3.16

Ceramic Fuel–Cladding Interaction

K. Maeda
Japan Atomic Energy Agency, O-arai, Ibaraki, Japan

ß 2012 Elsevier Ltd. All rights reserved.

3.16.1

Introduction and Overview of Ceramic Fuel–Cladding Interaction

444

3.16.2
3.16.2.1
3.16.2.2
3.16.3
3.16.3.1
3.16.3.1.1
3.16.3.1.2
3.16.3.2
3.16.4
3.16.4.1
3.16.4.1.1
3.16.4.1.2
3.16.4.1.3
3.16.4.1.4
3.16.4.1.5
3.16.4.2

3.16.5
3.16.5.1
3.16.5.2
3.16.5.2.1
3.16.5.2.2
3.16.5.2.3
3.16.5.3
3.16.5.3.1
3.16.5.3.2
3.16.5.3.3
3.16.6
3.16.7
3.16.7.1
3.16.7.1.1
3.16.7.1.2
3.16.7.1.3
3.16.7.1.4
3.16.7.2
3.16.7.2.1
3.16.7.2.2
3.16.7.2.3
3.16.8
References

Cladding Compatibility with Oxide Fuels and FPs
Formation of Protective Oxides on Cladding Materials
Chemical Interaction Among Oxide Fuels, FPs, and Cladding
Morphology of Cladding Attack in Oxide Fuel Pins
Observations of Cladding Attack
Deep localized cladding attack

FCCI at the top of the fuel column
Types and Characteristics of Cladding Attack
Occurrence of Interaction Between Oxide Fuels and Cladding
Key Parameters in FCCI Development
Fuel parameters
Effect of temperature
Effect of burnup
Effect of temperature difference between fuel and cladding
Effects of cladding materials
FCCI Model and Wastage Equation
Mechanism of Oxide Fuel and Cladding Interaction
Oxygen Potential of Irradiated Fuel
Characteristics of Major Corrosive FPs
Iodine
Cesium
Tellurium
Various Corrosion Reaction Mechanisms
Corrosion early in life
Iodine transport of cladding constituents
Cladding corrosion by Cs–Te mixture
Inhibition Methods for Oxide Fuel and Cladding Interaction
Nonoxide Ceramic Fuels and Cladding Interaction
FCCI of Carbide Fuel
Chemical reactions with FPs
Formation of intermetallic compounds
Clad carburization
Key parameters of clad carburization
FCCI of Nitride Fuels
Chemical reactions with FPs
Formation of intermetallic compounds

Clad nitriding
Outlook

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445
446
447
447
447
447
448
449
449
449
450
451
452
455
455
457
457
458
459
459
460
461
461
462
463
466

467
467
467
467
468
470
475
475
476
476
477
478

Abbreviations
AISI
ANL

American Iron and Steel Institute
Argonne National Laboratory

CCCT
C/M
DFR

Cladding component chemical transport
Carbon-to-metal
Dounreay fast reactor

443



444

Ceramic Fuel–Cladding Interaction

EBR-II
EPMA
FAE
FCCI
FCMI
FFTF
FP
FPLME
FR
GE
HEDL
JOG
LWR
MOX
MX
NMA
N/M
O/M
PFR
PIE
PNC

RIFF
SIMS
SNR


Experimental breeder reactor-II
Electron probe microanalysis
Fuel adjacency effect
Fuel and cladding chemical interaction
Fuel–cladding mechanical interaction
Fast flux test facility
Fission product
Fission product-induced liquid-metal
embrittlement
Fast reactor
General Electric Company
Hanford Engineering Development
Laboratory
Joint oxide-gaine (French)
Light water reactor
Mixed oxide
Nonoxide, where M stands for U þ Pu and
X stands for C or N
Nuclear microprobe analysis
Nitrogen-to-metal
Oxygen-to-metal
Prototype fast reactor
Postirradiation examination
Power Reactor and Nuclear Fuel
Development Corporation, currently
Japan Atomic Energy Agency
Re´action a` l’Interface Fissile Fertile (French)
Secondary ion mass spectrometer
Schneller Natriumgeku¨hlter Reactor

(German)

3.16.1 Introduction and Overview of
Ceramic Fuel–Cladding Interaction
Ceramic fuels used for fast reactors (FRs) are oxide
fuels and nonoxide ceramic (MX-type, where M stands
for U þ Pu and X stands for C and N) fuels such as
carbide and nitride fuels. Ceramic fuel–cladding interactions (FCCI, fuel and cladding chemical interaction)
are mainly divided according to oxide and MX-type
fuels. FCCI is more complicated in oxide fuels than
in MX-type fuels because oxide fuels lead to oxidation
of cladding materials and formation of various oxides
of fission products (FPs) by irradiation.
Chemical interactions between uranium and plutonium mixed oxide (MOX) fuels and/or FPs and
cladding materials are considered as one of the major
factors limiting the lifetime of fuel pins in FRs. This
limitation is especially important for long-term irradiation of (U,Pu)O2 fuel pins clad in stainless steel.

Fuel pins for FRs are generally designed with Type
316 stainless steel cladding to operate with a peak
cladding hot-spot temperature of 700  C.
Results and analyses of irradiation experiments
related to FCCI with oxide fuels have been reported
in various technical society conferences and topical
and periodic reports since FCCI was first reported in
the late 1960s. FCCI is nowadays recognized as one
of the major factors determining integrity and lifetime of oxide fuel pins as demonstrated in numerous
in-pile and out-of-pile tests. Some mechanisms of
cladding attack have been proposed from the results
of the many postirradiation investigations and thermodynamic analyses of the postulated chemical reactions.

Cladding and oxide fuels do not violently react,
even under a high oxygen potential condition; they
only form a protective layer on the inner wall of
the cladding. But when fuel pins are irradiated in a
reactor, the additional effect of the generated FPs
induces cladding attack. A number of experiments
have shown that both stoichiometric and hypostoichiometric oxide fuels react with stainless steel cladding when irradiated in typical FRs. On the other
hand, out-of-pile tests between (U,Pu)O2 or (U,Pu)
O2Àx and several stainless steels have shown that no
detectable reaction took place within the times and
temperatures of interest for FRs. When hyperstoichiometric fuel, (U,Pu)O2 þ x , was tested, cladding attack
was detected and the difference in reaction behavior
was ascribed to the excess oxygen provided in the
hyperstoichiometric fuel. Inability to reasonably
extrapolate the out-of-pile results to the in-pile results
is of concern for the design of oxide fuel pins, and it
indicates that the prediction of lifetime is complicated.
The thermochemistry of the fuel–cladding gap is
complex as well and difficult to predict because it
depends not only on concentrations of corrosive FPs,
but also on major parameters such as fuel–cladding
gap width, fuel oxygen-to-metal (O/M) ratio, cladding
temperature, fuel temperature, and radial temperature
gradient. FCCI is further regarded as sensitive to
linear heat rating and likely to change with fuel
burnup. When the swelling of the cladding is high
and the fuel–cladding temperature gap is large, the
probability of attack is enhanced. Thus, cladding attack
tends to be unpredictable, and it may be locally worse
compared to the overall condition. The possible consequence is complete penetration of the cladding by a

chemical mechanism alone. In addition to this, it may
be considered that there is some form of stress corrosion cracking in the cladding. Actual creep strain on
the cladding is from fuel and cladding mechanical


Ceramic Fuel–Cladding Interaction

interaction (FCMI) and/or internal FP gas pressure.
FCCI has been identified as a contributing factor in
the breaches of oxide fuel pins. Observations at an
axial location of a breach that was located at the
approximate original top of the fuel column have
shown extensive FCCI. The breach was a consequence
of FCMI and internal FP gas pressure.
As explanations of the observed effects of FCCI
have been speculative, fuel pin design could rest
only on empirical equations rather than on fundamental models. Cladding wastage equations by FCCI have
been developed for fuel pin designs. Most observations
of FCCI showed it to be the result of simple oxidation
of the inner surface of the cladding. Three principal
types of cladding attack in stainless steel can be distinguished. The first is a general oxidation of the inner
surface of the cladding. The second is intergranular
attack and is the most important. The third is advanced
attack which appears to be a transport of the cladding
constituents into the fuel. It is typically seen as wastage
of the cladding thickness in some local areas by mechanical or liquid-phase transport of cladding constituents
into the outermost oxide layer on the fuel pellets.
FCCI with oxide fuels has been recognized as an
important factor in the ability to achieve peak burnups in the range of 10 at.% in FRs while maintaining
high coolant bulk outlet temperatures. However, in

addition to cladding thickness losses due to FCCI,
oxide fuels and FPs have the potential for reducing
cladding load-bearing capabilities by mechanisms such
as liquid-metal embrittlement (FPLME, FP-induced
liquid-metal embrittlement).
The other type of FCCI occurs for nonoxide
ceramic (MX-type) fuels such as carbide and nitride
fuels and the cladding. MX-type fuels are chemically
unreactive to sodium coolant, so sodium may also possibly be used as a medium for bonding between fuel and
cladding instead of helium gas. MX-type fuels are
generally irradiated at lower temperatures and lower
radial temperature gradients than oxide fuels, although
at high linear heat rating, which results in low FPs
release rate. The volatile FPs (Br, I, Cs, and Rb) do
not form carbides or nitrides. In particular, MX-type
fuel pins are kept with low oxygen potential at the inner
cladding surface; therefore, severe oxidative FCCI of
the FPs is not expected. A number of irradiation experiments have been performed with MX-type fuels to
study FCCI. The compatibility with cladding materials
has been investigated in out-of-pile examinations and
thermodynamic analyses. As a consequence, unlike the
case of oxide fuels, FPs from MX-type fuels do not play
a major role in FCCI. Instead, the carburizing and

445

nitriding of cladding, and also the formation of intermetallic compounds of fuel and cladding, have been
investigated as a major FCCI of MX-type fuels.
The carbide fuels (U,Pu)C, which are designed to
be slightly hyperstoichiometric, will therefore be

in the two-phase region (U,Pu)C þ (U,Pu)2C3. The
presence of higher carbide phases carburizes the
claddings. Hyperstoichiometric carbides can embrittle the cladding by forming grain boundary carbides
which can lead to intragranular failure of the steel
after a moderate burnup. The creep and swelling
properties of stainless steels are sensitive to carburization and precipitation of M23C6. As sodium can act
as a transfer agent, carbon transport rates through the
gap in sodium-bonded fuel greatly increase. Hypostoichiometric mixed carbides contain (U,Pu) metal as
a second phase which may form low melting-point
eutectics with iron or nickel base cladding alloys.
Hyperstoichiometric MN1 þ x-containing sesquinitride phase can cause nitrogen penetration and
form a reaction layer at the cladding inner surface,
which results in the clad nitriding. The nitriding
of cladding generally decreases the ductility and
increases the mechanical strength. Hypostoichiometric MN1 À x contains free metal leading to a eutectic melting reaction between the free (U,Pu) metal and
the cladding, which results in formation of (U,Pu)Fe2and (U,Pu)Ni5-type intermetallic compounds.
At present, it is clear that the knowledge base for
MX-type fuels is much smaller and less detailed than
that for oxide fuels, and addition to the base is a work
in progress. However, MX-type fuels merit much less
concern regarding cladding–fuel compatibility than
oxide fuels.
In Sections 3.16.2–3.16.6, the causes of FCCI
with oxide fuels are reviewed, considering the dependence on irradiation conditions and fuel parameters
as well as types of cladding material. Furthermore, the
role of corrosive FPs in the FCCI, the mechanism of
FCCI, and the FCCI enhancement by oxygen potential are summarized. The MX-type fuel–cladding
interaction is briefly described in Section 3.16.7.

