the tubes in Case 1 and Case 2 are determined in the form of the
cumulative creep damage as expressed in Eq. (11). If the total of the
fractional damages is greater or equal to 1, the failure is predicted to
occur. It can be seen from Fig. 7, the estimated lives for both cases
are around 6–15% higher than that of the actual data. The estimated
results are reasonably shown to be in agreement with the actual
data. The estimated scale thickness as shown in Table 7 is also
shown to be in agreement with the actual data (Table 4). The
increasing of the average temperatures of Case 1 is significantly
higher than that in Case 2 as shown in Fig. 8. It results in the
estimated hardness values for Case 1 showing moderately less than
Case 2. It is believed that the geometry of the tube and mass flow
rate of steam influence the phenomena of heat transfer. It is
important to note that if the applied internal steam pressure is
higher, the mass flow rate of the steam has also to be considerably
high. Insufficient mass flow rate of steam may cause a significant
increase of the metal tube temperature and scale growth. This
feature may be seen from the conditions of Case 1 and Case 2. It can
be seen from Table 7 and Fig. 8, the increased tube temperatures
would accelerate the developments of the oxide scale on the inner
surface. The relation between the hardness values and the average
temperature of the tube over period of time can be used for future
references. The strength of the low-alloy steel will change with the
increasing of temperature over period of time.
The life estimations through computer simulations utilizing
parameters of the operational condition may provide an advanced
warning to take preventive actions prior to failure. Poor mass flow rate
of steam, higher operational steam temperature and higher convective
Fig. 6. Diagram of Larsen–Miller parameter with stress variation to rupture of
annealed material 2.25Cr–1Mo steel (ASTM) (Smith, 1971) (1 ksi ¼ 6.895 MPa).
104
J. Purbolaksono et al. / Journal of Loss Prevention in the Process Industries 23 (2010) 98–105
1.4
325
305
1.2
610
285
Tubes are predicted to failure
Hardness-Case 2
1
Hardness-Case 1
265
Hardness, HV
Case 1 - reheater tube
0.8
Case 2 - superheater tube
0.6
0.4
600
Temperature-Case 2
Temperature-Case 1
245
590
580
225
205
570
Temperature, °C
Cumulative creep damage
620
185
560
0.2
165
0
145
0
20000
40000
60000
80000
100000
550
120000
125
Service hours
0
20000
40000
Table 7
Estimations for scale thickness for Case 1and Case 2.
Case 1 (reheater tube)
Case 2 (superheater tube)
Service hours
Scale thickness, mm
Service hours
Scale thickness, mm
0
250
500
1000
2500
5000
10,000
20,000
40,000
60,000
80,000
100,000
117,522
0.0000
0.0556
0.0736
0.0965
0.1371
0.1783
0.2315
0.3008
0.3916
0.4584
0.5133
0.5610
0.5982
0
250
500
1000
2500
5000
10,000
20,000
40,000
60,000
80,000
100,000
115,494
0.0000
0.0229
0.0301
0.0392
0.0550
0.0707
0.0907
0.1163
0.1491
0.1728
0.1920
0.2085
0.2200
80000
100000
540
120000
Service hours
Fig. 7. The estimated cumulative creep damages.
coefficient and temperature of flue gas would lead to increasing tube
metal temperature and oxide scale thickness. Tube metal temperature
may increase slowly over many years or rapidly over a few hours.
Internal oxide scales usually result in long-term overheating that
gradually increase the tube metal temperature. The increasing tube
metal temperature and oxide scale growth have a reciprocal casual. As
the tube metal temperature increase, so does the rate of internal scale
formation. As the oxide scale thickness increases, so does the tube
metal temperature. The cycle continues progressively over period of
time. Loss of boiler feedwater or impaired steam flow usually results in
rapid overheating and often rapid failure. Localized high temperature
flue gas flow definitely causes localized overheating of the boiler tube.
These features potentially lead to rupture of the boiler tubes that
consequently leads to unscheduled and costly outages.
It is an essential element of condition-based maintenance in
which the equipment is maintained on the basis of its condition, or it
allows actions to be taken to avoid the consequences of failure, before
the failure occurs. It is typically much more cost-effective than
allowing the components to fail. By selecting a physical measurement
which indicates that deterioration is occurring, the readings on the
parameter need to be taken at regular interval. Since failure occurs to
individual components, the monitoring measurements need to focus
on the particular failure modes of the critical component. If a power
plant has been operating with breakdown maintenance or regular
planned maintenance, a change over to condition-based maintenance can result in significant improvements in plant availability and
60000
Fig. 8. The estimated hardness and the average metal tube temperatures over period
of time.
in reduced cost. Then, life estimations of the boiler tubes through the
computer simulations may be employed from time to time.