3.16.2 Cladding Compatibility with

Oxide Fuels and FPs
3.16.2.1 Formation of Protective Oxides on
Cladding Materials
Out-of-pile tests between (U,Pu)O2 or (U,Pu)O2 À x
and several stainless steels have shown that no
detectable reaction took place for exposure times


446

Ceramic Fuel–Cladding Interaction

and temperatures of a typical FR. However, in hyperstoichiometric fuel, (U,Pu)O2 þ x , cladding attack was
detected. Excess oxygen was provided by the hyperstoichiometric fuel, which was considered to cause
the difference in the reaction behavior. Therefore, if
the fuel surface O/M ratio can be maintained just
below exact stoichiometry, oxidation of the cladding
cannot take place.
Of the three major constituents of austenitic stainless steel cladding, Fe, Cr, and Ni, chromium has
the greatest affinity for oxygen and forms the most
stable oxide. Initially, chromium begins to get oxidized when the oxygen partial pressure satisfies the
equilibrium condition of the reaction
À 4=3CrðcladdingÞ þ O2 ðgÞ ¼ 2=3Cr2 O3 ðsÞ;
where the oxygen potential DG O2 ð¼ RT lnpO2 Þ of the
fuel surface reaches À554 kJ molÀ1 at 727  C.1 However, fuel and cladding do not severely react, even
when the oxygen potential is high; they only form a
protective layer on the inner wall of the cladding. The
stable protective Cr2O3 thin layer prevents the fuel
and cladding reaction from becoming thermochemically equilibrated.
In the initial stage of irradiation, the oxygen

potential of the fuel surface rises because of oxygen
redistribution. Excess oxygen, after uranium and plutonium fission in the fuel, leads to an increase in fuel
O/M ratio with burnup. Radial redistribution of oxygen along the fuel radial temperature gradient
enhances the increase of O/M ratio at the fuel surface. It appears unlikely that the oxygen potential at
the entire fuel–cladding interface can be kept low
enough to prevent cladding oxidation throughout
the entire lifetime of the fuel element. Thus, a thin
protective layer of oxide, mainly Cr2O3, soon forms
on the inside surface of the cladding, thereby physically separating the substrate metal from the oxidizing medium. Further growth of this layer requires
that chromium ions diffuse from the substrate metal
to the outer surface of the coating or that oxygen ions
migrate in the opposite direction. The rates of both
these processes are very slow at 727  C because of
the low values of the diffusion coefficients of the ions
in the oxide layer. If the thermochemically stable
uniform layer is breached by mechanical forces or is
dissolved by a component of the oxidizing environment, the substrate metal is exposed to rapid attack.
The integrity of the cladding relies on the kinetics
of the chemical attack in an environment where
oxidation is thermodynamically possible. In addition,
the inner wall temperature of the cladding in
an FR-MOX fuel pin reaches the range at which

the sensitization of stainless steel cladding occurs,
!500  C.2 This suggests that the corrosion resistance
of the stainless steel cladding might become degraded
because of chromium being held in carbide particles
in the cladding.
3.16.2.2 Chemical Interaction Among
Oxide Fuels, FPs, and Cladding

As long as the protective layer stays intact, the stainless
steel cladding is protected from further corrosion.
However, a number of irradiation experiments have
shown that both stoichiometric and hypostoichiometric
fuels reacted with stainless steel cladding. Unlike irradiated fuel, fresh fuel does not corrode stainless steel
cladding to the same extent. It was suggested that
irradiation damage might reduce the effectiveness of
this protective layer. But it was found that the extent
of oxidation did not sufficiently increase while
irradiation damage by fission fragments increased.3
The thermodynamic tendency of oxide fuels is to
oxidize the cladding, and not to violently attack it
in the absence of FPs because of the protection
provided by the oxide film formed on the surface of
the steel.4 However, the protective layer is impaired
by a chemical reaction of reactive FPs and oxygen
with chromic oxide. Such evidence suggests that one
or more of the FPs are responsible for accelerating
the chemical reactions between fuel and cladding in
irradiated fuel pins.3
FCCI is the FP-accelerated oxidative attack of the
cladding that is frequently observed in FR fuel pins
involving reactive FPs such as Cs, Te, and I.4 Specifically, cesium and tellurium are thought to contribute
to the most aggressive intergranular attack modes.4,5
The FCCI phenomenon is generally recognized to be
the result of the oxidation of chromium in the stainless steel cladding under the influence of FPs; cladding attack by Cs2Te has not been considered as an
oxidation mechanism of the cladding materials.
In irradiation experiments, however, the protective
oxide layer is breached in some places and cladding
attack takes place, usually in a few isolated patches

rather than uniformly. Whether a chemical reaction
between components of the irradiated fuel and constituents of the cladding can occur at all is determined by
the thermodynamics of the reactions involved. Local
breakdown of the protective layer and subsequent
corrosion appear to depend on the local accumulation
of observed major FPs, such as Cs and Te or I, which
are considered important corrosive elements.
The generated volatile FPs are released and accumulate at the fuel–cladding gap with increasing


Ceramic Fuel–Cladding Interaction

burnup. When fuel surface oxygen potential exceeds
the threshold necessary for oxygen transport to the
cladding inner surface, excess oxygen and corrosive
FPs can interact with the cladding inner surface leading
to FCCI. Internal wastage of the stainless steel cladding
is related to the complex phenomenon of corrosion
established by the presence of FPs (Cs, I, and Te) and
oxygen at the fuel–cladding interface. The threshold
temperature for cladding attack is around 500  C.

3.16.3 Morphology of Cladding
Attack in Oxide Fuel Pins
3.16.3.1

Observations of Cladding Attack

447


considered as intergranular corrosion accelerated by
sensitization which is seen as the loss of Cr by Cr23C6
precipitation in the grain boundary.9 As this was the
most aggressive form of FCCI observed along grain
boundaries deep in the cladding in the case of initial
O/M ratios above 1.98, this type of attack has been
largely eliminated by using fuel with O/M ratios of 1.98
and below. By utilizing an appropriate lower O/M fuel
associated with longer irradiation for excess oxygen
in the fuel pin, a more uniform matrix interaction
tends to take place.
A combined interaction form, consisting of matrix
FCCI proceeded by intergranular FCCI, occurs in
fuel with moderate O/M ratio and high burnup.2

3.16.3.1.1 Deep localized cladding attack

3.16.3.1.2 FCCI at the top of the fuel column

Regions of chemical reactions between the fuel and
cladding have been generally observed, especially the
hotter cladding temperature regions. Examination of
metallography samples showed occurrence of nonuniform and deep localized cladding attack in irradiated fuel pins.6,7 Cladding attack usually occurred
in an irregular manner over the inner surface of the
cladding and in the case of intergranular attack, its
depth of penetration varied from site to site. In addition, the observed deep localized interaction was
usually of a different type than in the rest of the
sample. When access to the substrate metal was established, cladding attack by FPs occurred, either uniformly or only locally, but in some cases it penetrated
more than 100 mm into the cladding. This would be a
significant reduction of the effective thickness of the

cladding. Despite the reduction in cladding thickness,
actual fuel pin failure has rarely been observed.
The occurrence of a deep localized interaction of
more than 100 mm in depth was observed in a sample
which had an initial O/M ratio larger than 1.98, and
was irradiated to less than 5 at.% burnup at cladding
temperature higher than 650  C.6 That suggests that
this type of interaction occurs primarily in hightemperature regions with relatively low burnup.
This interaction is called deep localized FCCI
and is an intergranular type of cladding attack, characterized by a highly localized reaction product.
Because cladding attack tends to be random, it becomes
locally worse compared to the overall condition. For a
fuel with an initial O/M ratio of 1.99–2.00, there is
evidence that intergranular attack of sensitized stainless
steel cladding occurs in the matrix around the carbide
particles in the grain boundaries.8 Microprobe examinations have shown this area to be depleted in chromium and manganese, with significant quantities of the
FP cesium present in the reaction product. This was

The top of the fuel column at a cladding temperature
near the maximum corresponds to the boundary of
fissile–fertile fuel pellets. At this location, axially
migrated and accumulated volatile FPs react with
the cladding material. Axial isotopic gamma scans
for high burnup pins have shown that there are larger
amounts of cesium in the area of the upper insulator
pellets than in the area of the fuel.10 Because of the
migration of cesium to the cold region in the irradiated fuel pins, cesium peaks are generally found at
both ends of the fuel column.
These accumulations were generally related to the
formation of a phase consisting of U–Cs–O (Cs2UO4)

at the UO2 blanket or insulator pellets, which caused
localized inelastic deformations of the cladding (up to
30 mm) at the fuel–blanket interfaces by a volumetric
change.10,11 But Kleykamp12 confirmed that instead
of Cs2UO4, a cesium uranoplutonate Cs2(U,Pu)4O12
layer was formed on the grain boundaries of the
UO2 blanket pellets in the irradiation experiment.
Furthermore, formation of compounds at the UO2
blanket or insulator pellets led to a severe intergranular attack of the cladding (up to 100 mm) in this
region.10,11 Figure 1 shows ceramographs for a longitudinal section removed from a fuel pin (maximum
burnup 14.5 at.%, 695  C cladding inner surface
temperature, and initial O/M ratio 1.984).6 Both
the depth and character of the FCCI had changed
significantly at the fissile–fertile transition zone.
The maximum depth of cladding attack at the fertile
and fissile fuel pellets was approximately 90 and
135 mm, respectively. A similar localized form of cladding attack occurred at higher temperatures of
>600  C at the fissile–fertile interface.13 This fissile–
fertile interface reaction, termed RIFF (Re´action a`
l’Interface Fissile Fertile, in French), is associated
with migration of volatile FPs to the end of the fuel


448

Ceramic Fuel–Cladding Interaction

column. There is no evidence of RIFF in PE-16 (high
Ni alloy) and EM-12 (ferittic–martensitic steel alloy).
The occurrence of RIFF appears to depend on the

choice of cladding materials.13
The change in character of the cladding attack at the
top of the fuel column suggests a change in the mechanism of chemical interactions at locations of fissile and
fertile fuel pellets. The reaction of fuel and cesium
suggests the presence of a high oxygen potential in
the fuel.14 Large cesium pressures, which are generally
expected in hypostoichiometric fuel, lead to the formation of cesium uranate in the UO2 blanket or insulator
pellets. The FP inventory and the radial temperature

gradient in the region of the UO2 blanket or insulator
pellets are significantly different from those at the
region of the fissile fuel column. The predominantly
radial heat transfer in the upper region of the fuel
column and the absence of heat-generating material
in the UO2 blanket or insulator pellet suggest little or
no thermal gradient across the UO2-cladding gap.
This effect has been taken into account in the design
of other irradiation experiments by reducing the volume
of the blanket pellets.15 The depth of chemical interaction between cladding and fuel outside the blanket–fuel
interface has always been lower than 60 mm.
3.16.3.2 Types and Characteristics of
Cladding Attack

UO2

(U,Pu)O2

Figure 1 Cladding attack in the vicinity of the fissile–fertile
fuel interface. Reproduced from Lawrence, L. A. Nucl.
Technol. 1984, 64, 139–153, with permission from ANS.


(a)

From the metallographic examination of stainless
steel–clad fuel pins irradiated to various burnup levels,
it is generally possible to observe the character of the
evolving cladding attack along the cladding temperature distribution in the fuel pin. The cladding attack
is classified into three types: (1) matrix, (2) intergranular, and (3) combined matrix and intergranular (also
called ‘advanced’ or ‘evolved’).16 Figure 2(a) and 2(b)
show typical photomicrographs of matrix and intergranular types of attack in a fuel pin with Type 316
stainless steel cladding (burnup: 50 MWd kgÀ1 M),
respectively. In addition, Figure 2(c) shows a severe
combined intergranular and matrix attack observed
in the fuel pin with Type 347 stainless steel cladding (burnup: 140 MWd kgÀ1 M).
The first type of cladding attack is a general oxidation and is confined mainly to the shallow inner
surface of the cladding. The entire body of the inner
wall of the cladding is converted to a reaction zone
containing the oxides of Fe, Cr, and Ni. In the regions
of matrix attack, EPMA (electron probe microanalysis) results show a depletion of iron and nickel and

(b)
Matrix

(c)
Intergranular

Combined or evolved

Figure 2 Types of representative cladding attack: (a) matrix, (b) intergranular, and (c) combined or evolved. Reproduced
from Perry, K. J.; Melde, G. F.; McCarthy, W. H.; Duncan, R. N., In Fast Reactor Fuel Element Technology, Proceedings of

Conference, New Orleans, Luisiana, Apr 13–15, 1971; Farmakes, R., Ed.; American Nuclear Society: Hinsdale, IL, 1971;
pp 411–429, with permission from ANS.