Failure analysis on the water-tube boiler can also be performed
by referring to Section 2, Part D of The ASME Boiler and Pressure
Vessel Code (ASME, 1998) for the maximum allowable stress values
and data from United States Steel Corporation (1972) as shown in
Table 8. The maximum allowable stresses and the operational hoop
stresses of the tubes (Case 1 and Case 2) over the service hours are
then plotted in Fig. 9. It can be illustrated that the critical conditions
may be predicted when the stresses in the tube exceeding the
maximum allowable stress. It can also be seen from Fig. 9 that the
tube can be estimated when the overheating condition occurs.
Since the operating hoop stresses are higher than the allowable
stresses, tubes in Case 1 and Case 2 are estimated to be critical at
the beginning of service. Even though the operating hoop stresses
remain much lower than the ultimate strength shown in Table 8,
a careful consideration needs to be taken. Monitoring on the
temperature increase over period of time is particularly important
when overheating in the tube is concerned. Thus, combination of
estimating the life expectancy using the cumulative creep damage,
considering the maximum allowable stresses and monitoring the
increasing of temperature in tube metal may provide better
monitoring of the boiler tubes. Appropriate preventive actions may
be taken during the scheduled outage for close inspections and
necessary examinations in order to avoid forced outage due to
a single failure of a component.
It is reported that the microscopic examination for Case 1 and
Case 2 were also carried out to support the life assessment. The
finding of Case 1 indicated that the tube metal microstructure had
a complete stage of spheroidization where the carbide particles
Table 8
The maximum allowable stresses and tensile strengths for seamless tube SA213-T22.
Temperature, C
Max. allow. stress,
MPa (ASME, 1998)
Tensile strength, MPa (United
States Steel Corporation, 1972)
537.78
565.56
593.33
621.11
648.89
784.00
55.16
39.30
26.20
16.55
9.65
–
377.16
–
282.69
–
–
151.69
J. Purbolaksono et al. / Journal of Loss Prevention in the Process Industries 23 (2010) 98–105
105
6. Conclusions
50
620
40
600
30
590
580
20
570
Temperature - Case 2
Temperature - Case 1
Max. allowable stress (Case 2)
Max. allowable stress (Case 1)
Operational pressure, 4 MPa
Operational pressure, 10 MPa
560
550
540
0
20000
40000
60000
80000
100000
Hoop stress, MPa
Average temperature, °C
610
10
0
120000
Finite element simulations incorporated with the iterative procedures may be used to estimate the increased temperature and
decreased hardness values of the tube metal and development of oxide
scale on the inner surface of boiler tubes over prolonged period of time.
Two failure cases in reheater and superheater tubes were evaluated.
Estimations obtained from the simulation were shown to be in
agreements with the actual data. Combination of estimating the life
expectancy using the cumulative creep damage, considering the
maximum allowable stresses and monitoring the increasing of
temperature in tube metal may provide an advanced warning to take
preventive actions prior to failure. A reciprocal casual of increasing
tube metal temperature and oxide scale growth in a hazardous environment of boiler operation would lead to tube leakage/rupture, by
which the safety and cost-effective of the thermal power plant is
influenced seriously. The ruptured boiler tubes in power plant generate
a serious problem that often leads to unscheduled and costly outages.
Service hours
Fig. 9. Operational hoop and maximum allowable stresses of the failed tubes for Case 1
and Case 2 at the corresponding average tube temperatures and service hours.
have coalesced and dispersed uniformly as shown in Fig. 10. Similar
feature of the microstructure was also reported for Case 2.
The change of microstructure confirmed that the failed tube had
operated at higher temperature or experienced overheating for
prolonged period of time. It agrees with the estimation obtained
from the simulations.
Abnormal governing parameters clearly cause loss of expected life
due to microstructural changes and creep failures as result of overheating. Well-informed data for heat transfer parameters according
to variations in the operating conditions from the monitoring system
will improve estimations obtained from the iteration procedure. In
other words, better estimation of the average temperature in the tube
metal could be obtained, provided that all the heat transfer parameters used are well specified. Early warning of failure may be monitored from time to time through simulations utilizing data of the
operational heat transfer parameters taken from the monitoring
devices. Regular measurements need to be taken, e.g. for months or
years, before a critical situation arises. It is expected that more
detailed evaluation can indicate the nature of the problem so that
rectification action can be planned.
Fig. 10. Complete stage of spheroidization with the coalesced carbide particles to form
chain of pores (Ahmad, 2004).
Acknowledgements
This work is supported by the Ministry of Science Technology
and Innovation, Malaysia through the research projects of IRPA
09-99-03-0033 EA001 and Sciencefund 04-02-03-SF0003. The
author wish to thank Universiti Tenaga Nasional and Kapar Energy
Ventures Sdn. Bhd for permission of utilizing all the facilities and
resources during this study.
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