Ceramic Fuel–Cladding Interaction

an enhancement of chromium and cesium. Trace
amounts of iodine and tellurium are also observed in
the region of matrix attack.16 When this type of cladding attack evolves, there is a definite segregation of
the cladding constituents in the reaction product layer.
The reaction product of the matrix attack is a mixture
of metal particles and nonmetallic compounds in the
fuel–cladding gap. In addition to the three major constituents of the cladding, the reaction zone contains the
FPs (Cs, Mo, and lesser amounts of I, Te, and Pd). The
reaction zone does not appear to contain the heavy
metals (U and Pu), and neither does the cladding.
The attack on the grains is uniform with no strong
preference for attack along the grain boundaries.
The second type of cladding attack is penetrating
the stainless steel cladding along grain boundaries
and it is the most relevant for fuel pin failure. Intergranular attack occurs where the steel is sensitized.
Such attack on the area of chromium depletion from
a steel layer adjacent to the grain boundaries by
precipitation of carbides at the grain boundaries is
in accord with metallographic observations. Opening
of the grain boundaries from the cladding inner surface indicates that attack has occurred along them.
In addition, metallic and nonmetallic reaction
products are also detected in the fuel–cladding gap.
This indicates that the grains have been chemically
attacked, as evidenced by the roughened surface.

The third type of cladding attack is the combined
matrix and intergranular attack that is characteristically observed in local areas, and is often accompanied
by wastage of the cladding thickness caused by
mechanical interaction or liquid-phase transport of
cladding constituents in the outermost oxide layer
adjacent to the fuel. The dissolution of iron, chromium, and nickel in a medium of liquid cesium and
tellurium present in the fuel–cladding gap is known as
cladding component chemical transport (CCCT). It is
interesting to note that the constituents of the cladding
are not uniformly distributed in the reaction zone.17–19

3.16.4 Occurrence of Interaction
Between Oxide Fuels and Cladding
3.16.4.1 Key Parameters in FCCI
Development
The thermochemistry in the fuel–cladding gap is
complex and is difficult to predict because it depends
not only on concentrations of corrosive FPs, but also
on major parameters such as the fuel–cladding gap
width, fuel O/M ratio, cladding temperature, fuel

449

temperature, and temperature gradient. It is essential
to develop correlations between the loss of cladding
strength and the various parameter groups such as
irradiation conditions and fuel specifications. The
qualitative characteristics of FCCI and the observed
effects of various fuel and irradiation parameters on
FCCI are described next.

3.16.4.1.1 Fuel parameters

Severity and frequency of internal cladding attack
appear to be independent of both fuel form and fuel
density. In the case of vibrocompacted fuel, reduction
in fuel density might contribute to increased FP
release and to eased radial migration via pores in
the fuel, which has a minor influence on the severity
of the cladding attack (see Chapter 2.02, Thermodynamic and Thermophysical Properties of the
Actinide Oxides).
Annular pellet fuel having a theoretical intrinsic
density of approximately 96%, which corresponds to
a smear density of 80%, showed no clear difference in
cladding attack in comparison with vibrocompacted
fuel of the same smear density.20 From the results, a
general increasing trend of the depth of cladding
attack at cladding temperatures above 500  C at
approximately 3 and 5 at.% burnup was indicated in
both pellet and vibrocompacted fuels. On the other
hand, Batey and Bagley21 showed that the cladding
attack in vibrocompacted fuel was significantly more
severe in comparison with pellet fuels. Also, vibrocompacted fuels pins that were irradiated at higher
power levels (57–79 kW mÀ1) showed two to eight
times the depth of cladding attack expected from
fuel irradiations having the same inner surface cladding temperatures.22,23 In contrast, sphere-packed
fuel pins which are loaded with spheres of mixed
UO2–PuO2 and UO2 have exhibited decreased
depths of cladding attack in comparison with pellet
fuels with a similar initial O/M ratio and irradiation
history.24–30 Thus, there is no clear explanation of the

evidence for different cladding attack behavior
caused by different fuel forms.
From the results of metallographic observations,
the influence of the initial O/M ratio on the severity
of cladding attack was emphasized in addition to the
influence of the cladding temperature,14 and it
appeared that the type of the attack was controlled
by the initial O/M ratio. The initial O/M ratio has a
significant influence on the depth of cladding
attack.31–34 Figure 3 shows the influence of initial
O/M ratio on the depth of cladding attack.31,32
The effects of initial fuel stoichiometry on the


450

Ceramic Fuel–Cladding Interaction

Maximum cladding temperature (ЊC)

Depth of cladding penetration (mils)

5.0

4.0

538

593


649

704
127

102

O/M
1.94
1.96
2.00

3.0

76
O/M 2.00

Threshold temperature

2.0

51

1.0

25

O/M 1.96

Depth of cladding penetration (mm)


482

O/M 1.94
0
1300

0
900

1100

1200

Maximum cladding temperature (ЊF)

Depth of cladding penetration (mm)

(a)

1000

Rapsodie I
£ O/M 1.96
³ O/M 1.98

150

(210)


100
³1.98

50
£1.96

0
450
(b)

500
550
600
650
Cladding inner surface temperature (ЊC)

700

Figure 3 (a) Maximum depth of cladding attack as a function of cladding inner surface temperature. The data represent
burnup ranging from 7 to 13 at.%. Reproduced from Weber, J. W.; Jensen, E. D. Trans. Am. Nucl. Soc. 1971, 14, 175–176,
with permission from ANS. (b) O/M influence on cladding attack as a function of cladding inner surface temperature.
Reproduced from Go¨tzmann, O.; Du¨nner, Ph. In Technical Committee Meeting on Fuel and Cladding Interaction, Proceedings
of the International Working Group on Fast Reactors, IWGFR/16, Tokyo, Japan, Feb 21–25, 1977; IAEA: Austria, 1977;
pp 43–48, with permission from IAEA.

characteristics of the FCCI have been examined and
the results suggested that hypostoichiometric fuel has
advantages with respect to cladding attack. Furthermore, effects of initial O/M ratio on the maximum
depth of cladding attack with increasing burnup have
been confirmed.35

The effects of other parameters could not be clearly
evidenced. This suggests that variations of the other
parameters are not so critical. However, fuel impurities
such as C, Si, Ni, or halogens might aggravate internal
cladding corrosion either by independent interactions

or through reinforcement of FP attack as catalysts.
Therefore, the changes in characteristics and the correlation of depth of FCCI have been determined as a
function of the initial O/M ratio.
3.16.4.1.2 Effect of temperature

The changes in characteristics and the correlation of
depths of FCCI have also been determined as a
function of the cladding inner surface temperature.
The experimental data for this and the results show a
remarkable increase in the depth of cladding attack


Ceramic Fuel–Cladding Interaction

200

Depth of cladding penetration (mm)

Experiment

(210)

O/M


³1.98

£1.96

Rapsodie I
RAPS.-MON.
MFBS 6
DFR 304
DFR 350
DFR 435
MOL 7A
MOL 7B

150

451

2s (³1.98)

100
s
X

50

X

O/M ³1.98

X


O/M £1.96

0
500

600
550
650
700
Cladding inner surface temperature (ЊC)

750

Figure 4 Measured depth of cladding attack and the results of statistical analysis as a function of cladding inner
surface temperature. Reproduced from Go¨tzmann, O.; Du¨nner, Ph. In Technical Committee Meeting on Fuel and Cladding
Interaction, Proceedings of the International Working Group on Fast Reactors, IWGFR/16, Tokyo, Japan, Feb 21–25,
1977; IAEA: Austria, 1977; pp 43–48, with permission from IAEA.

with temperature above a threshold of about 500  C.
Measured depth of cladding attack in specimens of
various irradiation experiments and the results of
statistical analysis as a function of cladding inner surface temperature are shown in Figure 4.32 The depth
of cladding attack increases with higher initial O/M
ratio. The threshold temperature of the cladding
attack is higher with lower initial O/M ratio. There
is general agreement that the temperature threshold
for cladding attack is an inner surface temperature of
500  C.4,17,21,32,36–52 From out-of-pile tests, the temperature threshold of cladding attack was identified
as between 450 and 500  C, which is consistent with

in-pile tests.
Random nonuniform cladding attack has been
observed at all cladding temperatures down to
$500  C in FR fuel pins. Although high temperature
appeared to promote more widespread attack, the
depth of penetration showed no consistent variation
with temperature in the range normally employed in
FRs. Certainly, the cladding attack generally increases
with temperature only in the interval of 500–600  C;
however, saturation and a decrease of the cladding
attack were observed above 600  C.37 Figure 5 shows
the temperature dependence of depth of cladding
attack and of neutron-induced swelling in stainless
steel cladding used for the fuel pins of Phenix.53 The
maximum swelling occurred near the relative axial
position of 0.7 in the Phenix fuel pins. Generally,

swelling of FR-MOX fuel increases with burnup and
results in fuel–cladding gap closure.18,54 At further
burnup, swelling of the cladding begins to occur
depending on the swelling properties of each material,
such as length of the incubation period. A large fuel–
cladding gap forms again, and the FPs are released into
the gap in accordance with the formation of the socalled JOG (joint oxide-gaine in French).18,19,54,55 As FPs
in the gap gradually migrate to a colder region of the
fissile column, the gap conductance should be
degraded.56,57 Therefore, the maximum temperature
increase occurred across the fuel–cladding gap at near
a relative axial position of 0.7 in the Phenix fuel pin as
shown in Figure 5. This large temperature difference

across the gap would lead to a thermodynamic driving
force for cladding attack.
3.16.4.1.3 Effect of burnup

With an increase in fuel burnup, there is significant
generation and migration of FPs such as Cs, Te, I, and
Mo. In addition, the oxygen potential increases with
irradiation in the fuel pin and fuel periphery. This
determines the thermochemically stable chemical
reactions which occur between FPs and cladding
constituents. A quasilinear relation exists between
the maximum cladding depth of FCCI and burnup.
However, the influence of burnup on the depth of FP
penetration into the cladding is not clear. Table 1
summarizes the characterization of FCCI as a


Swelling DV/V (arbitrary scale)

Ceramic Fuel–Cladding Interaction

Average depth of attack (mm)

452

40

Phenix clad attack
20
Swelling

0
700

800
900
Cladding inner surface temperature (K)

1000

Figure 5 Comparison of distribution of maximum depth of cladding attack with the profile of neutron-induced
cladding swelling of fuel pin. Reproduced from Fee, D. C.; Johnson, C. E. J. Nucl. Mater. 1981, 96, 80–104, with permission
from Elsevier.

Table 1

Characterization of cladding attack as a function of O/M ratio, burnup, and cladding inner surface temperature

O/M

Burnup
Low (0–3 at.%)

High (1.98–1.99)

Moderate (1.96–1.97)

Low (1.94–1.95)




Moderate (3–6 at.%)


>675 C intergranular
<675  C matrix
<500  C no FCCI
7 pins/36 samplesb
>550  C matrix
<550  C no FCCI
27 pins/68 samplesc

>675 C evolved matrix
<675  C matrix
<500  C no FCCI
2 pins/7 samples
>550  C matrix
<550  C no FCCI
3 pins/11 samples

>700  C shallow matrix

>625  C shallow
intergranular
<625  C no FCCI
3 pins/11 samples

a

<700  C no FCCI
4 pins/16 samples


High (>6 at.%)
>500  C evolved matrix
<500  C no FCCI
6 pins/19 samples
>650  C combined
>525  C matrix
<525  C no FCCI
7 pins/34 samples
>600  C shallow intergranular
and matrix
<600  C no FCCI
4 pins/16 samples

a

Cladding inner surface temperature.
Number of fuel pins and samples examined with irradiated O/M and burnups.
c
Includes 16 pins/30 samples from HEDL P-15 test.
Source: Lawrence, L. A. Nucl. Technol. 1984, 64, 139–153.
b

function of O/M ratio, burnup, and cladding inner
surface temperature.6 On the basis of the results of
postirradiation examinations,32,46 a slight burnup
dependency on the severity of cladding attack was
obtained from the beginning until approximately
10 at.% (Figure 6).
Sections from irradiated oxide fuel pins showed

no consistent variation in maximum cladding penetration with burnup in the range up to 10.0 at.%, as
shown in Figure 7.41
Perry et al.16 and Batey and Bagley21 investigated
cladding attack over a wide range of burnups; however, a clear dependency was not obtained. Go¨tzmann
et al.40 have evaluated the in-pile data from the viewpoint that the mass transport in the reaction layer was

the rate-determining step of the cladding attack, that
is, the corrosion rate depends on the square root of
the burnup. McCarthy and Craig58 have correlated
the cladding corrosion depths with cladding inner
wall temperatures assuming that the cladding corrosion is proportional to the burnup.
3.16.4.1.4 Effect of temperature difference
between fuel and cladding

With an increase in fuel burnup, the fuel pellet is
swollen, and the distance between the fuel and the
cladding is generally decreased. However, when the
cladding diameter increase is larger at high burnup, the
distance between the fuel and the cladding becomes
wider again. Thus, thermodynamic conditions might


Ceramic Fuel–Cladding Interaction

Microns Mils

30.5 1.2

800


Temperature of cladding inside surface (ЊF)
1000
1200
1400

Measured

Predicted

25.4 1.0
Depth of penetration

453

20.3 0.8

Avg. burnup
at.%
1.0
2.1
4.6

15.2 0.6
10.2 0.4
5.1 0.2
0.0

427

(a)


649
760
538
Temperature of cladding inside surface (ЊC)

Depth of cladding penetration (mm)

DFR 350 (5.6% Bu max)
DFR 435 (7–9.7% Bu max)

150

100

50

(420)

0
450

(b)

500

550
600
650
Clad inner surface temperature (ЊC)


700

Figure 6 (a) Measured depth of cladding attack on fuel pins of type 316 (20 % CW) cladding at different local burnup and
predicted depth of cladding attack using HEDL correlation.32 Reproduced from Roake, W. E.; Hilbert, R. F.; Adamson, M. G.;
et al. In Technical Committee Meeting on Fuel and Cladding Interaction, Proceedings of the International Working Group
on Fast Reactors, IWGFR/16, Tokyo, Japan, Feb 21–25, 1977; IAEA: Austria, 1977; pp 137–158, with permission from IAEA
and (b) measured depth of cladding attack as a function of cladding inner surface temperature at different local burnup.46
Reproduced from Go¨tzmann, O.; Du¨nner, Ph. In Technical Committee Meeting on Fuel and Cladding Interaction, Proceedings
of the International Working Group on Fast Reactors, IWGFR/16, Tokyo, Japan, Feb 21–25, 1977; IAEA: Austria, 1977;
pp 43–48, with permission from IAEA.

be rearranged with a widened fuel–cladding gap
because of the notably marked influence of cladding
deformation, in addition to the axial migrations of
FPs and oxygen.
The effect of cladding strain on the fuel–cladding
gap widening has been investigated.54,55 Figure 8(a)
shows the effect of cladding strain on the depth of
cladding attack with increasing burnup.55 The oxygen potential continues to increase with irradiation at
cladding strain (DD/D) less than 1%. The cladding

attack under normal conditions would be controlled
by rather slow corrosion kinetics. On the other hand,
at cladding strain (DD/D) beyond 1% cladding strain,
cladding attack exhibits much faster kinetics because
of the abundantly supplied reactive agents.
As illustrated schematically in Figure 8(b),59 the
interface region is considered as fuel (a) and cladding
(b) separated by a gas-filled or reaction product-filled

gap which supports the bulk of the temperature
difference. When the swelling of cladding is high


Ceramic Fuel–Cladding Interaction

Cladding corrosion (mm)

454

PNC (high density)
PNC (low density)
Coquerelle et al. 42
Weber et al.43

100

50

10

20

30

40

50

60


70

80

90

100

Burnup (MWD kg-1 M)
Figure 7 Depth of cladding attack as a function of local burnup. Reproduced from Koizumi, M.; Nagai, S.; Furuya, H.;
Muto, T. In Technical Committee Meeting on Fuel and Cladding Interaction, Proceedings of the International Working Group
on Fast Reactors, IWGFR/16, Tokyo, Japan, Feb 21–25, 1977; IAEA: Austria, 1977; pp 64–73, with permission from IAEA.

Fuel

Gap

Cladding

Fuel

Gap

Cladding

A

~30 mm


B

T (K)
1100

180
Corrosion depth (mm)

900
Fuel pins with
strained cladding
DD/D ³ 1%

140

100

Homogeneous fuel pins
Heterogeneous fuel pins

60

Fuel pins with
unstrained cladding
DD/D < 1%

20

(a)


0

2

4

8
6
Average burnup (at.%)

10

12

14

(b)

Figure 8 (a) Effect of cladding strain on depth of cladding attack with increasing burnup. Reproduced from Bailly, H.;
Menessier, D.; Prunier, C. The Nuclear Fuel of Pressurized Water Reactors and Fast Neutron Reactors: Design and Behavior;
Intercept Ltd.: UK, 1999; p 475, ISBN: CEA 2-7272-0198-2, with permission from CEA and (b) schematic illustration of
temperature gradient in fuel–cladding gap interface. Reproduced from Adamson, M. G.; Aitken, E. A.; Lindemer, T. B. J. Nucl.
Mater. 1985, 130, 375–392, with permission from Elsevier.

and the fuel–cladding temperature gap is large, cladding attack might be more enhanced by different
thermochemical conditions. Thus, the temperature
gradient (or lack thereof) between fuel and cladding
is a key parameter determining the driving force and
occurrence of cladding attack. The majority of irradiated fuel pins retain a significant fuel–cladding gap
over the greater part of the fuel column. But with


increasing burnup, this gap is either partly or fully
filled with reaction products.18,55–61 The reaction products in the gap are also seen to have extruded down
into radial cracks at the fuel periphery, and in some
cases islands of a phase having bright metallic ingots
are contained in the corrosion product. The thermal,
mechanical, and physicochemical properties of the
oxide fuels vary in a continuous manner owing to


Ceramic Fuel–Cladding Interaction

the presence of the reaction products, which thus
directly influence fuel temperatures, cladding attack,
and fuel–cladding interaction. The physical and
chemical states of the different FP compounds determine the volume occupied by these species and therefore swelling, that is, the volume increase of fuel pellets
resulting from fission.
3.16.4.1.5 Effects of cladding materials

Corrosion resistance of three major types of stainless
steel claddings, M316 (austenite), FV548 (ferrite), and
Nimonic PE16 (high Ni), have been investigated in an
irradiation experiment.21 The incidence and severity of
cladding attack were almost identical or indistinguishable in oxide fuel pins irradiated under similar conditions irrespective of cladding type. Furthermore,
corrosion of cladding materials in the fully solution
treated or 20% cold-worked conditions were investigated for M316 and M316L stainless steel cladding
pins. The influence of cladding heat treatment on corrosion sensitivity was not detected (see Chapter 2.08,
Nickel Alloys: Properties and Characteristics;
Chapter 2.09, Properties of Austenitic Steels for
Nuclear Reactor Applications; Chapter 4.02, Radiation Damage in Austenitic Steels; Chapter 4.03,

Ferritic Steels and Advanced Ferritic–Martensitic
Steels; Chapter 4.08, Oxide Dispersion Strengthened Steels; and Chapter 4.04, Radiation Effects in
Nickel-Based Alloys).
The corrosion susceptibility of M316 stainless
steel (carbon content $0.05%) and M316L stainless
steel (carbon content 0.02%) was also investigated.21
Although carbide precipitation and associated chromium depletion at grain boundaries should facilitate
intergranular attack, it was indicated that the initial
carbon content was not dominant among the factors
determining corrosion resistance. Although carburizing or nitriding of the cladding by transfer of carbon
and nitrogen from the fuel occurred, neither promotion nor prevention of FCCI was shown.21
On the contrary, out-of-pile results showed that
carbon may influence the nature of FP-induced cladding attack by saturating alloy grains near the
cladding surface in the form of intergranular carbide
precipitates.62 Heavily carburized regions of the cladding material were chemically attacked by Cs–Te
mixtures; the result was an attack zone with a more
uniform appearance than the deep intergranular type
observed with alloys of large grain size.
The effects of nitriding of cladding materials on
the attack were identified in an irradiation experiment.8 The cladding was irregularly nitrided on the

455

inner surface to a maximum depth of 4 mil (102 mm)
because of fuel pellets in the top half of the fuel
column which each contained 4800 ppm of nitrogen.
Cracks in the cladding were occasionally observed in
or near the nitrided layer, and they were caused by
the differential expansion of this layer. However, precipitates were observed along the cracks between
the nitrided layer and the remainder of the cladding,

which were suggested to be chromium nitrides.
Although nitriding of the cladding materials would
appear to complicate the situation of cladding attack,
the corrosion resistance of steels should be decreased
by the precipitation of chromium.63
3.16.4.2

FCCI Model and Wastage Equation

Several mechanism models have been proposed to
explain FCCI in MOX fuels based either on oxidation
or materials transport processes. However, the fuel pin
design only varies with respect to empirical equations
rather than fundamental models because explanations
of the observed effects of FCCI are speculative. Therefore, FCCI wastage equations for design have been
based on collected in-pile test data. Wastage correlations for fuel pin design were developed by relating the
loss of effective cladding thickness to irradiation and
fabrication parameters of the fuel pins. The following
wastage correlations, expressed using penetration
depth, have been proposed.
The ANL (Argonne National Laboratory) correlation, which was made by fitting the HEDL (Hanford
Engineering Development Laboratory) data, was
developed in order to evaluate in detail the relative
importance of the temperature difference across the
fuel–cladding gap. The average depth of cladding
attack was obtained by the following correlation:
Ave:WDepth ðANL; mmÞ ¼ 3 Â 105
À
Á
B½4:4 Â 104 exp À31200=TðfuelÞ

À
Á
À 2:7  103 exp À24600=TðcladdingÞ Š

½1Š

where B ¼ local burnup (at.%), T(fuel) ¼ temperature
of fuel outer surface (K), and T(cladding) ¼ temperature
of cladding inner surface (K).
Another wastage correlation developed at ANL64 is
shown in eqn [2] and is based on EBR-II (experimental
breeder reactor-II) and FFTF (Fast Flux Test Facility)
data. The EBR-II data included 305 data sets from
104 fuel pins with Type 316 stainless steel cladding
and 24 data sets from ten fuel pins with D9 stainless
steel cladding, irradiated to approximately 17 at.%
burnup at T values up to 730  C. FFTF data included


456

Ceramic Fuel–Cladding Interaction

78 data sets from a mixed total of 24 pins irradiated
to about 15 at.% burnup at T values up to 685  C.

cases that high initial O/M ≧ 1.98 and high burnup ≧
5 at.% are involved:

WDepth ðANL; mmÞ ¼ 0:5070 Á ðO=M À 1:935Þ


Max:WDepth ðGEðCCCTÞ; mmÞ ¼ 4:14 Â 105 B 0:5

½2Š

Á B Á ðT À 705Þ

Here, O/M > 1.935, B > 0, and T > 705; B ¼ local
burnup (at.%) and T ¼cladding inner surface temperature (K). The standard deviation of the correlation
was 12.5 mm.
A third correlation was obtained from an
experimental relationship for a low-burnup fuel pin
(1.2 at.%), designated P-23A EBR-II. The average
depth of cladding attack yielded eqn [3]65:
Ave:WDepth ðHEDL; mmÞ ¼ 3:36 Â 105 expð10063=T Þ

½3Š

where T ¼ cladding inner surface temperature (K).
Nonlinear regression analysis of the data from all
three fuel pins of two interim examinations of P-23A
HEDL fuel pins at 2.4 and 5.0 at.% yielded eqn [4]66:
Ave:WDepth ðHEDL; mmÞ ¼ 2:43 Â 106 B 0:517
expðÀ9806=T Þ

½4Š

½5Š

for B ! 0, O/M > 1.942 and T > 728; for conditions

outside these ranges, depth ¼ 0. Here, B ¼ local
burnup (at.%), C ¼ 12.23, O/M ¼ initial O/M ratio,
and T ¼ local time-averaged temperature of cladding
inner surface (K).
One GE (General Electric Company) correlation
of the cladding attack based on 68 irradiated fuel pin
sections was developed and included in fabrication
and operating parameters.4,66,67 This correlation is
given as eqn [6]:
Max:WDepth ðGE; mmÞ ¼ 2:36 Â 108
t ðO=M þ 0:001B À 1:96ÞQ
expðÀ14:185=1:987T Þ

2:33

½7Š

where B ¼ local burnup (at.%) and T ¼ cladding
inner surface temperature (K). In this equation, the
first term represents the conventional (oxidative)
FCCI and the second term represents the contribution from CCCT.
The SNR (Schneller Natriumgeku¨hlter Reactor) correlation52 which was based on DFR (Dounreay FR)
and Rapsodie (French experimental FR) data is
shown in eqn [8]. The correlation fits the average
corrosion depth and emphasizes the influence of the
initial O/M ratio on the corrosion depth (i.e., a strong
increase for O/M between 1.96 and 1.98)
Max:WDepth ðSNR; mmÞ ¼ 96:97½1 À 3:013 Â 105
ð2 À O=MÞ4 Š  exp½À76:920=ðT À 769ފ


where B ¼ local burnup (at.%) and T ¼cladding
inner surface temperature (K).
A later wastage correlation67 based on EBR-II data
is shown in eqn [5]:
Max:WDepth ðHEDL; mmÞ
¼ 0:4521ðB þ CÞ Â ðO=M À 1:942ÞðT À 728Þ

ÂexpðÀ9776=T Þ þ 0:86ðT À 873Þ0:5 ðB À 4Þ

½8Š

for 1.96 O/M 2.00 and 769 T 923 K; T ¼
cladding inner surface temperature (K) and O/M ¼
initial O/M ratio.
The PNC (Power Reactor and Nuclear Fuel
Development Corporation) correlation which suggests the maximum cladding corrosion depths for
the fuel pin design is shown in eqn [9]41:
Max:WDepth ðPNC; mmÞ ¼ 4:64 Â 10À4
Á ðO=M À 1:93Þ Á ðT À 360Þ Á B

½9Š

where T ¼ temperature( C), O/M ¼ initial O/M
ratio, and B ¼ burnup (MWd tÀ1); but T ¼ 620
(if cladding temperature ! 620), B ¼ 35 000 (if
burnup !35 000), and O/M ¼ 1.98 (if initial O/M
ratio !1.98).
A wastage correlation developed at PFR (prototype FR),68 which was based on data (up to 10 at.%
burnup), is shown in eqn [10].
Max:WDepth ðPFR; mmÞ


ð1 À pÞ
½6Š

where T ¼cladding inside surface temperature (K),
t ¼ equivalent full power days of irradiation, O/M ¼
initial O/M ratio, B ¼ local burnup (at.%), p ¼ smear
density (a fraction of the theoretical density), and
Q ¼ peak linear heating rate (W mmÀ1).
Another GE model involves CCCT as shown in
eqn [7]. This GE (CCCT) model applies only to the

¼ 10 Â Maximum burnup ðat:%Þ

½10Š

Relatively conservative wastage correlations [2], [5],
and [9] that employed major parameters such as cladding temperature, burnup, and O/M ratio are like the
envelopes of measured data points. Those parameters
are conservatively considered relating to threshold
cladding temperature, incubation period (burnup),
and threshold oxygen potential (O/M ratio). The
wastage correlation (10) was developed on the basis


Ceramic Fuel–Cladding Interaction

of experimental data, but it conservatively includes a
contribution from fabrication tolerances and external
corrosion (as well as FCCI).

Wastage correlations [1], [3], [4], and [6–8]
obtained by fitting experimental data with major parameters such as cladding temperature, burnup, and
O/M ratio were developed on the basis of the conventional (oxidative) FCCI. The form of the correlation
equations were mainly expressed as an exponential
temperature relationship. Thermally activated processes assumed in cladding attack exhibit exponential
temperature dependence. In addition, the wastage correlation [7] involves the term of CCCT contribution
expressed by threshold values (cladding temperature
and burnup). The wastage correlation [6] includes a
parameter of the smear density that has a large impact
on the fuel behavior, because of its large influence on
the thermal conductivity of the fuel. Although much
data is collected for FCCI, the range of parameters and
the uncertainty of the cladding temperature constitute
serious obstacles to developing consistent wastage
correlations. The form of correlations, except the
correlation [7], was not chosen to reflect current
knowledge of probable FCCI mechanism. The correlations have not successfully predicted and consisted
of actual corrosion behavior. Cladding wastage equations are often used to predict the overall condition of
cladding attack in fuel pins at each burnup level. The
current model used in cladding wastage equations
cannot sufficiently account for localized deep attack.
Cladding attack depends on various corrosion
mechanisms, reaction potentials, FP transport (release,
radial and axial migration), and condensation within
the fuel–cladding gap, which is always changing during
the course of an irradiation; thus severity of cladding
attack is not directly dependent on the above major
parameters in the correlations. The kinetics of FCCI
in fuel pins during irradiation is determined through
consecutive processes: (a) generation and release of

corrosive FPs and supply of free oxygen, (b) transport
of FPs and oxygen to the cladding, and (c) reaction
between the elements and cladding constituent elements. The process (c) is strongly dependent on cladding temperature. The process (b) is related to not
only fuel parameters and irradiation parameters but
also nonuniform temperature within the fuel–cladding
gap. In addition to change in cladding temperature
during irradiation, the nonuniform temperature associated with fuel behavior produces a change in the
character of FCCI. The amount of cladding wastage
will be different at each burnup level because of relative shifts in the processes.

457

3.16.5 Mechanism of Oxide Fuel and
Cladding Interaction
3.16.5.1

Oxygen Potential of Irradiated Fuel

The occurrence of FCCI depends on the excess oxygen available for interaction and on the cladding inner
surface temperature. The available oxygen is related
to the initial fuel O/M ratio and burnup. Thus, the
increase of the oxygen potential ðDG O2 ¼ RT lnpO2 Þ
due to burnup enhances the rate of the cladding
attack.50,53 The chemical state of the FPs depends on
the oxygen potential of the fuel, a potential which in
turn varies during irradiation, depending on the
amount of oxygen consumed by the FPs and therefore
on the nature of the compounds that they form. FCCI
is thermochemically controlled mainly by the oxygen
potential (see Chapter 2.02, Thermodynamic and

Thermophysical Properties of the Actinide Oxides; and Chapter 2.21, Fuel Performance of Fast
Spectrum Oxide Fuel).
In comparison with the fission of uranium, plutonium fission makes more oxygen available because of
higher yields of noble metals and lower yields of
oxide formers.39,49 The different fission yield curves
for plutonium and uranium result in greater increases
in oxygen potential with burnup for plutoniumcontaining fuels.69 Soon after reactor startup, before
significant amounts of FPs have been generated, the
oxygen in the fuel redistributes rapidly to make the
peripheral fuel nearly stoichiometric. Different techniques for direct measurement of the radial oxygen
profiles have been employed in postirradiation examinations.67–76 Figure 9(a) compares the oxygen profile
of irradiated samples obtained by an EMF method72
with similar profiles, which were obtained by using
different techniques (EPMA analysis of Mo and
MoO2 content in metallic and ceramic inclusions74
and X-ray measurements of lattice parameters75). In
the figure, three different zones may be postulated to
exist.72 Zone a is near the central hole, in which rapid
depletion of oxygen occurs; this zone coincides
approximately with the columnar grain region. Zone
b is an intermediate flat zone (or plateau). Zone c is
near the periphery of the oxide pellet in which the
O/M approaches the stoichiometric value.
A slight amount of carbon in MOX fuel is included
as an impurity during the fabrication process. It was
postulated that the oxygen potential at any point in a
fuel pin during use is controlled by a local equilibrium
with a CO/CO2 mixture of constant composition,77
from which it follows that redistribution of oxygen
should occur to give a nearly stoichiometric fuel at



458

Ceramic Fuel–Cladding Interaction

Position, r/r0 (–)
00.2 0.4

0.6

0.8

1.0

2.00
in zon

emf.

1.90

Pu/(U + Pu)
Initial O/M
BU, at.%
Final (O/M)

0.2

0.2

1.94
3.4
1.99

0.4

1.976
Low
1.976

0.6

Position (r/r0)2 (–)

0.3
1.98
7
1.994

0.8

1.0

1000

900

1100

1200


-700

-160
Unir

radia

-600

ted,

O/M

-140

= 1.9

82

-500

Burnup, at.%
Edge

Cente

r

( )

( )

-400

} 3.8
} 7.0

-100

11.2

-80

-300
900

(b)

-120

1000

1100

1200

1300

1400


2

Conte Kleykamp
X-ray Mo/MoO2

800

Oxygen potential, DGo (kcal mol-1)

2

1.94

1.92

(a)

Temperature (ЊC)
700

Oxygen potential, DGo (kJ mol-1)

1.96

nar gra

1.98

Colum


Oxygen to metal ratio (–)

e

Central
void

Temperature (K)

Figure 9 (a) Comparison of present profile obtained with EMF measurements with experimental results of Kleykamp and
Conte et al. using the Mo/MoO2 method and X-ray techniques, respectively. Reproduced from Ewart, F.; Lassmann, K.;
Matzke, Hj.; Manes, L.; Saunders, A. J. Nucl. Mater. 1984, 124, 44–55, with permission from Elsevier. (b) Oxygen potential
measurements with the EMF cell on irradiated Phenix fuel of initial composition (U0.8Pu0.2)O1.98. Reproduced from Matzke,
Hj.; Ottaviani, J.; Pellottiero, D.; Rouault, J. J. Nucl. Mater. 1988, 160, 142–146, with permission from Elsevier.

the fuel periphery. Therefore, the oxygen potential
of the fuel periphery in contact with the cladding has
the same stoichiometry in irradiated oxide fuels,
despite the different initial O/M ratios, after an equilibration period of oxygen redistribution. As a consequence, the influence of the initial stoichiometry on
cladding attack is not strong. From the data of Rand
and Markin,77 this stoichiometry at the fuel periphery
corresponds to an oxygen potential DG O2 in the fuel–
cladding gap of approximately À418 kJ molÀ1 at
700  C, which is more than sufficient for growth of a
Cr2O3-type protective oxide on the inner cladding
surface (the oxygen potential DG O2 in equilibrium
with chromium and Cr2O3 at 700  C is À576 kJ molÀ1).
The oxygen potential of irradiated Phenix fuel with an
initial composition Pu0.2U0.8O1.982 was measured78 and
the results are shown in Figure 9(b). The DG O2 value

increased with increasing burnup from 3.8 to 11.2 at.%,
and it also increased linearly with temperature. As for
oxygen stoichiometry of these irradiated fuels, the
DG O2 data measured for the pellet periphery
was regarded as giving the oxygen potential of the near-stoichiometric composition.72,78
In addition, for MOX fuel containing small
amounts of americium, the oxygen potential increases

with increasing americium content.79 In the Superfact
experiment, four oxide targets containing high and low
concentrations of 237Np and 241Am were irradiated in
the Phenix reactor.80,81 The maximum depth of corrosion was 50 mm in the SF13 (U0.74Pu0.24Np0.02O2 À x)
and SF16 (U0.74Pu0.24Am0.02O2 À x) fuel pins.
The depth of the cladding attack did not exceed the
limit set for standard MOX fuel pins with the same
burnup.
3.16.5.2 Characteristics of Major
Corrosive FPs
The chemical state of the corrosive FPs depends on
the oxygen potential of the fuel. Oxygen potential
varies during irradiation, depending on the amount of
oxygen consumed by the FPs and on the nature of the
compounds that they form.
Cesium, tellurium, and iodine are often called
volatile FPs. These elements and part of the compounds they form are gases at the fuel temperatures.
Therefore, they will undergo considerable radial and
axial migration to colder temperature regions. They
generally accumulate in the gap; in some cases, this
results in contact between fuel and cladding.



Ceramic Fuel–Cladding Interaction

The cesium is obtained in high yield. The large
excess of cesium over I and Te guarantees that not
only almost all the iodine and tellurium are bound as
CsI and Cs2Te, but also that there is still a large
amount of cesium left in the fuel pins. In FRs, the
Cs/I and Cs/Te ratios are approximately 10 and 4,
respectively.82 The other reactive FPs have very low
yields, and hence it is considered that cesium/oxygen
reactions are dominant in the pins.
As shown in many postirradiation investigations,
FP-enhanced oxidation is the main cause of cladding
attack in oxide fuel pins. Various out-of-pile experiments have shown that the FPs (I, Te, Sb, Cd, In, and
Sn) directly attack stainless steel.83 Understanding
the dependence of those reactions on the cladding
temperature and the oxygen potential in the system is
important to obtain a better understanding of the
reaction possibilities of the various FPs.
3.16.5.2.1 Iodine

Investigations with iodine revealed that reactions
with stainless steels take place at 400  C. Iodine reacts
predominantly with Cr. Preferential reactions along
the grain boundaries of the cladding material yield a
type of reaction similar to that of pitting corrosion.83
However, iodine is not present in elemental form, but
as CsI molecules which are thermodynamically stable
under the conditions in the fuel pins. The equilibrium pressures of I2 and I in the fuels for various

oxygen potentials are very low and much lower than
the CsI pressure (5Â10À7 atm at 727  C).84
In addition, dissociation of CsI in the strong
G-radiation field of the fuels has been postulated.
Cubicciotti and Davies85 have shown that free iodine
was released from thermochemically stable solid
iodides by G-radiation. The chemical activity of
iodine released into the fuel–cladding gap was associated with the amount of cesium which probably
could control the activity.
Iodine in the elementary state causes severe attack
on stainless steel cladding. However, when bonded
to Cs, it is not corrosive to the cladding materials.
Nevertheless, the state of iodine in the fuel has been
questioned for a long time in view of the dissimilar
transport of cesium and iodine in fuels,13 and their
different release rates.86 Investigations of 129I and
137
Cs radial profiles in light water reactor (LWR)
pin sections indicated that iodine migrates slightly
faster than Cs. This would contradict the possibility
for excessive CsI formation in the fuels. However,
experimental evidence from release characteristics
and postirradiation investigations87 has provided

459

support for the presence of CsI, rather than elemental
iodine, in the fuel pin. As other iodides, for instance, of
Zr, Mo, or Ru are less stable,88 it can be assumed that
CsI is the dominant iodide species in the fuel pin, and

that it will escape as such when lack of cooling causes
the fuel temperature to rise. Finally, as the atomic
ratio Cs/I is approximately 10, it can be expected
that all, or almost all iodine is present as CsI.
Hofmann and Go¨tzmann83 performed out-of-pile
reaction tests of cladding material with UO2 and UO2.08
in the presence of CsI or iodine. High-purity CsI did
not react with the cladding steels even at 800  C for
1000 h. Nevertheless, reactions took place with the
cladding when the oxygen potential was above a critical
value. The maximum depth attained by the reaction
with stoichiometric UO2 was about 10 mm at some
points. No reaction between CsI (simulated burnup
20 at.%) and stoichiometric UO2 at 800  C for 1000 h
was detected; however, marked reactions took place
with UO2.08 (reaction depth, about 20 mm). On the
other hand, reactions with UO2.08 were much reduced
in the absence of CsI. The addition to UO2.08 of free
iodine caused reactions up to 50 mm in depth in the
cladding material after 1000 h at 800  C. This indicated
that the oxidation of the cladding by hyperstoichiometric fuel was considerably accelerated in the presence of CsI and I.
3.16.5.2.2 Cesium

A large excess of cesium over I and Te is generated by
fission. Although almost all iodine and tellurium are
bound as CsI and Cs2Te, there is still a large amount
of cesium left in the fuel pin. Elemental cesium has a
high volatility, but formation of cesium uranate or
cesium molybdate lowers and determines the vapor
pressure of cesium.

Hofmann and Go¨tzmann83 conducted out-of-pile
tests of cladding material with UO2 and UO2.08 in
the presence of Cs. Elemental cesium was compatible
with stainless steels up to 3000 h at 1000  C. Even minor
impurities of cesium with oxygen caused a reaction with
the cladding. The depth of cesium reactions showed a
dependence on the O/M ratio of the fuel. The cladding
attack by hyperstoichiometric fuel was considerably
accelerated in the presence of cesium. While after
1000 h, reaction zones of less than 5 mm were observed
in contact with UO2.08 at 800  C, 100 mm was attained
after the addition of Cs (simulated burnup 10 at.%). If
molybdenum was added to the mixture of UO2.08 þ Cs,
the reactions with the cladding became weaker.
The Cs compound Cs2CrO4 reacted very violently with stainless steels at 800  C and 1000 h and


460

Ceramic Fuel–Cladding Interaction

grain boundary reactions took place, penetrating
more than 1000 mm. Cs2Cr2O7 also reacted with the
cladding. The chemical interactions covered some
50 mm at 800  C and 1000 h.83
Formation of various cesium compounds is determined by the oxygen potential. This indicates that
reactions between cesium compounds and cladding
occur only above a certain oxygen potential. The
change in chemical constitution within the gap with
variation of oxygen potential of a fuel (U0.7Pu0.3O2 Æ x)

in the presence of FPs (simulated burnups 2, 10,
and 20 at.%) has been examined by thermochemical
calculations of thermodynamic equilibria.89 From
thermodynamic studies, formation of the major cesium
compounds was determined at certain oxygen potentials of a fuel. The reactions which can buffer the
oxygen potential within the gap are given in Table 2.
Formation of cesium uranate and molybdate
was thermodynamically suggested.83 Postirradiation
examinations of highly irradiated fuel showed
Cs2MoO4 to be present on the fuel side of the fuel–
cladding gap. By contrast, the formation of Cs2U4O12
instead of Cs2UO4 or Cs2UO3.5 was observed in irradiated fuel.13 The formation of this uranate will cause
considerable swelling of the fuel90,91 and decrease the
fuel–cladding gap width.
3.16.5.2.3 Tellurium

Tellurium is a precursor of iodine. TeO2 cannot
be formed because the oxygen potential of the
Table 2
No.

Te/TeO2 system is much higher than that in
the fuel. Formation energy of stable Cs2Te appears
to be approximately 376 kJ molÀ1, which is just
enough for Cs2Te to be stable at the conditions in
the fuel. Cs2Te showed an appreciable volatility
at the conditions in a fuel.92 The vapor pressure
of Cs2Te is about 10–4 atm at 727  C, which means
that an appreciable amount of tellurium is already
present as Cs2Te in the gap at normal operating

conditions of the fuel pins. If gas-phase transport
is the main path for FPs to reach the fuel surface, as
has been stated,93 it is important to know the vapor
pressure of Cs2Te.
Pure tellurium showed marked penetration in
stainless steel up to depths comparable with those
observed in irradiated fuel pins at temperatures
below 800  C. Furthermore, out-of-pile experiments
showed that the reaction of tellurium does not
depend on the oxygen potential unlike in the case
of cesium.94 The reactions with tellurium are clearly
functions of irradiation period and temperature.
Go¨tzmann and Hofmann95 determined the maximum penetration depths of tellurium in stainless
steels as a function of the temperature and annealing
period. At 500 and 700  C, attack occurred along the
grain boundaries of the cladding. Below 500  C a
uniform attack was observed, and tellurium migrated
into the cladding by interfacial reactions. Above
700  C, the penetration depths of tellurium were
dependent on the grain size of the cladding material.

The reactions which can buffer oxygen potential within the fuel–cladding gap of fast reactor fuel pins

ÀGO2 ðkJ molÀ1 O2 Þ

Buffering reaction

Other condensed
phases present


900 K

1200 K

1

654

586

4Cs þ 2½U; PuŠO2Àx þ

2
3

602
595

535
544

4Cs þ Cr þ 2O2 ⇄Cs4 CrO4
2Cr þ 32O2 ⇄Cr2 O3

À3
2

Á
þ x O2 ⇄2Cs2 ½U; PuŠO3:56


CsI, Cs2, Te, Mo, Cr
CsI, Cs2Te, Mo
CsI, Cs2[U, Pu]O3,56
Cs2Te, Mo
CsI, Cs2Te Cs4CrO4

4

590

530

2Cs þ Mo þ 2O2 ⇄Cs2 MoO4

5

583

532

CsI, Cs2Te

6

549

494

4Mo þ 2Cs4 CrO4 þ 17
2 MoO4 þ Cr2 O3

2 O2 ⇄4Cs
À
Á
2Mo þ 2Cs2 ½U; PuŠO3:56 þ 52 þ x O2 ⇄2Cs2 MoO4 þ 2½U; PuŠO2Àx

7

525

351

2Cs2 ½U; PuŠO3:56 þ 12O2 ⇄2Cs2 ½U; PuŠO4

8
9

431
424

376
372

Mo þ Cs2 Te þ 2O2 ⇄2Cs2 MoO4 þ Te
Mo þ O2 ⇄MoO2

10

312

266


Cs2 Te þ ½U; PuŠO2Àx þ ð1 þ x=2ÞO2 ⇄Cs2 ½U; PuŠ4 O12 þ Te

11

287

260

12

219

180

Cs2 MoO4 þ MoO2 þ 12O2 ⇄Cs2 Mo2 O7
À
Á
2Cs2 MoO4 þ 4½U; PuŠO2Àx þ 12 72 À x O2 ⇄Cs2 Mo2 O7 þ Cs2 ½U; PuŠ4 O12

CsI, Cs2MoO4, Cs2Te,
Cr2O3
CsI, Cr2O3
CsI, Cs2MoO4, Te,
Cr2O3
CsI, Cs2MoO4, Cs2Te,
Cr2O3
CsI, Te, Cr2O3

CsI, Cs2Te, Cr2O3


CsI, Te Cr2O3

Source: Ball, R. G. J.; Burns, W. G.; Henshaw, J.; Mignanelli, M. A.; Potter, P. E. J. Nucl. Mater. 1989, 167, 191–204.


Ceramic Fuel–Cladding Interaction

The maximum penetration depth of tellurium was
greater in coarse grains than in fine grains.
When the oxygen potential DG O2 is approximately
376 kJ molÀ1, Cs2Te is stable in the fuel. When the
oxygen potential reaches a sufficiently high level
(>À418 kJ molÀ1), Cs2Te becomes less stable than
cesium molybdate or urinate.96 Cs2Te can attack
the cladding by reaction with the cladding constituents, forming cesium chromate, iron telluride, and
nickel telluride. Because of the dissolution of the
cladding materials into Te, the cladding constituents
diffuse into the fuel even at temperatures as low as
400  C. In contrast, tellurium diffuses into the cladding material and is found at the reaction front
together with Cr.96
However, the highly localized concentration of tellurium required for cladding attack in FR fuel pins
appears to be inconsistent with the postirradiation
investigations of numerous cladding attack regions in
which tellurium has not been detected.37,39,40,46 The
role of tellurium in cladding attack is largely based on
results from out-of-pile tests38,40,61,94,97–103 in which
the average tellurium content per unit area of cladding
inner surface was much greater than that encountered
in irradiation tests.

3.16.5.3 Various Corrosion Reaction
Mechanisms
3.16.5.3.1 Corrosion early in life

Generally, postirradiation examinations of specimens
irradiated at low power ($20 kW mÀ1) and low
burnup (2 at.%) have shown the absence of fuel–
cladding chemical attack.47,104 At low burnup, a radial
redistribution of oxygen occurs early in life. Oxygen
migration down the temperature gradient provides
oxygen activity at the cladding surface sufficient to

(a)

461

oxidize the Cr in the cladding material. Significant
cladding attack is not expected until FPs (Cs, Te,
and I) are available to participate in an enhanced
corrosion process. The quantities required are small,
however, and sufficient quantities are generated
by approximately 0.4 at.% burnup to initiate the
reactions. Therefore, it is expected that even at burnups as low as 1 at.%, significant reaction can be
observed.55 An internal cladding corrosion phenomenon due to Te and I has been observed,7 which is linked
to the oxide thermal behavior at the very beginning of
life. This intergranular type of corrosion is very deep
and it is characteristic of pins having been operated at
high linear power at the very beginning of irradiation.
The fission yield for cesium is generally higher
than is needed to react with Te and I. Potentials

of tellurium and iodine are greatly reduced unless
cesium is consumed by other reactions.
Figure 10 shows that tellurium-induced grain
boundary cladding attack in short-term irradiation
tests (burnup 0.1 at.%) of annular fuel occurred in a
low density fuel pellet of 86% T.D. (theoretical density) but not in a higher density fuel.7 The maximum
penetration depth of this type of cladding attack was
60 mm after 11 days of irradiation. Similar localized
intergranular cladding attacks are likely to appear at
the very beginning of life (from loading to an irradiation of 10 days) and show very deep penetrations. This
very beginning of life corrosion, termed ‘corrosion-dejeunesse’, in French, was in the form of spots of limited
area in fuel pins irradiated at high linear heating rate.55
This cladding attack is caused by corrosive FPs
(I and Te) accumulating in contact with the cladding.
Iodine and tellurium formed by fission generate Cs
and Rb through radioactive decay after roughly
10 days. Tellurium could rapidly form compounds
such as Cs2Te. However, because of the time required

(b)

Figure 10 Tellurium-induced early-in-life cladding attack (burnup 0.1 at.%): (a) chromium oxide scale on cladding wall
and (b) intergranular attack with grain boundaries ($60 mm). Reproduced from Go¨tzmann, O. In Fast Reactor Core and Fuel
Structural Behaviour, Proceedings of the International Conference, Inverness, June 4–6, 1990; BNES: London, UK, 1990;
pp 1–8, with permission from BNES.


462

Ceramic Fuel–Cladding Interaction


for the formation of rubidium and cesium, the iodine
and tellurium elements are in excess at the very
beginning of irradiation. Under the influence of
high linear heat rating, these FPs migrate radially
and axially to the colder temperature regions. If, in
these regions, the oxygen potential and the cladding
temperature are high enough, significant intergranular attack of cladding materials develops.
3.16.5.3.2 Iodine transport of cladding
constituents

Free iodine reacts with cladding materials. But
iodine is easily bound to Cs during irradiation, so
only insignificant quantities of I are available to react
with cladding materials. A reaction of the cladding
materials with I needs a sufficiently high partial pressure of I which is not possible from a thermodynamical
viewpoint. However, metallic inclusions of pure iron
from the cladding by iodine transport were observed
within the irradiated fuel in addition to intergranular
and matrix attack of cladding.16,105–107 Figure 11(a)
shows the appearance of metallic precipitates along
fuel cracks observed in irradiated MOX fuel.105
The radiation effect on the partial pressure of I,
involving the effect of recombination, has been

evaluated for gaseous CsI based on kinetic theory.108
The calculations were carried out for an oxygen
potential DG O2 ¼ À418kJmolÀ1 at the cladding
inner surface temperature of 600  C. The partial
pressure of iodine under the radiation condition was

approximately 10À2 Pa ($10À7 atm), which was
higher than that under the nonradiation condition
approximately 10À9 Pa ($10À14 atm). Radiation
effect on the partial pressures of FeI2, CrI2, and
NiI2 as a function of cladding temperature is shown
in Figure 11(b).108 These calculations indicated that
vapor transport of iron and chromium was possible
within a FR fuel pin. And the pressure of NiI2 (under
the radiation condition) was so low that significant
transport may not occur.
Aubert et al.107 proposed a vapor transport mechanism by iodine which is similar to the Van Arkel–de
Boer process. According to the vapor transport
mechanism, iodine reacts with stainless steel at the
cold region of the fuel–cladding gap to form metal
diiodides and they diffuse through the gap into the
fuel. The radiation-induced decomposition of metal
diiodides occurs and metal iodides are formed at the
fuel surface which corresponds to the hot regions.
This leads to transportation and precipitation of
0
PFel2 (rad)
-5
PCrl2 (rad)

100
200

log P (Pa)

-10


-15

PNil2 (rad)
PFeI (non-rad)
2

300
100x

800 mm

500

-20

PCrI2 (non-rad)

600
700

(a)

-25

(b)

PNil (non-rad)
2
800

900
1000
Cladding temperature (K)

Figure 11 (a) Appearance of ‘metallic rivers’ observed in MOX fuel. The rivers apparently originated at the fuel–cladding
interface. Reproduced from Fitts, R. B.; Long, E. L., Jr.; Leitnaker, J. M. In Fast Reactor Fuel Element Technology,
Proceedings of Conference, New Orleans, LA, Apr 13–15, 1971; Farmakes, R., Ed.; American Nuclear Society: Hinsdale, IL,
1971; pp 431–458, with permission from ANS. (b) Radiation effect on the partial pressures of Fel2, Crl2, and Nil2 as a function
of cladding temperature. Reproduced from Konashi, K.; Yano, T.; Kaneko, H. J. Nucl. Mater. 1983, 116, 86–93, with
permission from Elsevier.


Ceramic Fuel–Cladding Interaction

elements in the stainless steel and the release of
iodine available for further reaction with the steel.
According to the Van Arkel–de Boer process,
iodine reacts with cladding materials by the following
reversible reaction:
MðsÞ þ I2 ðgÞ Ð MI2 ðgÞ
where M ¼ Fe, Cr, or Ni and MI2 is the metal iodide
of M. Metal iodides decompose in the region of high
temperature. The free I2 molecules then diffuse back
to the cladding to react with it. As these reversible
reactions on both sides of the fuel and cladding are
cyclically continued, the cladding constituents are
transported to the fuel surface and inside cracks by
small quantities of iodine. But occurrence of this
transport is possible in only relatively high oxygen
potentials that are required to form a stable cesium

uranate.
In the high radiation field of FR fuel pins during
use, radiation decomposition of CsI would release
iodine for participation in the iron transport process.85,104 Some other reaction must remove cesium
from the gas phase in the fuel–cladding interface for
iodine to be freed. That only iron inclusions are
found within the fuel is reasonably understood in
consideration of the stability and volatility of possible
metal iodides. The free energy of formation of NiI2 is
too high for it to be formed at the prevailing iodine
partial pressure to transport the metal in significant
quantities. The transport of chromium would also be
limited by the low gas-phase pressure because stable
CrI2 is nearly three orders of magnitude less volatile
than the other iodides. The mechanism by vapor
transport of FeI2 in consideration of its stability and
volatility is consistent with the observation of only
iron inclusion in the fuel.
Sufficiently stable cesium–metal–oxygen compounds such as Cs–U–O and Cs–Mo–O decompose
CsI and release free iodine, which depends on oxygen
potential84 and supports the hypothesis of iron transport by the iodide process.109
3.16.5.3.3 Cladding corrosion by Cs–Te
mixture

When the temperature and oxygen potential are high
enough in the fuel–cladding gap, liquid Cs–Te dissolves constituents of the cladding; separate metallic
phases are then formed in the fuel cracks and on the
fuel surface. This nonoxidative mode of FCCI is
called CCCT.5,17–19 CCCT was seen in fuel pins
with high O/M ratios (i.e., >1.97), at cladding temperatures >600  C, and at burnups of 5 at.% or


463

greater, and it was characterized by a definite segregation of the cladding constituents of Fe, Cr, and Ni,
with the iron and nickel appearing in the fuel–
cladding gap as a metallic layer separated from the
reacted cladding by a nonmetallic layer containing
the chromium.
CCCT, as its name implies, is a chemical transport
process in which Fe, Ni, and Cr are selectively
removed from the cladding materials and transported
to the fuel surface by a melt of FPs consisting principally of Cs and Te species. Tellurium of the transport
medium is recycled in the melt and its activity in the
melt is related to the ratio of chemically associated Cs
and Te (Cs:Te).
Out-of-pile experiments have provided strong
evidence that the Cs:Te ratio is a critical parameter
that determines both severity and types of FCCI.62
The roles of Cs–Te mixtures in promoting both
FCCI and FPLME are essentially catalytic.4,62,110 In
addition, under certain local conditions of Cs:Te ratio,
stress, and temperature, FPLME may take place.111
CCCT will only occur when Cs:Te <2:1.62 At this
sufficiently high tellurium activity, iron and nickel
tellurides could be formed; however, formation of
chromium telluride depends on the activity of oxygen which determines the formation of other chromium oxide compounds such as Cr2O3 and Cs3CrO4.
The dominant feature of this form of interaction is
the segregation of the chromium from the Fe and Ni
in the reaction product layer.
The tellurium activity in the gap is determined by

an equilibrium reaction between condensed Cs2Te
and oxide fuels at its outer surface. This tellurium
activity-controlling equilibrium reaction is very sensitive to oxygen potential or fuel stoichiometry.
It requires a sufficiently high oxygen potential to
allow the decomposition of cesium telluride. This
corrosion process is determined by the molar ratio
of cesium to tellurium (Cs:Te) in a melt of the mixtures. The thermodynamic activity of tellurium and
the availability of tellurium and cesium for reactions,
are usually expressed as the Cs:Te ratio. The decomposition of cesium telluride occurs in the presence of
the oxygen supplied by the fuel as per the following
reaction.
Cs2 Te þ Cr þ 2O2 Ð Cs2 CrO4 þ Te
Because the above decomposition requires an oxygen
potential high enough to destabilize Cs2Te, a sufficiently high tellurium activity exists for severe
cladding attack. Because Cs2Te is less stable with
increasing oxygen potential than Cs–Mo–O and


Ceramic Fuel–Cladding Interaction

fuel–cladding gap.59 A purposely contaminated
Cs2Te specimen started melting at approximately
700  C, and the melting point of Cs2Te
(810 Æ 10  C) was found to be extremely sensitive to
the presence of impurity oxygen. The changes in
types of cladding attack behavior occurred over relatively constant tellurium activity where the Cs:Te
ratio was varied from 4:1 to 2:1. At Cs:Te ¼ 2:1,
where the change in tellurium activity showed the
most significant and steepest change, the cladding
attack behavior was a type of deep localized intergranular attack. The ratio of Cs:Te < 4:1 was sufficiently low for promotion of FCCI, and FPLME may

occur under conditions of Cs:Te < 2:1 and an oxygen
potential corresponding to stoichiometric or very
slightly hyperstoichiometric fuel.4,62,110 The Cs:Te
ratios in the gap of irradiated pins was low ($2:1)

Cs–U–O, the free tellurium in the gap can lead
to cladding attack by formation of tellurides of iron,
nickel, and chromium according to the following
reactions.
Fe þ 0:9 Te Ð FeTe0:9
Ni þ ð1 À Y ÞTe Ð NiTeð1ÀY Þ

withð1 À Y Þ % 0:6

2Cr þ 3Te Ð Cr2 Te3
From the results of out-of-pile experiments62,94,97,110
with various Cs:Te mixtures, a diagram of the tellurium potential and Cs–Te binary solidus/liquidus
boundaries versus Cs:Te ratio (or composition) is
drawn in Figure 12(a) at 677  C (950 K) which
was taken as a representative temperature for the

T = 950 K
-100

[Cr]ss + <Cr2Te3>

-200

Equivalent fuel O/M
(reaction 1)


aTe

FCCI
Cs: Te < 4:1
Bu ³ 1 at.%
O/M > 1.96
CCCT
Cs: Te < 2:1
Bu ³ 5 at.%
O/M > 1.98

1.9995
1.9990
-300
800

Temperature (ЊC)

10-1
10-2
10-3
10-4

[Fe]ss + <Fe3Te0.9>
[Ni]ss + <Ni3Te2>

2.0005
2.0000


4:1

Liquid

2:13:2 1:1

2:3 1:2

600

(a)

800

T = 950 K
Liquid

Liquid
(Cs-rich)
+
Cs2Te(s)

Liquid
+
Te(s)

Cs2Te Cs Te Cs Te
2 3
5 4
Cs3Te2 CsTe

Cs(s) + Cs2Te(s)

0
Cs

500
600
700
Temperature (ЊC)

800
Liquid
(Te-rich)
Cs2Te(s)

400

400

1:4Cs:Te ratio

Temperature (ЊC)

Oxygen potential DGTe (kJ mol-1)

0

FPLME/FAE
Cs: Te < 2:1
Bu ³ 0.2 at.%

é ³ 3ϫ10-3 s-1

Stress or fracture strain

464

60
40
Composition (at.%)

20

CCCT

700

600
FCCI
500
FPLME/FAE

Te

400
(b)

0

2


4
6
Burnup (at.%)

8

10

Figure 12 (a) Diagram illustrating the relationships between tellurium potential (DGTe), condensed phase Cs: Te ratio in gap,
and the solidus–liquidus phase boundaries of Cs–Te binary mixtures. Reproduced from Adamson, M. G.; Aitken, E. A.;
Lindemer, T. B. J. Nucl. Mater. 1985, 130, 375–392, with permission from Elsevier and (b) schematic illustrations of the
primary environmental domains for the various fuel–cladding (AISI 316) interaction mechanisms involving Cs and Te fission
products. Reproduced from Adamson, M. G.; Aitken, E. A. J. Nucl. Mater. 1985, 132, 160–166, with permission from Elsevier.


Ceramic Fuel–Cladding Interaction

relative to the fission yield ratio ($6:1). It was proposed that the tellurium activity in the gap was determined by an equilibrium reaction between condensed
Cs2Te and oxide fuels at their outer surface. This
tellurium activity-controlling equilibrium is very sensitive to oxygen potential or fuel stoichiometry, in
addition to the temperature.
The cesium telluride reaction with the cladding
constituents, which is a reaction that forms cesium
chromate, iron telluride, and nickel telluride, requires
a high oxygen potential ðDG O2 > À418 kJ molÀ1 Þ.55
Cladding attack proceeds until Cs2Te reacts with the
FPs and forms Cs2Mo2O7 and PdTe at slightly higher
oxygen potential.
Slotted ring tests in Cs–Te mixtures showed that
stress had no significant influence on the grain

boundary corrosion rate.112 In contrast, certain Cs–
Te mixtures have been found to strongly embrittle
stainless steels.111 Corroded cladding materials were
found to have considerably lower properties of
strength and ductility after contact with fuel compared with intact cladding materials irradiated to the
same neutron dose.113 Schematic illustrations of the
primary environmental domains for the various fuel–
cladding (AISI 316 stainless steel) interaction
mechanisms involving Cs and Te are shown in
Figure 12(b). The upper graph shows relationships
between stress level (or strain) and temperature
for FPLME, FCCI, and CCCT, and the lower one

Table 3

465

shows relationships between the temperature and
fuel average burnup.114
It has been postulated that contact of a crack tip
and an agent would cause reduction of the energy
required for propagation of a crack.115 Out-of-pile
tests have been used to try to explain this type of
cladding embrittlement which is termed FAE (fuel
adjacency effect). A chemical state of Cs:Te ratios <1
for embrittlement might be required for FAE.116
As for the dependence of oxygen potential, the
corrosion of cladding materials PE16 and M316
stainless steels by Cs:Te mixtures (1:1, 2:1, and 4:1)
was examined in sealed capsules under partial pressures of O2 set by metal/metal oxide couples at

675  C for 168 h.117 Results of the nature of the corrosion of alloys are shown in Table 3. Detailed features of cladding attack were classified into three
types as follows. Type 1: deep intergranular penetration, embrittlement, and dissolution of the metal;
Type 2: uniform attack of the alloy matrix causing
alternate layering; Type 3: intergranular penetration,
embrittlement, and no dissolution of the metal. As for
the dependence of oxygen potential on cladding
attack, the 2:1 and 4:1 mixtures produced a layering
matrix type of cladding attack and the depth of attack
increased with increasing oxygen potential. On the
other hand, cladding materials experienced severe
intergranular attack for the 1:1 mixtures and the
attack seemed independent of oxygen potential.

The nature of the corrosion of cladding alloys at 948 K of temperature by mixtures of Cs, Te, and oxygen

Alloy

Cs:Te ¼ 1:1 liquid

Cs:Te ¼ 2:1 solid

Cs:Te = 4:1a liquid

Cs:Te = 4:1b liquid

PE16

Deep intergrain

Matrix layer; layers of Cs, Te, Cr

between Ni, Fe

Matrix layer; outer layer
of Fe, Ni; inner layer of
Cs, Te, Cr, Ti

Deep intergrain + matrix
layer; Cs, Cr in gb

M316

Dissolution
Cs, Te, Ni in gb
!150 mm (Cr/Cr2O3)
!150 mm (Mo/MoO2)
!150 mm (Ni/NiO)
Type 1
Deep intergrain
Dissolution
Cs, Te, Ni in gb
!200 mm (Mo/MoO2)
Type 1

a

Corrosion by Cs + O2
0 mm (Cr/Cr2O3)
12 Æ 5 mm (Mo/MoO2)
40 mm (Ni/NiO)
Type 2

Matrix layer; outer layers of Fe, Ni;
inner layer of Cs, Te, Cr

30 Æ 5 mm (Mo/MoO2)
120 mm (Ni/NiO)
Type 2

With Mo/MoO2 buffer only.
With Ni/NiO buffer only.
gb, grain boundary.
Source: Pulham, R. J.; Richards, M. W. J. Nucl. Mater. 1990, 172, 206–219.
b

15–20 mm
Type 2
Matrix layer; layer of Cs,
Te, Cr
Between Fe, Ni

90 mm
Type 2 þ 3
Deep intergrain;
Cs, Cr in gb
Corrosion by Cs þ O2

48 Æ 5 mm

!200 mm

Type 2


Type 3


Ceramic Fuel–Cladding Interaction

3.16.6 Inhibition Methods for Oxide
Fuel and Cladding Interaction

coexistent phases all have the potential to prevent
oxidation of alloyed chromium: Ti/TiO, TiO/Ti2O3,
V/VO, Nb/NbO, NbO/NbO2, and Cr/Cr2O3.
Figure 13 shows the oxygen potential of these candidate inhibitors as a function of temperature in addition to those of the oxide fuels with various O/M
ratios.62 Cesium reacts with the oxides of Nb, Ti, Cr,
and V, and affinity of cesium is in the following order:
Nb > Ti > Cr ¼ V.118
Experimental verification of inhibitor effectiveness was done to confirm other pertinent physical
and chemical properties, neutronics, and oxidation
kinetics. Fuel pellets coated with getter materials
would be reasonable as they efficiently inhibit FCCI.
Meanwhile in addition to the kinetics of oxygen
trapping, the behavior of fuel and cladding contacting
with getters has to be technologically considered.
When the oxygen potentials increase, formation of
cesium uranate causes a deformation of the cladding
near the fissile–fertile pellet transition zone.8,119,120
The cladding attack near the fissile and fertile
pellet transition zone has been reported to be significant.40,95,121,122 For these reasons, the oxygen potential should be kept at a sufficiently low value.
Therefore, placing oxygen getters in the form of
pellets at the end of the fissile column could be a

possible method to control the oxygen potential.

The principal function of buffer–getter materials in
fuel pins is as an inhibitor acting as an oxygen sink,
thereby reducing oxidation of the cladding. When
oxygen is transported from the fuel to the buffer–
getter material, oxidation of the buffer–getter occurs
instead of cladding materials. Candidate materials
can be separated into buffers and getters by considering their relative affinity for oxygen. Selection and
placement of buffer–getter materials in fuel pins are
basically evaluated relative to the chemical and neutronic characteristics of the materials.
Buffers from elements such as V, Nb, and Cr are
considered, which will maintain a fixed fuel O/M
ratio under given conditions. On the other hand,
getters are also considered from elements such as
Zr, U, Ti, and certain rare earth metals, which
will reduce the fuel O/M ratio from the initial
value. Thermal properties of the fuel are generally
degraded; as a consequence, getters lower the O/M
ratio of the oxide. At typical temperature ranges of
the cladding in FR fuel pins, a buffer–getter additive
must maintain the fuel O/M ratio of the periphery
below the initial value in order to protect Cr in
the cladding from oxidation. The following pairs of

Oxygen potential, ΔGo2 (kcal mol−1)

−50

−209


Fuel outer surface
temperature

1.99
1.98
1.97
1.96
1.95
1.94
1.93

2.000

4 2
99 1.99

−100
K)

O5
/Nb 2

−150

−418

1.

O2

Nb
O3
Cr 2
Cr/
O2
/Nb
NbO O
b
N
Nb/

6 (U

96

9

1.

9
1.9

1.92

−628
Fuel center line
temperature

O
V/V


Oxygen potential, ΔGo2 (kJ mol−1)

466

O3
/Ti 2

−200

TiO

−837

iO

Ti/T

1000

1500
2000
Temperature (K)

2500

Figure 13 Oxygen potential diagram showing data for MOX (25% Pu) fuel and candidate buffers. Reproduced from
Adamson, M. G. In Technical Committee Meeting on Fuel and Cladding Interaction, Proceedings of the International Working
Group on Fast Reactors, IWGFR/16, Tokyo, Japan, Feb 21–25, 1977; IAEA: Austria, 1977; pp 108–136, with permission
from IAEA.



Ceramic Fuel–Cladding Interaction

Furthermore, niobium foils were successfully applied
to reduce or suppress oxidizing grain boundary attack
of the cladding, when they were placed between
the fuel and cladding in out-of-pile experiments
performed at 700 and 800  C for 1000 h.40,121 Porous
niobium and titanium metal pellets were used in
feasibility irradiation tests as oxygen absorbers at
the ends of the fuel column.118 However, they were
ineffective in gettering excess oxygen of the fuel
pin or in reducing the average O/M ratio of the
fuel column.
The effectiveness of niobium metal coatings on
the inner cladding surface to act as an oxygen getter
was examined in an in-pile test.123 The niobium
coating on the inner cladding surface worked well
in reducing the oxygen potential. However, it was
found that niobium reacted with noble metal FPs
and formed relatively stable intermetallic phases.

3.16.7 Nonoxide Ceramic Fuels and
Cladding Interaction
Nonoxide ceramic (MX-type) fuels are generally
irradiated at lower temperatures and lower radial
temperature gradients than oxide fuels, although at
high linear heat rating, which results in low FP
release rates. The volatile FPs (Br, I, Cs, and Rb)

do not form carbides or nitrides. In particular,
MX-type fuel pins are kept with a low oxygen potential at the inner cladding surface during irradiation;
therefore, severe oxidative chemical interaction of
cladding with the FPs is not expected. A number
of irradiation experiments to study FCCI have
been done with MX-type fuels (carbide124–133 and
nitride134,135). The compatibility with cladding materials has been investigated in out-of-pile examinations136,137 and thermodynamic analyses of fuel and
cladding have been performed.138,139 Although there
are some exceptions, FPs from MX-type fuels have
not been shown experimentally to have an important
role in FCCI. As a consequence, a little contribution
of FPs to FCCI of MX-type fuels has been considered. Unlike the case of oxide fuels, the carburizing
and nitriding of cladding, and also the formation of
intermetallic compounds of fuel and cladding, have
been investigated as major FCCI of MX-type fuels.
3.16.7.1

FCCI of Carbide Fuel

FCCI in fuel pins containing carbide fuel consists
of clad carburization, slight reaction with FPs, and

467

formation of intermetallic compounds. Although
mixed carbides are more impure and inherently less
stable chemically than MOX, carbide fuel presents
fewer compatibility problems than oxide fuels (see
Chapter 2.04, Thermodynamic and Thermophysical Properties of the Actinide Carbides; Chapter
3.03, Carbide Fuel).

3.16.7.1.1 Chemical reactions with FPs

In carbide fuel pins, the oxygen potential is too low
for the formation of complex oxides. In addition,
an increase in carbon potential during burnup is
not to be expected. As the FPs chemically react
with carbon, the total carbon-to-metal (C/M) ratio
decreases with increasing burnup. FP attack on the
cladding of a carbide fuel pin cannot cause severe
interaction.
Out-of-pile compatibility tests using the burnup
simulated hyperstoichiometric uranium carbide
(UC1 þ x) fuel containing UC2 as the second phase
were performed with Type 1.4988 steel cladding.139
The annealings were carried out up to 5000 h at
temperatures from 500 to 900  C. Significant FPcladding material interactions were not observed in
the simulated specimen of 10 at.% burnup. Tellurium
caused some intergranular attack at a high concentration of 20 wt% in the fuel, and it reduced carbide
precipitation in the cladding.
Reactions with the cladding in the presence of the
carbide fuel occurred only with free iodine. Most of
the iodine available in the irradiated fuel is present as
cesium compounds such as Cs2I2 and CsI.140 Because
of the formation of cesium and iodine compounds,
the corrosive effects of iodine will be reduced drastically because gaseous cesium iodide molecules are
thermodynamically stable. Larger additions of other
FP elements did not cause severe reactions with
cladding at temperatures up to 800  C.
Thin reaction layers ($10 mm) containing nickel,
iron, uranium, plutonium, and rare earths were found

at high specific linear heating rates.140–142 Severe
interactions were thus not expected nor were they
observed. The phases observed were in agreement
with experiments simulating different stoichiometries and burnups.143
3.16.7.1.2 Formation of intermetallic
compounds

The fuel surface temperature of carbide fuels is
generally below 1000  C. Hypostoichiometric carbide
UC1 À x forms uranium carbide and free U metal on
the basis of the U–C phase diagram.144–147 Plutonium